Online proceedings - EDA Publishing Association
Online proceedings - EDA Publishing Association
Online proceedings - EDA Publishing Association
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COLLECTION OF PAPERS PRESENTED AT THE<br />
Symposium on<br />
Design, Test, Integration and Packaging of<br />
MEMS/MOEMS<br />
Chairs/Editors<br />
Bernard COURTOIS<br />
Jean Michel KARAM<br />
Ai-Qun LIU<br />
Ryutaro MA<strong>EDA</strong><br />
Pascal NOUET<br />
Peter SCHNEIDER<br />
11-13 May 2011<br />
Aix-en-Provence, France<br />
Sponsored by:
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Table of Contents<br />
Wednesday 11 May<br />
INVITED TALK 1: TRENDS AND CHALLENGES IN MODERN MEMS SENSOR PACKAGES<br />
Trends and Challenges in Modern MeMs sensor PaCkages .................................................................... 1<br />
Jiri Marek<br />
SESSION C1: COMPACT AND BEHAVIOURAL MODELLING<br />
dynaMiC Behavior of resonanT PiezoeleCTriC, CanTilevers ParTially iMMersed in liquid . ....... 4<br />
M. Maroufi, Sh. Zihajehzadeh, M. Shamshirsaz, A.H. Rezaie, M.B. Asgari<br />
Behavioural Modelling of MeMs osCillaTors and Phase noise siMulaTion .................................. 392<br />
Guillaume Papin, Raphael Levy, Gaelle Lissorgues, Patrick Poulichet<br />
reliaBle sysTeM-level Models for eleCTrosTaTiCally aCTuaTed deviCes<br />
under varying aMBienT CondiTions: Modeling and validaTion .............................................................. 8<br />
Gabriele Schrag, Martin Niessner, Gerhard Wachutka<br />
invesTigaTion on The effeCT of geoMeTriCal diMensions on The ConduCTive<br />
Behaviour of a MeMs ConveCTive aCCeleroMeTer . .....................................................................................14<br />
A.A. Rekik, B. Mezghani, M. Masmoudi, F. Azaïs, N. Dumas, F. Mailly, P. Nouet<br />
design and siMulaTion of an on-ChiP oversaMPling ConverTer<br />
wiTh a CMos-MeMs differenTial CaPaCiTive sensor ....................................................................................18<br />
Ma Li Ya, Anis Nurashikin Nordin, Sheroz Khan<br />
SESSION T1: FABRICATION AND PACKAGING<br />
faBriCaTion of high asPeCT raTio nanoPorous array on siliCon ..........................................................23<br />
Jing-Yu Ho, Gou-Jen Wang<br />
faBriCaTion MeThods for The ManufaCTure of saPPhire MiCroParTs . .................................................29<br />
David M. Allen, Roxana Redondo, Maximilien Dany<br />
CharaCTerisaTion and CoMParison of waTer and alCohol as CaTalysTs ..........................................35<br />
in vaPour Phase hf eTChnig of siliCon oxide filMs<br />
D. Drysdale, T. O’Hara, C. H. Wang<br />
agile MeMs PaCkaging for Mass CusToMized MeMs ProduCTs ................................................................41<br />
Jens G. Kaufmann, David Flynn, Keith Brown, Marc P.Y. Desmulliez<br />
wafer-level glass-CaPs for advanCed oPTiCal aPPliCaTions . ................................................................46<br />
Juergen Leib, Oliver Gyenge, Ulli Hansen, Simon Maus, Karin Hauck, Kai Zoschke, Michael Toepper<br />
SESSION C2: MODEL ORDER REDUCTION<br />
reduCed order Modelling of MeMs dynaMiCs . ............................................................................................53<br />
Stefano Mariani, Saeed Eftekhar Azam, Aldo Ghisi, Alberto Corigliano, Barbara Simoni<br />
a Model for Two-diMensional arrays of CanTilevers in The dynaMiC regiMe . ...............................59<br />
Hui Hui, Michel Lenczner<br />
sensiTiviTy analysis and adaPTive MulTi-PoinT MulTi-MoMenT Model<br />
order reduCTion in MeMs design . ......................................................................................................................64<br />
Andreas Köhler, Sven Reitz, Peter Schneider<br />
vii
SESSION T2: ENERGY HARVESTING<br />
inTegraTion of ferroeleCTriC BaTio 3<br />
on MeTalliC ni<br />
TaPes for Power generaTion . ..............................................................................................................................72<br />
Greg Collins, Emanuel Silva, Ming Liu, David Elam, Chunrui Ma, Andrey Chabanov, Arturo Ayon, Chonglin Chen,<br />
Jie He, Jiechao Jiang, Efstathios Meletis<br />
an eleCTroMeChaniCal Model for ClaMPed-ClaMPed BeaM TyPe PiezoeleCTriC TransforMer ... 75<br />
Chi-Shao Chen, Chia-Che Wu<br />
sTudy of BlaCk siliCon oBTained By deeP reaCTive ion eTChing ...........................................................81<br />
aPProaCh To aChieving The hoT sPoT of a TherMoeleCTriC energy harvesTer<br />
K.N Nguyen, D.Abi-Saab, M. Malak, P. Basset, E. Richalot, N. Pavy, F. Flourens, F. Marty, D. Angelescu, Y. Leprince-Wang, T.Bourouina<br />
SESSION C3: MODELLING AND VALIDATION<br />
PerforManCe evaluaTion of MeMs PiezoeleCTriC inerTial energy generaTor . ...............................85<br />
Aliza Aini Md Ralib, Anis Nurashikin Nordin, Raihan Othman, Hanim Salleh,<br />
ParaMeTer design of Triaxial MiCroaCCeleroMeTers wiTh PiezoeleCTriC Thin-filM ...................90<br />
Jyh-Cheng YU, Chungda Lee<br />
Modeling and exPeriMenTal validaTion of leviTaTing sysTeMs<br />
for energy harvesTing aPPliCaTions . ..............................................................................................................97<br />
Giorgio De Pasquale, Sonia Iamoni, Aurelio Somà<br />
MoleCular dynaMiC siMulaTion of nanoParTiCle size effeCT on MelTing PoinT of gold . .........103<br />
P. Nayebi, M.Shamshirsaz, K. Mirabbaszadeh, E. Zaminpeyma, M.B. Asgari<br />
sTress idenTifiCaTion of Thin MeMBrane sTruCTures By dynaMiC MeasureMenTs . ......................106<br />
Steffen Michael, Christoph Schäffel, Sebastian Voigt, Roy Knechtel<br />
SESSION T3: SENSORS AND ACTUATORS<br />
Meso-sCale aCTuaTor design for The inTegraTed dynaMiC<br />
alignMenT of a lensleT array wiThin a PaCkage ......................................................................................110<br />
Stefan Wilhelm, Robert W. Kay, Marc P.Y. Desmulliez<br />
iMPleMenTing MeMs resonaTors in 90 nM CMos . ..........................................................................................116<br />
J.E. Ramstad, J.A. Michaelsen, O. Soeraasen, D. Wisland<br />
The influenCe of adhesive MaTerials on ChiP-on-Board PaCking of MeMs MiCroPhone . ............122<br />
Cheng-Hsin Chuang, Yi-Hsuan Huang, Shin-Li Lee<br />
Model of a volTage driven CaPaCiTive CouPled MiCro eleCTro-MeChaniCal rf swiTCh. ............128<br />
P. Heeb, W. Tschanun, R. Buser<br />
a Closed-looP MiCroMaChined aCCeleroMeTer Based on TherMal ConveCTion . ..........................134<br />
Alexandra Garraud, Philippe Combette, Benoît Charlot, Pierre Loisel, Alain Giani<br />
viii
PANEL DISCUSSION TEXTILE MICROSYSTEMS ...................................................................................137<br />
PiezoeleCTriC Charging for sMarT faBriC aPPliCaTions ........................................................................138<br />
R. Hackworth , J. R. Moriera, R. Maxwell, R. Kotha, A.A. Ayon<br />
MeTer-sCale surfaCe CaPaCiTive TyPe of TouCh sensors<br />
faBriCaTed By weaving ConduCTive-PolyMer-CoaTed fiBers ................................................................142<br />
Seiichi Takamatsu, Takeshi Kobayashi, Nobuhisa Shibayama, Koji Miyake, Toshihiro Itoh<br />
POSTER INTRODUCTION SESSION<br />
on-wafer-PaCkaging of CrysTal quarTz Based<br />
devises using low-TeMPeraTure anodiC Bonding . .....................................................................................148<br />
Y. Zimin, T. Ueda<br />
a novel self-Powered MeThod for PiPe flow Measuring .......................................................................152<br />
Song Hao Wang, Ronald Garcia, Pei Hua Chang<br />
a MiCrofluidiC ChiP wiTh single-ParTiCle-Based arrays using eleCTroosMoTiC flow . ...............157<br />
Chun-Ping Jen, Ju-Hsiu Hsiao<br />
a novel full range vaCuuM Pressure sensing TeChnique using free daMPing deCay<br />
of MiCro-Paddle CanTilever BeaM defleCTed By eleCTrosTaTiC forCe . ...........................................160<br />
Guan-Lan Chen, Chi-Jia Tong, Ya-Chi Cheng, Yu-Ting Wang, Ming-Tzer Lin<br />
design and develoPMenT of viBraTional MeChanoeleCTriCal MeMs TransduCer<br />
for MiCroPower generaTion . ............................................................................................................................164<br />
Rolanas Dauksevicius, Genadijus Kulvietis, Vytautas Ostasevicius, Ieva Milasauskaite<br />
inTerfaCial ConfiguraTions and Mixing PerforManCes of fluids<br />
in sTaggered Curved-Channel MiCroMixers ..............................................................................................170<br />
Jyh Jian Chen, Chun Huei Chen, Shian Ruei Shie<br />
MiCro ProBe array faBriCaTion By using The MiCrolens array Mask<br />
Through ProxiMiTy PrinTing . ............................................................................................................................176<br />
Tsung-Hung Lin, Hsiharng Yang, Ching-Kong Chao<br />
CresCenT shaPed alignMenT Marks aPPliCaBle To self-alignMenT of MiCro-ParTs<br />
wiTh and wiThouT PosiTive and negaTive Poles . ........................................................................................180<br />
Shouhei Shiga, Dong F. Wang, Takao Ishida, Ryutaro Maeda<br />
siMulaTion of 3d soi-sTruCTures for MeMs eleMenTs . .............................................................................184<br />
Igor Kogut, Victor Holota, Victor Dovhij, Anatoliy Druzhinin<br />
sTudies of oPTiCal and CrysTal ProPerTies of ald grown zno . ............................................................185<br />
David Elam, Anastasiia Nemashkalo, Yuri Strzhemechny, Chonglin Chen, Arturo Ayon, Andrey Chabanov<br />
a MeThodology for The Pull-in volTage of ClaMPed diaPhragMs . ....................................................187<br />
Joseph Lardiès, Marc Berthillier<br />
oPTiMisaTion and realisaTion of a PorTaBle nMr aPParaTus and MiCro anTenna for nMr . .......193<br />
Patrick Poulichet, Latifa Fakri-Bouchet, Christophe Delabie, Lionel Rousseau, Abdenasser Fakri, Anne Exertier<br />
ix
Convex Corner CoMPensaTion for a CoMPaCT seisMiC Mass<br />
wiTh high asPeCT raTio using anisoTroPiC weT eTChing of (100) siliCon ............................................197<br />
Jyh-Cheng YU<br />
a PrograMMaBle neural MeasureMenT sysTeM for sPikes and loCal field PoTenTials . ...........200<br />
Jonas Pistor, Janpeter Hoeffmann, Dagmar Peters-Drolshagen, Steffen Paul<br />
Thursday 12 May<br />
PANEL DISCUSSION HERMETICITY TESTS IN MEMS ................................................................................206<br />
Marc Desmulliez, Suzanne Costello, Heriot-Watt University, Edinburgh, UK., Wolfgang Reinert, Fraunhofer Institute for Silicon Technology<br />
Fraunhofer, Germany, Steven Martell, Sonoscan Inc., USA<br />
herMeTiCiTy TesT MeThods for MeMs:where are we? ...............................................................................207<br />
Marc Desmulliez<br />
sTandards for herMeTiCiTy TesT MeThods . ..................................................................................................208<br />
Suzanne Costello<br />
q-faCTor MoniToring as a 100% leak sCreen in indusTrial aPPliCaTions . ..........................................209<br />
Wolfgang Reinert<br />
evaluaTing The seal inTegriTy of MeMs herMeTiC PaCkages ................................................................210<br />
Steven R. Martell<br />
PANEL DISCUSSION HIGH ADDED VALUE MEMS ........................................................................................211<br />
Jérémie BOUCHAUD, IHS iSuppli, Munich, Germany, Jean Michel Karam, MEMSCAP, Bernin, France, Sean Neylon, Colibrys, Neuchâtel,<br />
Switzerland, Thierry Brisard, NEOSENS, Toulouse, France<br />
SPECIAL SESSION ON NETWORKED MICROSYSTEMS FOR GREEN AND LIFE INNOVATION<br />
inTegraTed sensing sysTeMs for healTh and safeTy . ..............................................................................212<br />
Kiyoshi ITAO<br />
design, faBriCaTion, and inTegraTion of PiezoeleCTriC MeMs deviCes<br />
for aPPliCaTions in wireless sensor neTwork . ..........................................................................................217<br />
Jian Lu, Yi Zhang, Toshihiro Itoh, Ryutaro Maeda<br />
novel MeMs digiTal TeMPeraTure sensor for wireless avian-influenza<br />
MoniToring sysTeM in PoulTry farM . .............................................................................................................222<br />
Yi Zhang, Hironao Okada, Takeshi Kobayashi, Toshihiro Itoh<br />
aPPliCaTion of wireless sensor nodes To CoMMerCial Power ConsuMPTion MoniToring . .........227<br />
Toshihiro Itoh, Jun Fujimoto, Ryutaro Maeda, Takeshi Kobayashi, Toshihiro Itoh, Ryutaro Maeda<br />
develoPing MeMs dC eleCTriC CurrenT sensor for end-use MoniToring of dC Power suPPly .......231<br />
Kohei Isagawa, Dong F. Wang, Takeshi Kibayashi, Toshihiro Itoh, Ryutaro Maeda<br />
low Power analog To digiTal ConverTor wiTh digiTal CaliBraTion for sensor neTwork .......237<br />
Tsukasa Fujimori, Hiroshi Imamoto, Hideaki Kurata, Yasushi Goto, Toshihiro Ito, Ryutaro Maeda<br />
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Friday 13 May<br />
SESSION C4: APPLICATIONS I<br />
large area adaPTaTive fluidiC lens ..............................................................................................................241<br />
Solon Mias, Aurélien Bancaud , Henri Camon<br />
faBriCaTion and CharaCTerisTiCs of a fused siliCa-Based oPTiCal waveguide<br />
wiTh feMToseCond fiBer laser Pulses . ..........................................................................................................245<br />
Ting-Chou Chang, Chien-Hsing Chen, Wei-Hung Shih, Jian-Neng Wang, Chai-Yu Lee, Jaw-Luen Tang, Shau-Chun Wang,<br />
Lai-Kwan Chau, Wei-Te Wu<br />
a MulTilevel PolyMer ProCess for liquid direCT<br />
enCaPsulaTion for oPTo-fluidiC aPPliCaTion . .............................................................................................249<br />
Remy Bossuyt, Laurent Mazenq, Véronique Conédéra, Jérôme Ballet, Anne-Marie Gué, Jean-Paul Cano, Henri Camon<br />
MulTiPle-ouTPuT MeMs dC/dC ConverTer ......................................................................................................253<br />
A. Chaehoi, M. Begbie, D. Weiland, S. Scherner<br />
design of The siliCon MeMBrane of high fideliTy and high effiCienCy MeMs MiCrosPeaker .......258<br />
Iman Shahosseini, Elie Lefeuvre, Emile Martincic, Marion Woytasik, Johan Moulin, Souhil Megherbi, Romain Ravaud, Guy Lemarquand<br />
ModulaTion insTaBiliTy in rf MeMs deviCes . ...............................................................................................263<br />
Romolo Marcelli, Giancarlo Bartolucci, Giorgio De Angelis, Andrea Lucibello, Emanuela Proietti<br />
SESSION T4: EMBOSSING AND MOULD<br />
sTudy of sCreen-PrinTing MiCrolens array using eleCTroforMing Molds ....................................268<br />
Ming-Je Lin, Hsiharng Yang, Feng-Tsai Weng<br />
PolyMer-Based faBriCaTion TeChniques for enClosed<br />
MiCroChannels in BioMediCal aPPliCaTions ..................................................................................................273<br />
Annabel Krebs, Thorsten Knoll, Dominic Nussbaum, Thomas Velten<br />
hoT eMBossing of BiodegradaBle PolyMers . ..............................................................................................278<br />
Matthias Worgull, Alexander Kolew, Heilig Markus, Marc Schneider, Heinz Dinglreiter<br />
su-8-Based raPid Tooling for TherMal roll eMBossing. .........................................................................279<br />
Khaled Metwally, Laurent Robert, Roland Salut, Chantal Khan Malek<br />
MulTi-CoMPonenT hoT eMBossing of MiCro- and nanosysTeMs ............................................................284<br />
Alexander Kolew, Markus Heilig, Karsten Sikora, Daniel Muench, Matthias Worgull<br />
INVITED TALK 2: SUCCESS IN MEMS, «FROM DRIE TECHNOLOGY TO SOCIAL INNOVATION»<br />
suCCess in MeMs, «froM drie TeChnology To soCial innovaTion» . .......................................................288<br />
Susumu KAMINAGA<br />
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SESSION C5: APPLICATIONS II<br />
BrighTness enhanCeMenT of oleds By using<br />
MiCrolens array filM wiTh siliCon oil and ag ParTiCles .......................................................................294<br />
Shan-Shan Hsu, Tung-Yu Chang, Hsiharng Yang, Jen-Sung Hsu<br />
inTegraTion of hyBrid oPTiCal filTer wiTh Buried quad Pn-junCTion PhoTodeTeCTor<br />
for MulTi-laBeling fluoresCenCe deTeCTion aPPliCaTions. ..................................................................300<br />
Charles Richard, Patrick Pittet, Stéphane Martel, Luc Ouellet, Guo-Neng Lu, Vincent Aimez, Paul G. Charrette<br />
faBriCaTion and CharaCTerisTiCs of a fusedsiliCa-Based oPTiCal<br />
waveguide wiTh feMToseCond - fiBer laser Pulses ..................................................................................305<br />
Ting-Chou Chang, Chien-Hsing Chen, Wei-Hung Shih, Jian-Neng Wang, Chai-Yu Lee, Jaw-Luen Tang,<br />
Shau-Chun Wang, Lai-Kwan Chau, Wei-Te Wu<br />
CaPaCiTive MiCroPhone faBriCaTed wiTh CMos-MeMs surfaCe-MiCroMaChining TeChnology .....309<br />
Josué Esteves, Libor Rufer, Gustavo Rehder<br />
a novel inTegraTed soluTion for The ConTrol and diagnosis<br />
of eleCTrosTaTiC MeMs swiTChes . ....................................................................................................................315<br />
Carlo Trigona, Norbert Dumas, Laurent Latorre, Pascal Nouet<br />
an ulTra low Power TeMPeraTure sensor for sMarT PaCkaging MoniToring . ...............................320<br />
Souha Hacine, Frederick Mailly, Norbert Dumas, Laurent Latorre, Pascal Nouet<br />
SESSION T5: RELIABILITY, TESTING AND MEASUREMENT<br />
aCCuraTe TherMal CharaCTerizaTion of Power seMiConduCTor PaCkages<br />
By TherMal siMulaTion and MeasureMenTs . ............................................................................................324<br />
Andras Vass-Varnai, Robin Bornoff, Sandor Ress, Zoltan Sarkany, Sandor Hodossy, Marta Rencz<br />
linear energy ConTrol of laser drilling and iTs aPPliCaTion<br />
for TfT-lCd BrighT Pixel rePairing .................................................................................................................330<br />
Taco Chen, Ming-Tzer Lin<br />
MeasureMenT of eleCTriCal ProPerTies of MaTerials under The oxide layer<br />
By MiCrowave-afM ProBe . ......................................................................................................................................334<br />
Lan Zhang, Yang Ju, Atsushi Hosoi, Akifumi Fujimoto<br />
aMPliTude enhanCeMenT using viBraTion Mode loCalizaTion<br />
wiTh a single MiCro-MeChaniCally CouPled BeaM-shaPed resonaTor array . ................................339<br />
Keisuke Chatani, Dong F. Wang, Tsuyoshi Ikehara, Ryutaro Maeda<br />
uniaxial MeChaniCal sTress and nanoindenTaTion To CharaCTerize Thin MulTilayers ...........344<br />
Thibaut Fourcade, Jeremie Dhennin, Xavier Chauffleur, Mikaël Colin, Jean-Michel Desmarres, Joël Alexis<br />
eleCTriCal and MeChaniCal CharaCTerizaTion of laTeral neMs swiTChes . ...................................348<br />
R. Hinchet, L. Montès, G. Bouteloup, G. Ardila, R. Parsa, R.T. Howe, H.-S. Philip Wong, K. Akarvardar<br />
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SPECIAL SESSION OF BIO-MEMS/NEMS<br />
a dieleCTroPhoreTiC PreConCenTraTor wiTh CirCular MiCroeleCTrodes<br />
for BiologiCal Cells in sTePPing eleCTriC fields .....................................................................................352<br />
Chun-Ping Jen, Ho-Hsien Chang<br />
a novel su-8 MiCrogriPPer wiTh exTernal aCTuaTor for BiologiCal Cells ManiPulaTion . .......356<br />
M. Mehdi S. Mousavi, Giorgio De Pasquale, Aurelio Somà, Eugenio Brusa<br />
ParTiCle foCusing in a ConTaCTless dieleCTroPhoreTiC MiCrofluidiC ChiP<br />
wiTh insulaTing sTruCTures .............................................................................................................................362<br />
Chun-Ping Jen, Hsin-Yuan Shih, Yung-Chun Lee, Fei-Bin Hsiao<br />
inCreasing densiTy of anTiBody-anTigen Binding on a sensor surfaCe<br />
By ConTrolling MiCrofluidiC environMenTs . ............................................................................................366<br />
Chia-Che Wu, Ling-Hsuan Hung, Ching-Hsiu Tsai, Yao-Lung Liu<br />
faBriCaTion and aPPliCaTion of iron-oxide nanoParTiCle/PdMs Cone in laB on a ChiP . ...............372<br />
Cheng-Chun Huang, Ming-Dao Wu, Yu-Chi Wang, Wen-Pin Shih<br />
diaMond-Based TeChnology dediCaTed To MiCro<br />
eleCTrode arrays for neuronal ProsTheses . ............................................................................................378<br />
A. Bongrain, A. Bendali, G. Lissorgues, Lionel Rousseau, B. Yvert, E. Scorsone, P.Bergonzo, S. Picaud<br />
MeasureMenT of diffusiviTy in nanoChannels . ........................................................................................382<br />
Yu-Tze Tsai, Gou-Jen Wang<br />
energy harvesTing sysTeM for CardiaC iMPlanT aPPliCaTions ...........................................................387<br />
Martin Deterre, Bertrand Boutaud, Renzo Dalmolin, Sébastien Boisseau, Jean-Jacques Chaillout, Elie Lefeuvre, Elisabeth Dufour-Gergam<br />
index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 398<br />
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XIV
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Trends and Challenges in<br />
Modern MEMS Sensor Packages<br />
Jiri Marek<br />
Robert Bosch GmbH<br />
Tübinger Straße 123<br />
72762 Reutlingen, Germany<br />
Abstract- Modern MEMS sensors for automotive as well as<br />
consumer electronics face a continuous pressure for size<br />
reduction. This can be met not only by a consistent shrink of<br />
sensing elements and ASICs resulting in a smaller package but<br />
also through the integration of several sensors into one system.<br />
The trends and challenges of this steady shrink will be<br />
explained and several examples of current automotive and<br />
consumer MEMS sensors will be shown.<br />
Growth ~10% p.a.<br />
I. INTRODUCTION<br />
A car is skidding and stabilizes itself without driver<br />
intervention; a laptop falls to the floor and protects the hard<br />
drive by parking the read/write drive head automatically<br />
before impact; an airbag fires before the driver involved in<br />
an impending automotive crash impacts the steering wheel<br />
thereby significantly reducing physical injury risk; – all<br />
these systems are based exclusively on MEMS sensors.<br />
These crucial MEMS sensor components of electronic<br />
control systems are making system reactions to human<br />
needs more intelligent, precise, and at much faster reaction<br />
rates than humanly possible.<br />
Important prerequisites for the success of sensors are their<br />
size, functionality, power consumption and costs. This<br />
technical progress in sensor development is realized by<br />
micro-machining. The development of these processes was<br />
the breakthrough to industrial mass-production for microelectro-mechanical<br />
systems (MEMS). Besides leading-edge<br />
micromechanical processes, innovative and robust ASIC<br />
designs, thorough simulations of the electrical and<br />
mechanical behaviour, a deep understanding of the<br />
interactions (mainly over temperature and lifetime) of the<br />
package and the mechanical structures are needed. This was<br />
achieved over the last 20 years by intense and successful<br />
development activities combined with the experience of<br />
volume production of billions of sensors.<br />
II.<br />
MARKET AND DRIVERS<br />
A. Market Size<br />
The growth of the MEMS market and the market<br />
segmentation is shown in Fig. 1 (source: iSuppli). Today’s<br />
market size is around 7 billion US dollars with four main<br />
segments:<br />
Source – iSuppli Corporation MEMS market tracker, H2 2010<br />
Figure 1: MEMS market<br />
• data processing (mainly ink jet printer nozzles)<br />
• automotive<br />
• mobile and consumer electronics<br />
• industry and process control<br />
B. Market Drivers for MEMS Sensors<br />
For MEMS Sensors there are several drivers which push<br />
new developments. In different markets the drivers are<br />
similar but have a different ranking.<br />
For automotive MEMS sensors the main drivers are:<br />
1. high functional requirements (high accuracy, selftest,<br />
advanced safety concepts)<br />
2. high reliability and quality (reliability for 15 years<br />
with failure rates of less than 1ppm under extreme<br />
environmental conditions)<br />
3. low price<br />
MEMS sensors for consumer electronics applications face<br />
different drivers:<br />
1. low price (
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Figure 2: Package roadmap of automotive acceleration sensors<br />
III. ACCELERATION SENSORS<br />
The success story of MEMS acceleration sensors started<br />
nearly 20 years ago with first high-g sensors for airbag<br />
applications and continued with low-g sensors for ABS,<br />
ESP, etc. Today over 90% of all new passenger cars are<br />
sold with an airbag system with at least one<br />
micromechanical acceleration sensor inside. These<br />
impressing equipment rates are only possible due to a<br />
continous and massive reduction of the costs of an airbag<br />
system and therefore also of the integrated acceleration<br />
sensor. This is achieved by strong improvements in the<br />
micromechanical sensing elements, the ASIC and – last but<br />
not least – the packaging. Fig. 2 shows the roadmap of<br />
packages for airbag acceleration sensors at Bosch. The first<br />
sensor, issued in the late 1970s, was a mechanical sensor<br />
element in a metal can. In the mid 1980s the first<br />
mechanical sensor followed with the ASIC integrated in the<br />
same metal package. It was supplanted in 1996 by the first<br />
generation of micromechanical acceleration sensors in a<br />
PLCC28 package. The current generation of airbag<br />
accelerometers, starting in 2010, uses a SOIC8 package.<br />
This corresponds to a size reduction of more than 85% in 14<br />
years.<br />
The massive size reduction was achieved by several steps<br />
in technology development. Due to design and process<br />
progress the micromechanical sensor element could be<br />
drastically reduced in size. The use of modern technologies<br />
in IC processes led to a steady decrease of the ASIC size at<br />
the same time – despite enhanced sensor performance and<br />
higher self-test capabilities. With sophisticated state of the<br />
art simulations – fed by the experience of several sensor<br />
generations and of far more than 1 billion sensors produced<br />
– key parameters of the package are optimized. The most<br />
important of those are<br />
• overall geometry (package height, length and width vs.<br />
die size, symmetry, …)<br />
• leadframe design (size, thickness, structure, …)<br />
• die-attach (material parameters like E-modulus,<br />
thickness,…)<br />
• mold compound (CTEs, …)<br />
• mold coverage (overall portion of mold compound vs.<br />
Silicon content of the package)<br />
The main hurdles for a more aggressive package size<br />
reduction are the the capability for further processing and<br />
the extreme environmental conditions automotive sensors<br />
have to withstand.<br />
Figure 3: Footprint of automotive and CE acceleration sensors<br />
Figure 4: Crosssection and SEM picture of BMA220 (© Chipworks)<br />
The consumer electronics (CE) industry has even higher<br />
constraints regarding package size (footprint as well as<br />
height). Bosch’s first acceleration sensor for CE in 2006<br />
reduced the automotive package size to a 4 4 mm² QFNpackage<br />
by half. Already one year later the size was further<br />
reduced to 3 3 mm². At the beginning of 2010 Bosch<br />
introduced the BMA220 - world’s first digital acceleration<br />
sensor in a 2 2 mm² LGA package. Fig. 3 shows the<br />
footprint development of automotive and CE sensors.<br />
One major step towards the 2 2 cm² package was the<br />
transition from side-by-side assembly to 3D stacked<br />
assembly. With this 3D Integration approach the ASIC is<br />
stacked on the micromechanical sensor element. Fig. 4<br />
depicts insights into the construction of the BMA220.<br />
IV. INERTIAL COMBI-SENSORS<br />
Sooner or later the further size reduction will become<br />
increasingly difficult. A new trend arises for sensors used in<br />
systems with a standard combinations of different sensors.<br />
An example are the inertial sensors used for vehicle<br />
dynamics control systems like ESP®. A typical ESP system<br />
needs the signals of a yaw rate sensor and an one or two<br />
axial low-g acceleration sensor.<br />
The first ESP systems were using a macro-mechanical<br />
yaw rate sensor, which was based on a piezoelectrically<br />
actuated, vibrating metal cylinder with piezo’s as sensing<br />
element of the Coriolis force [1], for detection of the car´s<br />
rotation along its vertical axis. In addition a mechanical<br />
single-axis low-g accelerometer has been applied to detect<br />
the vehicle´s dynamical state and for plausibilization of the<br />
yaw rate signal.<br />
2
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May 2011, Aix-en-Provence, France<br />
<br />
Figure 7: Package roadmap of yaw rate sensors for ESP<br />
specific readout circuit (ASIC) in a SOIC16w package (Fig.<br />
6). With this approach the footprint of the sensor could be<br />
reduced by 70% to the two predecessor sensors.<br />
Figure 5: SEM picture of Bosch’s first micromechanical yaw rate sensor<br />
(combination of bulk and surface micromachining)<br />
In 1998, as ESP systems were starting to gain broader<br />
market share, Bosch introduced its first silicon<br />
micromachined yaw rate sensor [2]. The sensing elements<br />
were manufactured using a mixed bulk and surface<br />
micromachining technology and have been packaged in a<br />
metal can housing (Fig. 5).<br />
Growing demand for new additional functions of ESP and<br />
of future vehicle dynamics systems – like Hill Hold Control<br />
(HHC), Roll Over Mitigation (ROM), Electronic Active<br />
Steering, and others – required the development of<br />
improved inertial sensors with higher precision at lower<br />
manufacturing costs. These goals have been achieved by the<br />
3 rd generation ESP sensors [3], a digital inertial sensor<br />
platform based on cost effective surface micromachining<br />
technology, which was released in 2005.<br />
Fig. 7 shows the development of the first mechanical yaw<br />
rate sensor to the current combined inertial sensor SMI540<br />
REFERENCES<br />
[1] A. Reppich, R. Willig, “Yaw Rate Sensor for Vehicle Dynamics<br />
Control Systems”, SAE Technical Paper 950537 (1995).<br />
[2] M. Lutz, W. Golderer. J. Gerstenmeier, J. Marek, B. Maihöfer, S.<br />
Mahler; H. Münzel, U. Bischof, in Proceedings of Transducers '97,<br />
Chicago, IL, June 1997, p. 847-850.<br />
[3] U. Gómez, B. Kuhlmann, J. Classen, W. Bauer, C. Lang, M. Veith,<br />
E. Esch, J. Frey, F. Grabmaier, K. Offterdinger, T. Raab, R. Willig,<br />
R. Neul, “New Surface Micromachined Angular Rate Sensor for<br />
Vehicle Stabilizing Systems in Automotive Applications”, in<br />
Proceedings of Transducers ’05, Seoul, June 2005, p. 184-187.<br />
Recent development at Bosch resulted in the world’s first<br />
integrated inertial sensor modules, combining different<br />
sensors (angular rate and low-g acceleration sensors) and<br />
various sensing axis (x, y) into one single standard mold<br />
package at low size and footprint (SMI540). In detail, the<br />
sensor consists of a combination of two surface<br />
micromachined MEMS sensing chips – one for angular rate,<br />
one for 2-axis acceleration – stacked onto an application<br />
Figure 6: Combined inertial sensor SMI540 for ESP<br />
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May 2011, Aix-en-Provence, France<br />
<br />
Dynamic Behavior of Resonant Piezoelectric<br />
Cantilevers Partially Immersed in Liquid<br />
M. Maroufi 1,2 ,Sh. Zihajehzadeh 1,3 , M. Shamshirsaz 1 , A.H. Rezaie 3 , M.B. Asgari 4<br />
1 New Technologies Research Center, 2 Mechanical Engineering Department, 3 Electrical Engineering Department<br />
Amirkabir University of Technology (Tehran Polytechnic), 4 Niroo Research Institute<br />
424 Hafez Ave., P.B. 15875-4413. Tehran, Iran<br />
E-mail: shamshir@aut.ac.ir<br />
Abstract<br />
Resonant Piezoelectric-excited Millimeter-sized Cantilevers<br />
(PEMC), has attracted many researchers' interest in the<br />
applications such as liquid level and density sensing. As in<br />
these applications, the PEMC are partially immersed in liquid,<br />
an appropriate analytical model is needed to predict the<br />
dynamic behavior of these devices.<br />
In this work, a PEMC has been fabricated for liquid level<br />
sensing. An analytical model based on Euler-Bernoulli theory<br />
and energy method is developed and applied to evaluate the<br />
performance of this device with respect to different tip<br />
immersion depth. To validate this model, the theoretical<br />
results are compared with the experimental results for the tip<br />
immersion depth from 0.5 mm to 9 mm in water. The<br />
simulation results are in almost good agreement with<br />
experimental data. The difference in natural frequency<br />
obtained by the theoretical model for different immersion<br />
depth remains less than 8%. The linear region of the natural<br />
frequency shift versus immersion depth has been identified to<br />
be from the depth of 9 to 11 mm.<br />
I. INTRODUCTION<br />
Nowadays, resonant Piezoelectric-excited Millimeter-sized<br />
Cantilevers (PEMC) have many applications as sensors.<br />
Among these diverse applications, are the ones where the<br />
cantilever is partially immersed in the liquid environment. In<br />
these cases, PEMC are used for online measuring of liquid<br />
density [1], [2], [3] or online determination of liquid level at<br />
micron resolution [4]. <strong>Online</strong> level detection of liquid is a<br />
powerful tool in many analytical processes where solvent<br />
concentration has to be monitored.<br />
Even though, there exists different tools for liquid level<br />
sensing such as ultrasonic, acoustic and optical methods, none<br />
of them is competent with PEMC, considering their ease of<br />
fabrication, small size and high performance [4].In fact, the<br />
performance of these devices for sensing application in liquid<br />
environment depends on many factors such as dimension of<br />
the cantilever and the piezoelectric layer, the immersion depth<br />
of the cantilever into liquid and so on.<br />
To evaluate the performance of PEMC partially immersed in<br />
liquid, a theoretical model is needed. Analytical model for the<br />
piezoelectric driven macro cantilever in air in introduced in [5]<br />
and also a model for the thermal driven cantilever wholly<br />
immersed in liquid with application in AFM is presented in<br />
[6].<br />
In this work, a PEMC has been fabricated for liquid level<br />
sensing. The motivation is first to develop an analytical model<br />
to predict the dynamic behavior of PEMC partially immersed<br />
in liquid. This model is derived here based on Euler-Bernoulli<br />
theory and energy method. Further, this model could be<br />
utilized to investigate the effect of the different geometrical<br />
and material properties on the performance of these devices as<br />
future work.<br />
Second objective in this work is to identify the appropriate<br />
immersion depth range in which the resonant frequency<br />
changes due to immersion depth variation show a linear<br />
behavior in liquid level sensor application.<br />
To validate this model, the theoretical results are compared<br />
with the experimental results for the tip immersion depth from<br />
0.5 mm to 9 mm in water. The simulation results are in almost<br />
good agreement with experimental data. The difference in<br />
natural frequency obtained by the theoretical model for<br />
different immersion depth remains less than 8%. The linear<br />
region of the natural frequency shift versus immersion depth<br />
has been identified to be from the depth of 9 to 11 mm.<br />
II. THEORETICAL MODEL<br />
The fabricated PEMC is depicted schematically in Fig. 1. This<br />
structure consists of a millimeter sized steel beam as a<br />
cantilever on which a piezoelectric patch is attached. The<br />
cantilever is immersed partially in the fluid. Applying<br />
electrical AC voltage on the piezoelectric patch, the cantilever<br />
is forced to vibrate.<br />
To model the resonant cantilever partially immersed in liquid ,<br />
three regions on the cantilever has been considered; a) first<br />
part where piezoelectric patch is bonded on the cantilever, b)<br />
middle part of cantilever where it vibrates freely ignoring air<br />
damping effect c) end part where the cantilever vibrates in the<br />
liquid. Also, three coordinate systems are assumed in each<br />
region (Fig. 1).<br />
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<br />
(3)<br />
In (3), d is the piezoelectric constant. To obtain E along<br />
piezoelectric thickness, it is assumed that an AC electrical<br />
voltage V is applied to piezoelectric patch. Using piezoelectric<br />
constitutive equation, the electrical field can be calculated by<br />
[5]:<br />
2 2 2 <br />
<br />
2 <br />
(4)<br />
Fig.1.Configuration of the PEMC partially immersed in liquid; three<br />
regions considered in theoretical model<br />
To achieve vibration equation of the cantilever accompany by<br />
piezoelectric patch the strain distribution along the thickness<br />
of the cantilever must be obtained. In Fig.2 the neutral axis<br />
position is demonstrated. In this figure the distance between<br />
neutral axis of the cantilever-piezoelectric from piezoelectric<br />
bottom layer is denoted by . It is assumed that the<br />
distribution of the strain along cantilever thickness is linear, so<br />
the strain at a distance x from the neutral axis is:<br />
<br />
<br />
(1)<br />
In which w is the transverse displacements of the cantilever.<br />
(w ) denotes twice derivation with respect to X 1 along the<br />
cantilever.<br />
In which t is the thickness of the piezoelectric. After<br />
Substitution (4) and (1) in (2), the energy of the piezoelectric<br />
layer can be obtained. To drive vibration equation, the<br />
Lagrangian for the piezoelectric layer and cantilever is<br />
calculated. So, the kinetic energy has to be solved. The mass<br />
per length for the first region of the piezoelectric cantilever in<br />
the kinetic energy calculation is defined as:<br />
(5)<br />
After calculation of the Lagrangian for both layers, the<br />
variation of the (6) is set to zero:<br />
<br />
<br />
0 (6)<br />
<br />
<br />
In which are the Lagrangian for each region. Using (6), the<br />
equation of motion and the boundary condition of the first<br />
region are determined. The equation of the motion is:<br />
2 2 0 (7)<br />
In (7), I , I are the moment of inertia of the piezoelectric<br />
patch and cantilever with respect to neutral axis respectively.<br />
A is the piezoelectric cross section area and Y is the Young<br />
modulus of the cantilever. a and a are defined as [5]:<br />
1 2 <br />
<br />
(8)<br />
Fig.2:The position of the neutral axis with respect to piezoelectric patch<br />
and the coordinate X 3 along the cantilever thickness<br />
<br />
2 <br />
8<br />
(9)<br />
To drive vibration equation of the first region of the cantilever,<br />
the energy stored in the piezoelectric can be determined by<br />
[5]:<br />
1 2 1 2 (2)<br />
Where Y is the Young modulus of the piezoelectric layer, E <br />
is the electrical filed along piezoelectric layer thickness due to<br />
applied voltage, is the dielectric constant. e is defined as:<br />
Using energy method also for the second region, the equation<br />
of the motion becomes [6]:<br />
0 (10)<br />
In the above equation, (w ) is the transverse displacement of<br />
the cantilever in the second region. To obtain solution in<br />
frequency domain, both (7) and (10) should be rewritten in<br />
frequency domain. (11) and (12) are the frequency domain<br />
expression of the (7) and(10) respectively:<br />
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May 2011, Aix-en-Provence, France<br />
<br />
, 2 2 <br />
(11)<br />
, 0<br />
, TABLE 1<br />
, 0 (12)<br />
TABLE OF MATERIAL AND GEOMETRICAL PROPERTIES<br />
Property<br />
Cantilever<br />
Piezoelectric<br />
(Steel St304) (PZT5H4E)<br />
In the third region the vibration of the cantilever is affected by<br />
Young Modulus (GPa)<br />
the liquid presence. The force which is exerted by liquid to a<br />
193 62<br />
long vibrant cantilever in frequency domain is given in [7] Density (Kg/m 3 ) 8000 7800<br />
as:<br />
Length (mm) 36 6<br />
Width (mm) 3 3<br />
4 Γ , <br />
(13)<br />
Thickness(mm) 0.1 0.267<br />
Piezoelectric<br />
In (13), W is the transverse displacement of<br />
the cantilever in<br />
--- --- -320×10<br />
constant(m/V)<br />
-12<br />
the third region in frequency domain, ρ F is the density of fluid<br />
and b is the width of the cantilever. Γ is the hydrodynamic Relative Dielectric<br />
--- --- 3800<br />
function considering the viscosity and density of the displaced<br />
constant<br />
liquid [7]. Regarding the liquid force on the vibration of the<br />
cantilever, equation of the motion of the cantilever in<br />
The geometrical and material properties of the PEMC and<br />
frequency domain can be presented by:<br />
piezoelectric patch are given in TABLE 1.<br />
The natural frequency of the piezoelectric cantilever has been<br />
, , (14) determined by an impedance evaluation board, where the<br />
impedance phase angle attaints the maximum [4]. The<br />
To obtain the frequency response of the cantilever three experiments have been carried out to obtain natural<br />
equations developed above are solved simultaneously with frequencies of the PEMC for different immersion depth in<br />
appropriate boundary conditions.<br />
liquid. To vary this immersion depth, for each test, a known<br />
volume of water equivalent to 500μm liquid level change, is<br />
III. EXPERIMENTAL SETUP AND PROCEDURE<br />
added to container.<br />
The schematic of the experimental setup for liquid level<br />
sensing is shown in Fig. 3. The PEMC is mounted on a holder<br />
IV. RESULTS AND DISCUSSION<br />
which keeps PEMC at a fixed position in the<br />
liquid container<br />
during the experiments.<br />
The experimental and theoretical natural frequencies for each<br />
For the fabrication of the millimeter size cantilever, Electrical immersion depth in water are shown in Tab 2. As it can be<br />
Discharge Machining (EDM) is used to achieve the seen, the natural frequency decreases as the liquid level<br />
dimensional tolerances below millimeter. The piezoelectric increases. This decrease in natural frequency is due to a<br />
layer is cut by diamond knife and is attached to cantilever by greater displaced mass of liquid<br />
with the immersed cantilever<br />
cyanoacrylate adhesive. The schematic of the PEMC is shown part.<br />
in Fig. 4.<br />
The deviation percentage of theoretical results from<br />
experimental data defined as <br />
100 is also reported<br />
<br />
in this table. This deviation remains less than about 8% for<br />
different immersion depth.<br />
To examine the performance of<br />
the PEMC as a liquid level<br />
sensor, the curve of natural frequency shift versus immersion<br />
depth is presented in Fig. 5. In this figure, the middle region;<br />
from 9 to 11 mm immersion depths, not only the curve is<br />
linear but also it has the highest slope.<br />
Fig.3.Schematic of experimental setup for liquid level sensing<br />
Fig.4.Schematic of the PEMC<br />
TABLE 2<br />
COMPARISION OF THEORITICAL AND EXPERIMENTAL RESULTS IN WATER<br />
Immersion<br />
<br />
depth (mm)<br />
(Hz)<br />
6 4211<br />
7 4193<br />
8 4184<br />
9 4133<br />
10 4009<br />
11 3850<br />
12 3744<br />
13 3710<br />
14 3702<br />
(Hz)<br />
Deviation<br />
(%)<br />
4095.5 2.73<br />
4087 2.53<br />
4015 4.04<br />
3850 6.85<br />
3677.5 8.27<br />
3577 7.09<br />
3553 5.1<br />
3549 4.34<br />
3502 5.402<br />
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May 2011, Aix-en-Provence, France<br />
<br />
Fig.5.Experimental and theoretical natural frequency shift vs. immersion<br />
depth for water<br />
The sensor should be utilized in the region with the highest<br />
slope or highest sensitivity i.e. the highest frequency shift for<br />
each increment change in the immersion depth.<br />
It can be seen from Fig.5 that even though there is a slight<br />
difference between the theoretical and experimental curve, as<br />
the trend of the curves are similar, the dynamic behavior of the<br />
PEMC in liquid can be satisfactorily predicted by this<br />
theoretical model. The slight difference of theoretical results<br />
from the experimental data; a positive vertical shift<br />
accompanied with a negative horizontal shift, can be described<br />
as follow. First, theoretical model has been developed based<br />
on the assumptions for the simplicity of calculation such as the<br />
cantilever length and the container dimensions are assumed<br />
too long comparing with cantilever width and thickness, …[7].<br />
Moreover, some of the mechanical and dimensional<br />
parameters have been ignored due to lack of measurement<br />
parameters. These parameters are the thickness of the adhesive<br />
layer and its mechanical properties, the effect of the clamp and<br />
the dissipation factor in piezoelectric and cantilever. Also, the<br />
values of the steel cantilever and the piezoelectric material<br />
properties such as the Young modules, density …are provided<br />
from literature, so there exists some uncertainties in<br />
parameters' values given in Tab. 1. Furthermore, in the<br />
experimental test there can be some errors in determining the<br />
exact volume of added fluid, and consequently in determining<br />
the immersion depth exactly.<br />
REFERENCES<br />
[1] Kishan Rijal, Raj Mutharasan, “Piezoelectric-excited millimetersized<br />
cantilever sensors detect density differences of a few<br />
micrograms/mL in liquid medium”, Sensors and Actuators B 124<br />
(2007) 237--244<br />
[2] Christian Riesch, Erwin K. Reichel,Franz Keplinger, and Bernhard<br />
Jakoby, “Characterizing Vibrating Cantilevers for Liquid Viscosity<br />
and Density Sensing”, Hindawi <strong>Publishing</strong> Corporation Journal of<br />
Sensors Volume 2008, Article ID 697062, 9 pages<br />
doi:10.1155/2008/697062<br />
[3] Wan Y. Shih,Xiaoping Li, Huiming Gu, Wei-Heng Shih and Ilhan<br />
A. Aksay,” Simultaneous liquid viscosity and density<br />
determination with piezoelectric unimorph cantilevers”, Journal of<br />
Applied Physics, Volume 89, Number 2, January 2001<br />
[4] Gossett A. Campbell, Raj Mutharasan,” Sensing of liquid level at<br />
micron resolution using self-excited millimeter-sized PZT<br />
cantilever”, Sensors and Actuators A 122 (2005) 326–334<br />
[5] Sudipta Basak, Arvind Raman, Suresh V. Garimella, “Dynamic<br />
Response Optimization of Piezoelectrically Excited Thin Resonant<br />
Beams”, Journal of Vibration and Acoustics FEBRUARY 2005<br />
Vol. 127 / 19<br />
[6] Singiresu .S. Rao, “ Vibration of continuous systems”, Wily 2007,<br />
PP. 321-338.<br />
[7] John Elie Sader, “Frequency response of cantilever beams<br />
immersed in viscous fluids with applications to the atomic force<br />
microscope”, Journal of Applied Physics, Volume 84, Number 1,<br />
July 1998<br />
[8] K. Fukuda, H. Irihama, T. Tsuji, K. Nakamoto, K. Yamanaka,<br />
Sharpening contact resonance spectra in UAFM using Q-control,<br />
Surf. Sci. 532, 535 (2003) 1145–1151.<br />
V. CONCLUSION<br />
A PEMC with a test set-up have been fabricated for liquid<br />
level sensing. An analytical model to predict the dynamic<br />
behavior of partially immersed PEMC in liquid environment is<br />
developed. The validity of the model is examined by<br />
comparison of simulation results with the experimental data<br />
for different immersion depth of the PEMC in water. The<br />
difference in natural frequency obtained by the theoretical<br />
model for different immersion depth remains less than 8%. A<br />
linear region for sensing related to immersion depth from 9 to<br />
11 mm is identified where the sensitivity is maximum.<br />
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<br />
Reliable system-level models for electrostatically actuated<br />
devices under varying ambient conditions:<br />
Modeling and validation<br />
Gabriele Schrag, Martin Niessner, Gerhard Wachutka<br />
Institute for Physics of Electrotechnology, Munich University of Technology<br />
Arcisstr. 21, D-80290 München, Germany<br />
email: schrag@tep.ei.tum.de<br />
Abstract- We present a physics-based multi-energy domain<br />
macromodel that allows – in general – for the efficient simulation<br />
of any electrostatic actuator within standard IC frameworks and<br />
apply it exemplarily to an RF-MEMS switch. The predictive<br />
power of this macromodel, which depends crucially on the<br />
quality of the applied damping and contact models, has been<br />
evaluated by white light interferometer and laser vibrometer<br />
measurements. It turned out that the models for viscous damping<br />
as well as for the electromechanical energy domain are in very<br />
good agreement with the experiments while the applied standard<br />
contact model fails in reproducing the measured contact<br />
phenomena. Based on these findings suggestions for improved<br />
system-level contact models are discussed.<br />
investigation of the performance of the models for varying<br />
pressure conditions and during the phase of initial contact,<br />
since these phenomena have decisive impact on the closing<br />
behavior of the considered switch and all MEMS actuators<br />
operating in contact mode.<br />
I. MOTIVATION AND PROBLEM DESCRIPTION<br />
A key prerequisite for the routine use of<br />
microelectromechanical actuators like radio frequency (RF-<br />
MEMS) switches, e.g., as standard circuit elements is the<br />
availability of computationally efficient, but yet physics-based<br />
and, thus, predictive models, which correctly describe their<br />
operation. Furthermore, these models should be compatible<br />
with a framework that allows for an integrated design of<br />
semiconductor-based circuits with MEMS hybridization. The<br />
simulation of the switching behavior, i.e. of the pull-in and<br />
pull-out transients of such devices, is, however, a challenging<br />
task because multiple energy domains and their nonlinear<br />
interactions have to be taken into account, i.e. the electrostatic<br />
actuation of the mechanically moving parts, viscous air<br />
damping and contact forces during impact. The preferred<br />
method for enabling the fast simulation of such pull-in/-out<br />
transients is therefore not the use of complex and<br />
computationally expensive finite element models but of multienergy<br />
domain coupled macromodels with a highly reduced<br />
number of degrees of freedom that are by far more efficient and<br />
can be simulated within standard integrated circuit (IC) design<br />
frameworks.<br />
In the following, we present physics-based macromodels<br />
suited, in general, for the design of electrostatically actuated<br />
and viscously damped actuators which operate under dynamic<br />
pull-in conditions and apply it to an RF-MEMS switch. The<br />
derived models, which can be directly used for co-simulation<br />
with electronic circuits in standard IC design frameworks, are<br />
evaluated w.r.t. measurements carried out with a white light<br />
interferometer (WLI) and a laser Doppler vibrometer (LV),<br />
respectively. Special emphasis has been placed on the<br />
Figure 1. Measured (WLI) 3D profile of the RF switch without bias.<br />
Figure 2. Measured (WLI) profile of the electrodes and the 12 elevated<br />
contact pads. The membrane was removed for this measurement.<br />
II.<br />
contact<br />
pads<br />
DEMONSTRATOR AND EXPERIMENTAL SETUP<br />
A 3D white light interferometer profile of the considered<br />
RF-MEMS switch is depicted in fig. 1. The switch has been<br />
fabricated at Fondazione Bruno Kessler (FBK) in Trento [1]<br />
and consists of a movable perforated gold membrane<br />
suspended above a fixed ground electrode through four straight<br />
beams. The fixed ground electrode acts as actuation electrode<br />
of the switch and consists of several lateral fingers that are<br />
connected in parallel (cp. fig. 2). By applying a voltage, the<br />
8
suspended membrane can be pulled towards the ground<br />
electrode, collapses onto 12 elevated contact pads and closes an<br />
ohmic contact so that the RF signal path is closed.<br />
The topography of the switches has been analyzed by applying<br />
a white light interferometer (Veeco NT1100 DMEMS) and the<br />
dynamics has been characterized by recording the transient<br />
deflection of the moving membrane by a single spot laser<br />
Doppler vibrometer (Polytec OFV-5000). An on-purpose<br />
developed vacuum chamber with pressure control enables the<br />
characterization of the microstructures under varying pressure<br />
conditions in order to evaluate the applied models for viscous<br />
damping. The experimental set up is shown in fig. 3.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
using the single point laser Doppler vibrometer depicted in<br />
fig. 3, parameters like the pull-in/pull-out voltages (and from<br />
that the actual gap height) or the resonance frequency of the<br />
mechanical structure could be extracted. The parameters of the<br />
investigated switch, which are the basis of our models, are<br />
summarized in detail in table 1 below.<br />
Table 1. Technical data of the investigated RF-MEMS switches. For the<br />
electrode and the substrate, the gap width is given between the membrane and<br />
the dielectric layers.<br />
Membrane<br />
Suspensions<br />
Thickness 5.2 µm Thickness 2.0 µm<br />
Length 260 µm Length 165 µm<br />
Width 140 µm Width 10 µm<br />
A<br />
Side length of<br />
holes<br />
Spacing between<br />
holes<br />
20 µm<br />
20 µm Other<br />
Resonance frequency<br />
14.7 kHz<br />
C<br />
B<br />
C<br />
Gap widths<br />
Membrane to<br />
contact pad<br />
Membrane to<br />
electrode<br />
Membrane to<br />
substrate<br />
Thickness of 700 nm<br />
dielectric on<br />
electrode<br />
1.7 µm Pull-in voltage 29-30 V<br />
2.7 µm Release voltage 22-26 V<br />
3.4 µm Effective residual<br />
air gap (g min )<br />
~20-50 nm<br />
Figure 3. Photograph of the laser vibrometer (A) and the on-purpose<br />
developed pressure chamber (B). Two pressure sensors (C) are used to<br />
control the pressure inside the chamber .<br />
Displacement [μm]<br />
0.4<br />
0.2<br />
0<br />
-0.2<br />
-0.4<br />
-0.6<br />
-0.8<br />
-1<br />
-1.2<br />
-1.4<br />
-1.6<br />
-1.8<br />
-30 -20 -10 0 10 20 30<br />
Voltage [V]<br />
Figure 4. Quais-static pull-in/pull-out characteristic of the RF MEMS<br />
switch. A trinangular waveform with zero mean voltage and 70 V amplitude<br />
(peak to peak) at a frequency of 1 Hz has been applied. The pull-in voltage<br />
lies between 29 V and 30 V, the release voltage between 23 V and 26 V.<br />
As a first guess, the parameters of the switch have been taken<br />
from the technical data given by the process description and<br />
the design. In order to include also the manufacturing<br />
tolerances in our model and, thus, to enhance its accuracy,<br />
optical measurements appling a white ligth interferometer (see<br />
fig. 1 and 2) have been carried out in order to extract the exact<br />
dimensions of the device (electrode, contact pads, membrane<br />
thickness, e.g.). From the quasi-statically measured pull-in and<br />
pull-out characteristics (see fig. 4) and dynamic measurements<br />
III. MODELING AND THEORETICAL BACKGROUND<br />
The macromodel of the switch is derived on the basis of the<br />
hierarchical modeling approach as reported in [2], which is<br />
strictly based on flux-conserving reduced-order and/or compact<br />
modeling techniques, so that the resulting system-level models<br />
are rigorously formulated in terms of conjugated variables<br />
(”across”- and ”through”-variables) and the generalized<br />
Kirchhoffian network theory can be used as a theoretical<br />
framework for the formulation of the entire system model.<br />
Starting point of the modeling procedure is the decomposition<br />
of the device into tractable subsystems. In this particular case,<br />
these are the mechanical subsystem represented by the<br />
perforated membrane and the four flat suspension springs, the<br />
electrostatic subsystem, accounting for the electric field<br />
between the perforated membrane and the actuation electrode<br />
(see Fig. 2), and the fluidic subsystem comprising the ambient<br />
air that exerts damping forces on the moving parts of the<br />
structure. Additionally, adequate compact models have to be<br />
derived that describe the closing phase of the switch properly.<br />
The basis for the mechanical submodel of the suspended<br />
membrane is the modal superposition technique described in<br />
[3]. The eigenmode shapes and frequencies of the suspended<br />
membrane are calculated in a FEM simulation tool. The most<br />
significant modes – in the case of the considered switch the<br />
fundamental and the next higher completely symmetric<br />
eigenmode – are identified and used to formulate a<br />
macromodel in terms of modal amplitudes consisting of only<br />
one second-order differential equation per included eigenmode.<br />
9
11-13 <br />
May 2011, Aix-en-Provence, France<br />
Residual stress in the suspended membrane induced by the<br />
<br />
Here, q denotes the vector of modal amplitudes, φ<br />
pad , n<br />
the<br />
fabrication process has been taken into account by calibrating<br />
the fundamental eigenfrequency to the measured one.<br />
averaged modal shape factor for the n-th contact pad,<br />
The submodel for the electrostatic forces exerted by the<br />
ground electrode is derived in two steps. First, the electrostatic<br />
energy, which is stored between a single electrode finger and<br />
the membrane, is determined in terms of the modal amplitudes.<br />
Second, Lagrangian energy functionals are calculated for each<br />
eigenmode and are included as electrostatic actuation term in<br />
the respective eigenmode equation of the mechanical model.<br />
In order to take into account the viscous damping forces the<br />
mixed-level approach as presented in [4] is applied. It is based<br />
on the Reynolds equation, which is evaluated by applying a<br />
fluidic Kirchhoffian network distributed over the device<br />
geometry. At perforations and outer boundaries lumped<br />
physics-based fluidic resistances are added accounting for the<br />
additional pressure drops at these locations (see fig. 5).<br />
Consequently, this mixed-level model does not constitute a<br />
pure lumped element model and – depending on the granularity<br />
of the finite network – still exhibits a rather larger number of<br />
degrees of freedom. However, the advantage of this approach is<br />
that it can be tailored to the topography of the real structure, i.e.<br />
take into account all perforations and – in the case of the<br />
considered switch – also locally varying gap heights which<br />
occur due to the elevated contact pads and electrode fingers.<br />
ambient pressure<br />
P 0<br />
moving<br />
plate<br />
{<br />
finite network<br />
R O<br />
R C<br />
R T<br />
fixed plate<br />
moving<br />
plate<br />
{<br />
finite network<br />
Figure 5. Illustration of the mixed-level model. Models for holes are<br />
embedded in the finite network solving the Reynolds equation. The resistor R T<br />
models the region, where the fluid enters the channel. The resistor R C models<br />
the channel resistance; R O models the orifice flow [5,6].<br />
The mechanical contact that occurs, when the switch is<br />
closed, is included into the mixed-level model by adding<br />
contact forces at the respective locations above the contact<br />
pads. The modal formulation of these forces reads as follows:<br />
12<br />
⎧<br />
⎪<br />
( )<br />
∑ φpad , n<br />
⋅kcontact , n<br />
⋅gn ( q) if gn<br />
( q)<br />
≤0<br />
Fcontact, total,<br />
i<br />
q = ⎨ (1)<br />
n=<br />
1<br />
⎪⎩ 0 else<br />
k<br />
contact,<br />
n<br />
the lumped contact stiffness of the n-th pad and gn<br />
( )<br />
q the<br />
locally averaged displacement at the n-th pad.<br />
This model enables to simulate also bouncing during the<br />
landing phase of the membrane. In order to avoid numerically<br />
undesired discontinuities resulting from the if-then-else<br />
construct proposed in previous work [7], we now use a<br />
function Θ<br />
n<br />
based on the tanh-function instead, in order to<br />
implement a more stable transition into the contact state:<br />
⎛ ⎛ gn<br />
Θ<br />
n ( q)<br />
= 0.5⋅⎜1−tanh<br />
⎜<br />
⎜ ⎜ β<br />
⎝ ⎝<br />
( q)<br />
⎞⎞<br />
⎟⎟<br />
⎟⎟<br />
⎠⎠<br />
β denotes a parameter controlling the smoothness of this<br />
transition. The complete contact formulation then reads:<br />
12<br />
( ) = Θ ( ) ⋅φ<br />
⋅ ⋅ ( )<br />
F q ∑ q k g q (3)<br />
contact , total , i n pad , n contact , n n<br />
n=<br />
1<br />
The compact model of the entire switch is then assembled<br />
by formulating all submodels in terms of the modal amplitudes<br />
and combining them with the mechanical submodel:<br />
2 7<br />
2 Vb<br />
∂Ck( q)<br />
T<br />
i<br />
+ ωi i<br />
= ∑ + θi 2 k = 1 ∂qi<br />
<br />
ext, i<br />
<br />
0<br />
F el<br />
q<br />
q F ( qq , , p)<br />
Here, q i and ω i denote the amplitude and the frequency of<br />
the i-th eigenmode, θ i<br />
denotes the vector of the respective<br />
discretized mode shape, Ck<br />
( q ) stands for the capacitance of<br />
the k-th electrode finger and V k for the respective applied<br />
voltage. F ext represents the vector of external forces comprising<br />
in this case the models for damping and contact forces.<br />
Finally, the derived macromodels of the subsystems are<br />
formulated in terms of conjugated variables (”across”- and<br />
”through”-variables) and interlinked to form a generalized<br />
Kirchhoffian network, which inherently governs the exchange<br />
of energy and other physical quantities through Kirchhoff’s<br />
laws and can be implemented easily in any standard system<br />
simulator of an IC framework (in this work: Spectre from the<br />
Cadence IC design suite).<br />
IV. EXPERIMENTAL VALIDATION OF SIMULATED RESULTS<br />
The macromodel was evaluated w.r.t. measurements<br />
performed with a laser Doppler vibrometer (LV) and a white<br />
light interferometer (WLI). A pressure chamber as depicted in<br />
fig. 3 was used to enable measurements at different pressure<br />
levels.<br />
First, the combined electro-mechanical model was validated<br />
against the quasi-statically measured pull-in/pull-out<br />
characteristic of the membrane (see Fig. 6). It shows good<br />
agreement with the pull-in characteristic, but yields an<br />
incorrect pull-out voltage. Since the electromechanical model<br />
works quite accurately for the pull-in curve, this discrepancy is<br />
(2)<br />
(4)<br />
10
most likely due to not yet considered adhesion forces and other<br />
contact-related phenomena.<br />
Displacement [μm]<br />
0.4<br />
0.2<br />
0<br />
-0.2<br />
-0.4<br />
-0.6<br />
-0.8<br />
-1<br />
-1.2<br />
Measurement<br />
-1.4<br />
MLM<br />
-1.6<br />
-1.8<br />
0 5 10 15 20 25 30<br />
Voltage [V]<br />
Figure 6. Measured and simulated quasi-static pull-in/-out characteristics of<br />
the membrane. The curves have been taken using the laser vibrometer and<br />
actuating the membrane electrostatically with a voltage of 70 V (peak to peak)<br />
of triangular wave form at a frequency of 1 Hz.<br />
Second, we evaluated the transients of the device at an<br />
ambient pressure ranging from 960 mbar to 1 mbar. Figure 7<br />
shows the measured and simulated response of the switch to a<br />
square wave voltage of 25 V, a voltage which is lower than the<br />
pull-in voltage. The membrane responds in the first half of the<br />
period (0..2.5ms) with a damped oscillation at a mean<br />
displacement of about 420nm. After voltage has been turned<br />
off (t > 2.5ms), it releases to its original position undergoing<br />
damped oscillations.<br />
0.4<br />
0.2<br />
Measurement<br />
MLM<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Additionally, we evaluated the limits of the fluidic damping<br />
model by extracting the quality factor Q as a measure for<br />
viscous damping from the transients recorded for varying<br />
ambient pressure values. Since for pressures lower than<br />
100 mbar the Q-factor is limited by another mechanism than<br />
squeeze film damping, we extracted the value Q LIMIT for this<br />
damping mechanism from experiments and added it to the Q-<br />
factor Q SQFD calculated from our mixed-level damping model<br />
according to<br />
Q TOTAL -1 = Q LIMIT<br />
-1<br />
+ Q SQFD<br />
-1<br />
Fig. 8 reveals that our model is in very good agreement –<br />
within an error of about 7% to 10% – to the measured data for<br />
moderately high pressure values from normal pressure down to<br />
about 100 mbar. Below pressures of about 100 mbar a 15%<br />
error limit between simulated and measured Q-values is<br />
exceeded. This is certainly due to the large variance of the<br />
correction factors accounting for gas rarefaction effects, which<br />
can be found in literature. Systematic investigations on<br />
dedicated test structures are focus of on-going work in order to<br />
get a better data base for the physical understanding of gas<br />
damping in this regime [9].<br />
Quality factor<br />
Q MLM<br />
Q MEAS<br />
(5)<br />
10 2 Pressure [mbar]<br />
Displacement [μm]<br />
0<br />
-0.2<br />
-0.4<br />
-0.6<br />
-0.8<br />
0 1 2 3 4 5<br />
Time [ms]<br />
Figure 7. Measured and simulated response of the membrane.<br />
Actuation: rectangular voltage waveform (200 Hz; amplitudes 25<br />
V (on) and 0 V (off)). Ambient pressure: 960 mbar.<br />
The very good agreement of simulation and measurement<br />
in the second half of the time period (2.5..5ms) proves the<br />
accuracy of the damping model, while the good agreement in<br />
the first half of the period proves that the model correctly<br />
reproduces the electrostatic spring softening, which essentially<br />
decreases the resonance frequency of the switch during<br />
actuation, and the increased damping due to the decreasing gap<br />
height. Models of switches without physically-based<br />
description of gas damping fail at this point [8].<br />
10 1 10 0 10 1 10 2 10 3<br />
Figure 8. Simulated and measured Q values calculated from the frequency<br />
spectrum using the half power method (“3dB bandwidth”). Remark: The<br />
measured sample exhibited an increased gap (plus about 300 nm ).<br />
In order to check the contact model, we actuated the<br />
device with a step voltage of 35 V, a voltage higher than the<br />
pull-in voltage (see fig. 9), in order to force the structure into<br />
contact and, subsequently, to release it again.<br />
Displacement [μm]<br />
1.4<br />
1<br />
0.6<br />
0.2<br />
-0.2<br />
-0.6<br />
-1<br />
-1.4<br />
Measurement<br />
MLM<br />
-1.8<br />
0 1 2 3 4 5<br />
Time [ms]<br />
Figure 9. Response of the membrane to a rectangular waveform (200Hz):<br />
amplitudes 35V (on) and 0V (off). Pull-in/contact occurs.<br />
11
A detailed analysis of the frequency spectrum of the<br />
landing phase (fig. 10) as well as of the release phase (with and<br />
without contact, i.e. actuation voltages of 35 V and 25 V,<br />
respectively, fig. 11) reveal that these phases are dominated by<br />
an intricate interplay of different mechanical vibrations.<br />
Peak-normalized<br />
amplidute (dB)<br />
Peak-normalized<br />
amplitude (db)<br />
0<br />
-10<br />
-20<br />
-30<br />
-40<br />
-50<br />
87 kHz 218 kHz<br />
-60<br />
0 50 100 150 200 250 300<br />
Frequency (kHz)<br />
Figure 10. Frequency spectrum of the measured landing phase of the<br />
membrane. Applied voltage 35V, frequency 250Hz.<br />
0<br />
-10<br />
-20<br />
-30<br />
-40<br />
-50<br />
14.7 kHz<br />
136 kHz<br />
35V (after pull-in)<br />
25V (no pull-in)<br />
-60<br />
0 50 100 150 200 250 300<br />
Frequency (kHz)<br />
Figure 11. Frequency spectrum of the measured release phase of the<br />
membrane. Applied voltage 35V and 25V, resp.; frequency 250Hz.<br />
The modes at 14.7 kHz and 136 kHz occurring during<br />
release after the membrane was in contact with the contact pads<br />
(see black curve in fig. 11) correspond to the natural<br />
eigenfrequencies of the membrane. The notable contribution of<br />
the mode at 136 kHz compared to the spectrum where no pullin<br />
occurred (blue curve in fig. 11) leads us to the assumption<br />
that kinetic energy of the fundamental mode is transferred to<br />
the next higher symmetric mode during impact. Analyzing the<br />
landing phase of the switch by FFT (fig. 10) gives evidence of<br />
several superimposed vibrations at higher frequencies (see two<br />
modes at 87 and 218 kHz), which are obviously involved in<br />
the contact process. The high frequencies are supposed to be<br />
due to the high contact stiffness which now couples with the<br />
stiffness of the suspended membrane. Thus, the contact<br />
physics can only be captured correctly, when multiple and also<br />
higher modes and the coupling between them are implemented<br />
in the macromodel. Fig. 12 (top) shows a zoom of the closing<br />
transient displayed in fig. 9, where only the first 100μs are<br />
shown. It can be observed that our simulation model captures<br />
the landing phase already remarkably well, while models of<br />
Iannacci [10] and the commercial simulation tool Architect3D<br />
[11] that have been applied for benchmarking cannot<br />
reproduce the landing phase as well and, in particular, the<br />
Architect3D model overestimates the closing time<br />
considerably (see fig. 12, bottom). We assume that the good<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
agreement of our model is due to the fact that the mechanical<br />
part of our model, by its nature, already includes different (and<br />
also higher) mechanical eigenmodes.<br />
Displacement [μm]<br />
Displacement [μm]<br />
0<br />
-0.2<br />
-0.4<br />
-0.6<br />
-0.8<br />
-1<br />
-1.2<br />
-1.4<br />
-1.6<br />
Measurement<br />
MLM<br />
-1.8<br />
0 10 20 30 40 50 60<br />
Time [μs]<br />
0<br />
-0.2<br />
-0.4<br />
-0.6<br />
-0.8<br />
-1<br />
-1.2<br />
-1.4<br />
-1.6<br />
Measurement<br />
Architect3D<br />
Iannacci<br />
-1.8<br />
0 20 40 60 80 100<br />
Time [μs]<br />
Figure 12. Top: Zoom of the closing transient of fig. 9. Only the first 60μs are<br />
shown. The measurement is compared with simulated data from the MLM<br />
model. Bottom: Zoom of the closing transient of fig. 9. Only the first 100μs<br />
are shown and compared to two alternative macromodels (Iannacci[10] and<br />
Architect3D[11]), which have been applied for benachmarking.<br />
V. CONCLUSIONS<br />
We presented a physics-based multi-energy domain<br />
macromodel for an electrostatically actuated RF MEMS switch<br />
working under ambient pressure. It shows very good agreement<br />
with the measured quasi-static pull-in characteristic, with the<br />
non-contact transient measurements for ambient pressures<br />
down to 100 mbar and the pull-out transient after contact.<br />
However, the macromodel fails in reproducing the quasi-static<br />
pull-out characteristic and the contact phase properly. A FFT of<br />
the measured transients revealed that during impact of the<br />
membrane multiple structural modes are involved so that this<br />
phase can only be captured correctly by taking them and their<br />
interactions properly into account. Obviously, standard contact<br />
models are not yet accurate enough to reproduce this phase<br />
correctly. However, it showed that a mechanical model based<br />
on modal superposition techniques, i.e. a model where, by its<br />
nature, several eigenmodes are included, reproduces the<br />
landing phase best compared to other standard models. A<br />
sensitivity analysis showed that, together with the proper<br />
12
11-13 <br />
May 2011, Aix-en-Provence, France<br />
description of the contact phase, accurate, physics-based<br />
<br />
damping models are a prerequisite for reliable and predictive<br />
modeling of MEMS actuators. An on-going comparative study<br />
given in [9] demonstrates this impressively. Future research<br />
will therefore focus on the physics during the impact of the<br />
membrane and how it can be captured in a system-level<br />
macromodel. Additionally, systematic experimental and<br />
theoretical investigations on viscous damping in the rarefied<br />
gas regime have to be carried out in order to gain a better data<br />
base and, thus, a better understanding of the underlying physics<br />
in order to include this effect correctly in our models.<br />
[1] J. Iannacci, F. Giacomozzi, S. Colpo, B. Margesin and M. Bartek,<br />
“A General Purpose Reconfigurable MEMS-Based Attenuator for<br />
Radio Frequency and Microwave Applications”, in Conf. Proc. of<br />
EUROCON, 2009, pp: 1201-1209<br />
[2] G. Schrag, R. Khaliliyulin, M. Niessner, and G. Wachutka,<br />
“Hierarchical Modeling Approach for Full-System Design and<br />
control of Microelectromechanical Systems,” in Conf. Proc. of<br />
Eurosensors XXII, 2008, pp. 528-531.<br />
[3] L. Gabbay, J. Mehner, and S. Senturia, “Computer-aided<br />
generation of reducedorder dynamic macromodels - I:<br />
Geometrically linear motion,” J. Microelectromechanical Systems,<br />
vol. 9, 2000, pp. 262–269.<br />
[4] Schrag G., and Wachutka G., “Physically based modeling of<br />
squeeze film damping by mixed-level system simulation,” Sensors<br />
and Actuators A, vol. 97-98; 2002 , pp. 193–200.<br />
[5] R. Sattler, “Physikalisch basierte Mixed-Level Modellierung von<br />
gedämpften elektromechanischen Systemen“, Shaker Verlag,<br />
Aachen, 2007.<br />
[6] M. Niessner, et al., “Experimentally validated and automatically<br />
generated multi-energy domain coupled model of a RF-MEMS<br />
switch”, Proc. EuroSimE 2009, Delft, NL, April 27-29, 2009, pp.<br />
595-600.<br />
[7] M. Niessner, G. Schrag, G. Wachtuka and J. Iannacci, “Modeling<br />
and fast simulation of RF-MEMS switches within standard IC<br />
design framework,” Proc. SISPAD 2010, Bologna, Italy,<br />
September 6-8, 317-320 (2010).<br />
[8] L. del Tin, et al., Digest Tech. Papers of 14 th Int. Conf. on Solidstate<br />
Sensors, Actuators and Microsystems (Transducers’07),<br />
Lyon, June 10-14, pp.635-638, 2007..<br />
[9] M. Niessner, G. Schrag, J. Iannacci, G. Wachutka, “Mixed-level<br />
modeling of squeeze film damping in MEMS: Simulation and<br />
pressure-dependent experimental validation, accepted for<br />
publication at 16 th Int. Conf. on Solid-state Sensors, Actuators and<br />
Microsystems (Transducers’11), Beijing, China, June 5-9, 2011.<br />
[10] J. Iannacci, R. Gaddi and A. Gnudi,, “Experimental Validation of<br />
Mixed Electromechanical and Electromagnetic Modeling of RF-<br />
MEMS Devices Within a Standard IC Simulation Environment,”<br />
Journal of Microelectromechanical Systems 19 (3), 526-537<br />
(2010).<br />
[11] Coventor, Inc. [Coventorware Architect Version 2008.10<br />
Reference] (2008).<br />
13
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Investigation on the effect of geometrical dimensions<br />
on the conductive behaviour of a MEMS convective<br />
accelerometer<br />
A.A. Rekik 1,2 , B. Mezghani 2 , F. Azaïs 1 , N. Dumas 1 , M. Masmoudi 2 , F. Mailly 1 , P. Nouet 1<br />
1 LIRMM - CNRS/Univ. Montpellier 2 - 161 rue Ada, 34095 Montpellier, France<br />
2 ENIS - University of Sfax - Route Soukra, BP 1173 3038 Sfax, Tunisia<br />
Abstract— This paper presents an investigation on the effect<br />
of geometrical dimensions on the conductive behaviour of a<br />
CMOS MEMS convective accelerometer. Numerous FEM<br />
simulations are conducted to prove the validity of a previously<br />
developed model. This model was firstly developed to<br />
represent the effect of only one geometrical parameter; the<br />
etched cavity depth. We prove here that this model may<br />
represent also the effect of other geometrical parameters, the<br />
cavity half-width and the package height, on the conductive<br />
behaviour of the sensor.<br />
Keywords— MEMS, convective accelerometer, heat<br />
conduction, modelling.<br />
I. INTRODUCTION<br />
CMOS MEMS convective accelerometer is one example<br />
of a monolithically integrated sensor system. Such<br />
accelerometers present several advantages compared to the<br />
commonly used capacitive ones [1-2]. In addition to low<br />
cost batch fabrication, convective accelerometers are able to<br />
survive and sense very high shocks [3]. Mechanical rigidity<br />
and low mass are the main reasons that insure robustness to<br />
large accelerations. Unfortunately, and as it is generally the<br />
case for monolithic structures, strict requirements for<br />
process compatibility usually limit the performance level<br />
and potential applications [4-5]. To take advantage of the<br />
benefit of such a sensor, one use a system level design<br />
approach that requires an accurate and complete model of<br />
the sensor. This is considered as a vital and important step<br />
to help in predicting and/or increasing the overall system<br />
performance. One possible strategy is then to optimize the<br />
sensor dimensions to cope with the specifications and to<br />
adapt the global architecture accordingly.<br />
Existing models allow understanding the behaviour of a<br />
given sensor or studying geometry variations in a single<br />
dimension [6]. For multiple parameter variations, evaluation<br />
of geometrical and material properties in convective<br />
accelerometers is only possible through Finite Element<br />
Modelling (FEM) [7-8], which is very time consuming and<br />
meaningless. At the contrary, compact models allow a quick<br />
exploration of broad design spaces and a physical<br />
understanding of parameter influences. In a previous work<br />
[9], we have proposed a sensor model in which physicallybased<br />
expressions were developed to describe the<br />
conductive behaviour of a thermal accelerometer. This<br />
preliminary model was only including the effect of a single<br />
geometrical parameter (the depth of the bottom cavity, h 1 in<br />
Fig.1) since this parameter is the most difficult to control; it<br />
depends on the concentration of the etching solution and on<br />
the etching time. Initially developed to validate test<br />
methods, this model was then considered to validate system<br />
level architectures. Additional geometrical parameters<br />
appeared to be critical parameters for the design of the<br />
sensor. Therefore, the impact of these parameters has been<br />
investigated and a fully parameterized model is today<br />
introduced in this paper.<br />
In section 2, we briefly describe the accelerometer. In<br />
section 3, we present the previously published model of the<br />
sensor and we focus on the conduction part (heater source<br />
and common mode). Using FEM simulations, the effect of<br />
sensor dimensions on the conductive behaviour of the<br />
device is investigated in section 4. Finally, the proposed<br />
model validity is demonstrated over a large range of<br />
geometrical dimensions.<br />
II. ACCELEROMETER PRESENTATION<br />
The device under study is a convective accelerometer<br />
obtained by Front-Side Bulk Micromachining (FSBM) of a<br />
CMOS die fabricated in a 0.8 µm technology from Austria<br />
Microsystems® (Fig.1). Main lateral dimensions are the<br />
half-width of both the heater beam (r 1 ) and the cavity (r 2 ),<br />
and the distance between the heater and one detector (d).<br />
The three thin bridges are composed of the CMOS process<br />
back-end layers (oxide, polysilicon, aluminium, and<br />
nitride). In particular, polysilicon is used to implement<br />
resistors, for both heating (R H ) and temperature sensing<br />
(R D1 , R D2 ). The heater R H is powered by an electrical voltage<br />
(U H ) to create a temperature gradient in the bottom (i.e.<br />
etched silicon) and top (i.e. package) cavities: the<br />
temperature is then maximum at the heater location and<br />
minimum at the cavity boundaries.<br />
In absence of acceleration, the temperature detectors (R D1 ,<br />
R D2 ) are located on an identical isotherm for symmetry<br />
reasons. Under acceleration along the AA’-axis, the free<br />
convection deforms the cavity temperature distribution so<br />
14
that detectors may measure the differential temperature.<br />
Indeed, polysilicon resistivity exhibits high temperature<br />
dependence (Temperature Coefficient of Resistance,<br />
TCR=9×10 -4 /°C) and the thermal signal is easily converted<br />
into a voltage by means of a Wheatstone bridge. This<br />
voltage is then amplified by an instrumentation amplifier.<br />
For more details on sensor manufacturing and<br />
characterization, please refer to previous works from some<br />
of the authors [8-11].<br />
Heater<br />
Detectors<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
The model is composed of four main blocks (Fig.2). First,<br />
the heater temperature (T H ) is calculated from the external<br />
power supply (U H ). Then, the common mode temperature is<br />
calculated from the heater temperature. The heater<br />
temperature (T H ) and the common mode temperature (T CM )<br />
are governed by the heat conduction in the fluid. Regarding<br />
fluid convection, the differential temperature (ΔT D ), resulting<br />
from acceleration, is calculated with an empirical expression<br />
that involves fitting parameters. Finally, regarding<br />
transduction, both detector resistances (R Di ) are deduced<br />
from their TCR factor.<br />
In this paper, we will focus on the conduction part only<br />
(heater and common mode).<br />
Amplifier A1<br />
Sensing<br />
direction<br />
(AA)<br />
A. Heater source<br />
During normal operation, the heating bridge is powered by<br />
a DC voltage (U H ) in order to set an initial temperature<br />
distribution in the cavity. The average heater temperature<br />
(T H ) is directly linked to the electrical power (P H ) with a<br />
linear relationship:<br />
Figure 1. SEM picture of the prototype and geometrical parameters:<br />
r 1=20μm, r 2=350μm, d= 125μm, h 1=300μm, h 2=1000μm.<br />
III. ACCELEROMETER MODEL<br />
For defect simulations, a previously developed high level<br />
model [11] has been improved to include the effect of<br />
etching defects [9]. This model only takes into account the<br />
effect of the bottom cavity depth h 1 . No other sensor<br />
dimensions were included. In addition, this model was<br />
developed for fixed values of cavity half-width (r 2 =350µm)<br />
and package height (h 2 =10mm).<br />
T<br />
T<br />
Rth<br />
P<br />
T<br />
Rth<br />
2<br />
UH<br />
HAHHA<br />
⋅+=⋅+<br />
(1)<br />
RH<br />
where T A is the ambient temperature, R H is the electrical<br />
resistance of the beam and Rth H is the thermal resistance of<br />
the heater beam.<br />
The temperature dependence of the electrical resistance is<br />
taken into account in the power dissipation with:<br />
R<br />
R<br />
( 1 TCR T )<br />
Δ ⋅+=<br />
(2)<br />
0 HH<br />
where TCR is the temperature coefficient given by the<br />
foundry and R H0 is the nominal value of the heater resistance<br />
at a reference temperature T 0 . This electrical resistance is<br />
obviously independent of the cavity depth but only depends<br />
on the heating resistance size and material properties.<br />
In contrast, the thermal resistance of the beam depends not<br />
only on the beam dimension but also on its geometrical<br />
environment:<br />
Rth<br />
H<br />
=<br />
1<br />
⋅ Sh<br />
where S=54.7×10 3 µm² denotes the exchange surface<br />
between the beam and the fluid given by 2(2r 1 +e)L, with L<br />
the beam length and e the beam thickness.<br />
The heat transfer coefficient h H depends on the heater<br />
temperature, the ambient temperature, the geometry of the<br />
heater beam and the geometry of its environment.<br />
Assuming a cylindrical heater with a radius R 1 at a<br />
temperature T H surrounded by a cylindrical cavity with a<br />
radius R 2 at a temperature T A , the heat transfer coefficient h H<br />
is given by [9]:<br />
H<br />
(3)<br />
Figure 2. Block diagram of the sensor model.<br />
h<br />
H<br />
=<br />
1<br />
(4)<br />
⎛<br />
2 2<br />
3 3<br />
4 4<br />
λ<br />
⎞<br />
0 ⎜<br />
δ H1 −TT<br />
A δ H2 −TT<br />
A δ3<br />
H −TT<br />
A<br />
1+⋅<br />
+ ⋅ + ⋅ ⎟<br />
⎛ R ⎞<br />
⎝<br />
2 − AH 3TT<br />
− AH 4TT<br />
−<br />
AH<br />
⋅<br />
⎠<br />
⎜<br />
2<br />
l ⎟ nR<br />
⎝ R1<br />
⎠<br />
15
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
where λ 0 =-3.9333×10 -4 Wm -1 K -1 is the air conductivity<br />
extrapolated at 0K. The three parameters δ 1 =-0.2589 Wm -1 K -<br />
2 , δ 2 =1.2349×10 -4 Wm -1 K -3 and δ 3 =-3.8662×10 -8 Wm -1 K -4<br />
represent the coefficients of thermal conductivity variation<br />
for the air. We have shown in [9] that, for the complex<br />
geometry of our accelerometer, expression 4 is still valid if:<br />
• R 1 is a trade-off between the radius of a cylinder with<br />
a perimeter equivalent to the heater bridge and the radius of<br />
a cylinder that contains the heater beam.<br />
• R 2 is a function of the bottom cavity depth h 1 :<br />
h<br />
rR ⋅=<br />
(5)<br />
1<br />
22<br />
4<br />
4 ⎛ r<br />
4<br />
2 ⎞<br />
h1<br />
+ ⎜<br />
⎝ 2 ⎟ ⎠<br />
where the polynomial degree and the root order were<br />
chosen to obtain the best fit in two regions : a linear one<br />
obtained for low values of h 1 and a saturation region at r 2<br />
obtained for high values of h 1 . This model was developed for<br />
fixed values of r 2 and h 2 (350µm and 10mm, respectively).<br />
Also, it is found that the transition distance between the two<br />
regions seemed to be equal to r 2 /2. However, a parametric<br />
study on r 2 and h 2 should be conducted to confirm that it was<br />
not only a coincidence.<br />
B. Common mode<br />
Considering the same cylindrical geometry assumption,<br />
the relationship between the common mode temperature<br />
T CM , at a distance d from the center (R 1
2) Cavity width effect<br />
To verify the influence of the cavity half-width r 2 on the<br />
conductive behaviour of the sensor, a FEM simulation is<br />
performed for a new set of parameters: r 2 =700µm,<br />
h 1 =600µm, h 2 =10mm and T H =600K. The heat transfer<br />
coefficient h H and the common mode temperature T CM are<br />
extracted from this simulation and compared with the results<br />
obtained from expressions 4, 5, 6 and 7. The obtained FEM<br />
and model values are given in table II, which shows that the<br />
results are approximately equal.<br />
TABLE II. MODEL AND FEM RESULTS OF THE HEAT TRANSFER<br />
COEFFICIENT AND THE COMMON MODE TEMPERATURE<br />
r 2=700µm, h 1=600µm, h 2=10mm, T H=600K<br />
Conduction Model FEM<br />
h H (W.m -2 .K -1 ) 612 616<br />
T CM (K) 369 371<br />
Therefore, it seems that the model is valid when r 2 varies<br />
but it should be verified further with other sets of<br />
parameters.<br />
3) Model validity using FEM simulations<br />
(a)<br />
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May 2011, Aix-en-Provence, France<br />
<br />
heater temperature may be strongly impacted for a given<br />
biasing voltage.<br />
In Figure 4, we plot the heat transfer coefficient h H (Fig.<br />
4.a) and the common mode temperature T CM (Fig. 4.b)<br />
extracted from FEM simulations versus those calculated<br />
from the model. We clearly notice that the model and FEM<br />
results are in good agreement. In both figures, the points<br />
which diverge from the ideal curve correspond to very low<br />
values of the cavity depth h 1 0.2r 2 ), the relative error between the model and<br />
FEM results remains below 4%. To conclude, the validity<br />
domain of the model is then verified for 200µm
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Design and Simulation of an On-chip Oversampling<br />
Converter with a CMOS-MEMS Differential<br />
Capacitive Sensor<br />
Abstract- This paper presents the design and analysis of an<br />
integrated oversampling converter with MEMS capacitive<br />
sensor. The MEMS capacitive sensor is a comb-drive which<br />
provides change in capacitance when change in acceleration is<br />
detected. The analog-to-digital converter (ADC) is a first-order<br />
1-bit sigma-delta (Σ-Δ) converter. Σ-Δ ADCs are suitable for<br />
MEMS sensors since their output voltage are in mV and are low<br />
frequencies. Both the Σ-Δ ADC and MEMS capacitive sensor<br />
were designed in Silterra’s 0.13μm CMOS process. Simulation<br />
of the Σ-Δ ADC was conducted using Cadence TM Spectre, while<br />
the MEMS sensor was simulated by COMSOL Multiphysics ® .<br />
Results indicate that the Σ-Δ ADC can process small voltage<br />
output of the MEMS sensor and convert it into digital signals<br />
satisfactorily.<br />
Ma Li Ya, Anis Nurashikin Nordin, Sheroz Khan<br />
Department of Electrical and Computer Engineering<br />
International Islamic University Malaysia<br />
Jalan Gombak 53100, Kuala Lumpur, Malaysia<br />
I. INTRODUCTION<br />
With the development of the circuit and fabrication<br />
technology, great progress has been made in the effort to<br />
integrate MEMS devices with Integrated Circuit (IC)<br />
together on the same chip. Most of the MEMS sensors on the<br />
market can be divided into two groups; piezoresistive<br />
solutions and, the most used solution today, capacitive<br />
solutions [1]. The accelerometer sensor, for example, detects<br />
changes in acceleration and reflects it as capacitance.<br />
Low-cost and small-footprint MEMS accelerometers with<br />
high sensitivity, high resolution, and low power consumption<br />
are required for applications ranging from GPS-augmented<br />
inertial navigation systems to guidance and stabilization of<br />
satellites and spacecrafts [2]. Fig.1 illustrates accelerometer<br />
sensors in the range of different acceleration and bandwidth.<br />
Most acceleration sensors operate in low frequency with<br />
small acceleration values. This requires a very high<br />
resolution interface circuit to process such signals. The<br />
oversampling Σ-Δ modulator is extremely suitable for low<br />
frequency; weak input applications [3], as shown in Table I.<br />
This paper presents a first-order Σ-Δ converter with a<br />
CMOS-MEMS differential capacitive sensor designed on the<br />
same chip using Silterra 0.13μm CMOS process technology.<br />
Section II describes the MEMS capacitive sensor’s design.<br />
Section III illustrates the design of first-order Σ-Δ converter.<br />
Section IV displays the simulation results and frequency<br />
domain analysis.<br />
Fig.1. Accelerometer sensors with different range of acceleration and<br />
bandwidth [4].<br />
II. DESIGN OF CMOS-MEMS DIFFERENTIAL CAPACITIVE<br />
SENSOR<br />
Fig.2 shows a simplified schematic of a differential<br />
capacitive sensor which is under (a) zero acceleration and (b)<br />
non zero acceleration conditions. A differential capacitive<br />
sensor is used to minimize the systematic offset, improve<br />
power supply rejection and reduce drift of the sensor. This<br />
accelerometer consists of a moveable proof mass and two<br />
fixed parts. The proof mass is made by the Ultra Thick Metal<br />
(UTM) of copper in Silterra 0.13μm CMOS process, and it is<br />
suspended on the substrate using four tethers. The four<br />
TABLE I<br />
SIGNAL BANDWIDTH AND CONVERSION RESOLUTION TRADEOFF<br />
ADC<br />
Applications<br />
Conversion<br />
Resolution<br />
High<br />
Middle<br />
Low<br />
Sub<br />
lower<br />
Signal Bandwidth<br />
Low Middle High<br />
Sigma-Delta<br />
Successive<br />
Approximation<br />
Sub ranging /<br />
Pipelined<br />
Flash<br />
18
11-13 May 2011, Aix-en-Provence, France<br />
<br />
TABLE II<br />
ACCELEROMETER SPECIFICATIONS<br />
Model’s total dimensions 370μm×414μm×3.5μm<br />
Proof mass<br />
1.2μg<br />
Number of finger pairs 50<br />
The range of Δd<br />
-1.5μm~1.5μm<br />
The range of acceleration -3g~3g (1g=9.8m/s 2 )<br />
The range of capacitive change 132.81fF~929.67fF<br />
C + = n⋅[ε 0 ε r A/(d 0 -Δd)] (4)<br />
C - = n⋅[ε 0 ε r A/(d 0 +Δd)] (5)<br />
where, n is the number of the finger pairs;<br />
ε 0 = 8.854×10 −12 F m –1 is the electric constant;<br />
ε r = 1 is the dielectric constant of the material<br />
between the plates;<br />
A is the area of the overlapping of the two plates;<br />
d 0 = 2μm is the distance of two plates when there’s<br />
no acceleration existing.<br />
Fig.2. The simplified schematic of a differential capacitive model, (a)<br />
acceleration is zero (a=0); (b) acceleration is non zero (a>0); (c) the<br />
structures of four tethers.<br />
tethers work like mechanical springs. When the substrate<br />
undergoes any external acceleration (a) in its sense direction<br />
(for this model is horizontal direction), the proof-mass exerts<br />
a force (F) on the suspension, according to Newton second<br />
law. At the same time, for frequencies below the mechanical<br />
resonance of the spring-mass system, this force causes the<br />
suspension to deflect a distance (Δd), according to Hooke’s<br />
law. The relationship of these values can be showed using<br />
following (1) and (2).<br />
F = ma (1)<br />
where, m is the total mass of the proof;<br />
a is the external acceleration;<br />
F is the force of the proof generated.<br />
Δd = F/k = ma/k = a(1/ω 2 n ) (2)<br />
where, Δd is the distance for proof-mass moving;<br />
k is the overall spring constant;<br />
ω n is the natural frequency of the sensor in the<br />
direction of applied acceleration.<br />
The overall spring constant is,<br />
k = x⋅(12EI/L 3 ) = x⋅(Ewt 3 /L 3 ) (3)<br />
where, x = 4 is the number of the tethers;<br />
E is the Young’s modulus of UTM;<br />
w = 3.5μm is the width of the tether;<br />
t = 2μm is the thickness of the tether;<br />
L = 170μm is the length of the tether.<br />
The equivalent circuit of the differential capacitive sensor<br />
for the MEMS model is shown in Fig.3. Table II describes<br />
the specifications of this accelerometer. The values of these<br />
two capacitors can be derived from the distance changing of<br />
Δd, as (4) and (5) shown.<br />
III.<br />
DESIGN OF A SIGMA-DELTA ANALOG-TO-DIGITAL<br />
CONVERTER (Σ-Δ ADC)<br />
A. Overall Design<br />
Fig.4 demonstrates the whole system block diagram. This<br />
CMOS monolithic chip mainly contains two parts, sensor<br />
and interface circuit. The sensor circuit converts the physical<br />
parameter to an analog signal, which is sent to the ADC. The<br />
CMOS monolithic chip provides a 1-bit digital output which<br />
is fit for signal processing or transmission. In order to make<br />
design simple, low cost, and high resolution, first-order<br />
sigma-delta analog-to-digital converter was chosen as the<br />
interface circuit.<br />
The circuit level design is presenting in Fig.5. The MEMS<br />
sensor part is represented as two sensor capacitances C + and<br />
C - which are both in the range of pF or fF. The sensor circuit<br />
converts the sensing capacitances to analog voltage, as<br />
described in (6) below.<br />
V a = V 1 [(C + -C - )/(C + +C - )] = V 1 (Δd/d 0 ) (6)<br />
where, V 1 is the amplitude of the sinusoid.<br />
Fig.4. Integrated micro system block diagram.<br />
Fig.3. The equivalent circuit of differential capacitive sensor.<br />
Fig.5. The whole chip circuit design.<br />
19
B. First-order Σ-Δ ADC Design<br />
The term of Σ-∆ ADC has become almost synonymous<br />
with noise shaping ADC. Oversampling reduces the<br />
quantization noise power in the signal bandwidth by<br />
spreading the quantization noise power over a larger<br />
frequency range [5]. Noise shaping attenuates this noise in<br />
the signal bandwidth and amplifies it outside of the signal<br />
bandwidth. A low-pass filter is used to attenuate the<br />
out-of-band quantization noise. A down sampling circuit is<br />
added to obtain the Nyquist rate output.<br />
As shown in Fig.5, the first-order Σ-∆ ADC consists of a<br />
discrete-time integrator, a 1-bit ADC (comparator), a D<br />
flip-flop, and a 1-bit digital-to-analog converter (DAC) in the<br />
feedback path. Here, the internal ADC and DAC are both low<br />
resolution. Insertion of a buffer in series with a comparator is<br />
done in order to make the D flip-flop work properly.<br />
The analog signal processing circuit comprises of<br />
switch-capacitor circuits [6]. In comparison with the<br />
continuous time circuits consisting of resistors, capacitors<br />
and op amp [7], this technique produces a more accurate<br />
frequency response, good linearity and dynamic range. The<br />
most important component which is displayed in Fig.6 is the<br />
two-stage op amp. Table III describes the op amp design<br />
specifications while Table IV summarizes the op amp<br />
simulation results. The design of comparator is almost same<br />
as the op amp, except there’s no compensation capacitor<br />
between differential stage and inverter stage.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
TABLE IV<br />
THE OP AMP DESIGN RESULTS<br />
Transistor M1 M2 M3 M4 M5 M6 M7 M8<br />
(W/L) 36 36 6 6 39.38 39.6 107.8 78.77<br />
Component I dc C c<br />
Value 80μA 8pF<br />
IV.<br />
SIMULATION AND RESULTS ANALYSIS<br />
A. MEMS Capacitive Sensor Simulation<br />
The CMOS-MEMS differential capacitive sensor was<br />
simulated using COMSOL Multiphysics ® , and the 3D model<br />
is presented in Fig.7.<br />
Equations (4) and (5) provide the relationship between the<br />
distance changes (Δd) and the values of differential<br />
capacitors; and the simulated results are shown in Fig.8. It<br />
can be observed that the distance changes of within 1.5μm<br />
causes the capacitance to change in hundredfold of fF. The<br />
changing output voltage however, is linear to the sensing<br />
displacement of as shown in Fig.9.<br />
Fig.7. The CMOS-MEMS differential capacitive model in 3D.<br />
Fig.6. The schematic of two-stage op amp.<br />
Specification<br />
TABLE III<br />
THE OP AMP DESIGN SPECIFICATIONS<br />
Power Band Phase<br />
Supply Width Margin<br />
Open<br />
Loop Gain<br />
Design Value ±1.2V 10kHz ≥75° ≥100<br />
Input<br />
Slew Settling<br />
Specification Signal<br />
Rate Time<br />
Range<br />
Minimum<br />
Length of<br />
Transistor<br />
Design Value -0.7V~0.7V 4.45μs ≤1kΩ 130nm<br />
Fig.8. Calculated and simulated results for the differential capacitive sensor.<br />
20
Fig.9. Capacitive sensor output, Va as the function of the sensing distance,<br />
Δd.<br />
B. First-Order Σ-∆ Converter Simulation<br />
The overall design of the Σ-∆ ADC shown was simulated<br />
using Cadence TM Spectre. The input test analog signal is a<br />
100mV, 500 Hz sinusoid with the oversampling frequency is<br />
160 kHz. The oversampling ratio (OSR) is 160 and the<br />
output resolution is around 10 and the dynamic range is over<br />
60dB. Fig. 10 illustrates the waveforms of each stage output.<br />
C. Results Analysis<br />
Normally there are two ways to test the output of ADC: i)<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
using an ideal DAC to convert the digital signal back to an<br />
analog signal and compare it with the original one, however<br />
this requires to design another very high resolution DAC; ii)<br />
the most popular method is using fast Fourier-transform<br />
(FFT) to analysis the output signals in frequency domain.<br />
FFT is commonly used to estimate the power spectral<br />
density (PSD) of Σ-∆ converter. Data obtained from Cadence<br />
circuit simulation is transferred to MATLAB. This is done<br />
for signal processing analysis of the final output of the signal.<br />
Fig. 11 shows the low frequency portion (0 to 10 kHz) of an<br />
FFT based PSD estimate of the output with a sinusoidal input<br />
frequency of 500Hz and f s of 160 kHz. The spectrum of Fig.<br />
11 (a) consists of one large spike representing the input<br />
signal sine wave, plus many smaller spikes distributed out of<br />
the base-band frequency along the frequency axis,<br />
representing white noise. The PSD of the output digital<br />
signal is shown in Fig. 11 (b). It demonstrates that this<br />
designed first-order sigma-delta converter really does shape<br />
quantization noise. From this figure it can be seen that the<br />
frequency of the largest spike signal is still 500Hz, which<br />
represents the input signal. The following smaller signals are<br />
the quantization errors; and their amplitudes are obviously<br />
lower than the main signal. The Σ-Δ converter performs<br />
noise-shaping and forces all the quantization errors to be<br />
outside the frequency of the base band signal.<br />
l<br />
Fig.10. Each stage output of the first-order Σ-∆ ADC.<br />
21
(a)<br />
PSD of the input signal.<br />
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May 2011, Aix-en-Provence, France<br />
Anis Nurashikin Nordin received the B. Eng. Degree in Computer and<br />
Information Engineering from the International Islamic University Malaysia<br />
(IIUM), Kuala Lumpur, Malaysia in 1999, and the M.S degree in Computer<br />
Engineering from the George Washington University (GWU), Washington<br />
DC, in 2002, and the D. Sc. Degree in Electrical and Computer Engineering<br />
at GWU. Currently, she is a lecturer at International Islamic University<br />
Malaysia (IIUM).Her main research interests are VLSI.<br />
Sheroz Khan is currently working as a faculty member within the<br />
Department of Electrical and Computer Engineering, here at the<br />
International Islamic University Malaysia since January 2002. At the<br />
moment he is leading a research group, called Wireless Communication and<br />
Signal Processing where a group of over twenty research students working<br />
for their M Sc/Ph Ds under the supervision of qualified colleagues as faculty<br />
members. Sheroz Khan's areas of interest include sensors and transducer<br />
interfacing, electronic instrumentation, embedded systems. He has worked<br />
on UWB signals generation, shaping and modulation. Also, his interests<br />
include those for biomedical medical electronic application, including areas<br />
such energy scavenging and devices such as smart and intelligent<br />
transducers designed and developed for contactless data acquisition from<br />
inaccessible points.<br />
ACKNOWLEDGMENT<br />
This work is supported by International Islamic University<br />
Malaysia’s Endowment Fund (EDW B 0905-304).<br />
(b)<br />
PSD of the output signal.<br />
Fig.11. The power spectral density of input and output signals.<br />
V. CONCLUSION<br />
This paper presents a design, simulation and analysis of an<br />
Σ-∆ interface circuit for a CMOS-MEMS differential<br />
capacitive sensor. The results demonstrate that the usage of<br />
an Σ-∆ modulator allows very weak analog signals to be<br />
converted to an extremely high resolution digital output. The<br />
usage of Silterra 0.13μm CMOS process allows the on-chip<br />
area to be kept to a minimum. This Σ-∆ modulator has an<br />
input signal frequency of 500 Hz, oversampling frequency of<br />
160 kHz, oversampling ratio (OSR) of 160; and<br />
signal-to-noise ratio of 62.98 dB.<br />
REFERENCES<br />
[1] Karianne Qysted and Dag T. Wisland, University of Olso, Dep. of<br />
Informations, Norway, “Piezoresistive CMOS-MEMS Pressure<br />
Sensor with Ring Oscillator Readout Including Δ-Σ<br />
Analog-to-Digital Converter On-chip”, IEEE Custom Integrated<br />
Circuits Conference, 2005.<br />
[2] Babak Vakili Amini, Student Member, IEEE, and Farrokh Ayazi,<br />
Member, IEEE, “A 2.5-V 14-bit ΣΔ CMOS SOI Capacitive<br />
Accelerometer”, IEEE Journal of Solid-state Circuits, Vol. 39, No.<br />
12, December 2004.<br />
[3] A. N. Nordin and M. Zaghloul, “CMOS design and implementation<br />
of sigma-delta analog-to-digital data converter suitable for MEMS<br />
devices”, 2003.<br />
[4] Bernhard E. Boser, “Surface Micromachining An IC-Compatible<br />
Sensor Technology”, Berkeley Sensor & Actuator Center Dept. of<br />
Electrical Engineering and Computer Sciences University of<br />
California, Berkeley.<br />
[5] Shlomo Engelberg, “Instrumentationnotes: Sigma-Delta<br />
Converters: Theory and Simulations,” IEEE Instrumentation &<br />
Measurement Magazine, pp. 49-53, December 2007.<br />
[6] E. Dallago, P. Malcovati, D. Miatton, T. Ungaretti, and G. Venchi,<br />
“Analysis of sigma-delta converter for MEMS sensors using power<br />
supply voltage as reference”, Circuits, Devices and Systems, IEE<br />
Proceedings -, vol. 153, pp. 473-479, 2006.<br />
[7] R. Jacob Baker, “CMOS Mixed-Signal Circuit Design”, second<br />
edition, IEEE Press, 2009, New York.<br />
VI. BRIEF BIOGRAPHY OF THE AUTHOR<br />
Ma Li Ya received the B. Eng Electronic and Information Engineering<br />
from Changchun University, Jilin Province, China in 2007 and currently<br />
working toward Master Degree in Electronic Engineering. Her main<br />
research interests are Mix-signal Integrated Circuit Design and MEMS. She<br />
currently involves in the design and simulation of analog-to-digital<br />
converter for low-frequency MEMS sensor applications.<br />
22
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May, 2011, Aix-en-Provence, France<br />
<br />
Fabrication of High Aspect Ratio Nanoporous Array<br />
on Silicon<br />
Jing-Yu Ho 1 1, 2*<br />
and Gou-Jen Wang 1 Department of Mechanical Engineering<br />
2 Graduate Institute of Biomedical Engineering<br />
National Chung-Hsing University, Taichung 40227, Taiwan<br />
Tel:+886-4-22840725 x 320<br />
Email: gjwang@dragon.nchu.edu.tw<br />
Abstract- In this study, a simple method for the fabrication of high<br />
aspect ratio silicon nanoporous arrays is developed. A N-type<br />
silicon wafer is used as the material; a micro-scale pattern of the<br />
desired porous array is transferred to the front surface of the<br />
silicon wafer by photolithography; the wafer is placed in a<br />
home-made fixture to efficiently expel the etching generated air<br />
and promptly hold the back-side illumination light; a halogen<br />
lamp is used as the light source for backside illumination to<br />
enhance the electron-hole pairs generation; anodization is then<br />
processed using a new etchant which consists of the hydrofluoric<br />
acid and the EtOH and EMSO mixed surfactant to effectively<br />
polish the pore surface and sharp the tips of the etched pores. A<br />
nanochannel array with nano-tip being 61.4 nm is obtained.<br />
I. INTRODUCTION<br />
Porous silicon has attracted increasing interest owing to the<br />
advantages such as easily being fabricated in large area, pores<br />
being able to be arranged in order and patterned, and pores’<br />
shape being able to be adjusted by the process parameters. In<br />
addition, the semiconducting characteristic enables a porous<br />
silicon device to reveal different physical and/or chemical<br />
properties due to variation of porosity. Porous silicon can be<br />
found variety applications in light emitting diode (LED) [1],<br />
photodetector [2], solar cell [3], photonic crystal, photo sensor<br />
[4], biosensor [5], and field emission display [6].<br />
It has been nearly 50 years since the invention of porous<br />
silicon. However, the formation mechanism of porous silicon is<br />
still being developed. Parkhutik et al. [7] proposed that the<br />
relatively thinner oxide layer between the pores and the<br />
substrate enhances the electric field at each pore end; hence the<br />
silicon atom can easily acquire electric hole then dissolves. A<br />
high aspect ratio structure can thus be obtained. The basic<br />
principle is similar to that of the anodic aluminum oxide.<br />
Unagami [8] proposed that the porous silicon layer (PSL) is<br />
constructed by the local dissolution of silicon which happens<br />
only at the base of the pores. The divalent and the tetravalent<br />
reactions of silicon with hydrofluoric acid (HF) are the main<br />
driving forces for the dissolution of silicon in the pores.<br />
Theunissen et al. [9] reported that at the tip of a pore in N-type<br />
material, silicon atoms are likely to reach the breakdown<br />
voltage then dissolve due to the concentration of electric field.<br />
Beale [10] suggested that the depletion layer influences the<br />
distribution of the electric field in the silicon substrate, hence<br />
affects the porous structure. Smith et al. [11] proposed that the<br />
diffusion of the electric holes toward the interface between the<br />
silicon material and the etching solution determines the etching<br />
process. Lehmann et al. [12] pointed out that the band gap of<br />
the whole system is increased after the growth of the porous<br />
silicon such that the charge concentration in the porous silicon<br />
is reduced. A high aspect ratio pore array thus can be produced.<br />
Several factors affect the structure of a silicon porous array.<br />
Rönnebeck et al. [13] investigated the etching behavior in<br />
N-type and P-type materials. It was pointed out that the degree<br />
of doping in silicon substrate influences the etching results<br />
since it determines the thickness of the depletion layer.<br />
Carstensen et al. [14] compared the effects of HF concentration,<br />
different surfactants, and illumination intensity. Tsuboi et al.<br />
[15] studied the influences of the applied field induced<br />
polarization of the surfactant on the etching results. Among<br />
those reported studies, the size of the pores ranges from<br />
submicron to several micron. For delicate applications such as<br />
the detection of DNA sequences, nano-scale pores are desired.<br />
In this study, a simple process for the fabrication of<br />
nano-size silicon porous array is proposed. A N-type silicon<br />
wafer is used as the material; a micro-scale pattern of the<br />
desired porous array is transferred to the front surface of the<br />
silicon wafer by photolithography; the wafer is placed in a<br />
home-made fixture; the anodization is then processed using HF<br />
as the etching solution mixed with two surfactants to smoothen<br />
the etched surface under backside illumination.<br />
2.1 Materials<br />
II.<br />
MATERIALS AND METHODS<br />
(1) N-type silicon wafer<br />
As mentioned, diffusion of the electric holes plays an<br />
important role in the etching process of silicon [11]. Since the<br />
major carriers in P-type material are electric holes, the<br />
dissolution rate of silicon in P-type material is higher than that<br />
in N-type material. However, the etching process in P-type<br />
material is difficult to be well controlled. Lateral etching is the<br />
general problem. In this study, N-type material with thickness<br />
and resistivity being 525 ± 25μm and 1~100Ω-cm 2 respectively<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
23
is hence used. To enhance the efficiency of electron-hole<br />
separation, the back-side illumination is implemented.<br />
(2) Etching solution<br />
Hydrofluoric acid is the commonly used etching solution<br />
for silicon etching. However, extra high surface tension of HF<br />
prevents the pore surface from being uniformly wetted during<br />
etching downward. Resultantly, lateral etching can always be<br />
observed. The surface tension effect can be softened by either<br />
adding surfactants to the solution [16] or increasing the<br />
conductivity of the solution [17]. Ethyl alcohol (EtOH),<br />
dimethyl sulfoxide (DMSO), N-dimethylformamide (DMF),<br />
and tetrabutylammonium perchlorate (TBAP) are the<br />
commonly used surfactants. EtOH can soft the surface tension<br />
of the etching solution so the pore surface can be better wetted<br />
to prevent the lateral etching effect. Both DMSO and DMF are<br />
polar aprotic solvent which can enhance the downward<br />
diffusion of fluorine ions such that a smoother etching can be<br />
obtained. TBAP can increase the conductivity of the etching<br />
solution.<br />
In this study, HF is employed as the etching solution.<br />
EtOH is used to soft the surface tension of the etching solution.<br />
DMSO is added to increase the conductivity of the etching<br />
solution. DMSO is preferred because it is less toxicity than<br />
DMF.<br />
(3) Fixture<br />
Figure 1 illustrates the cross-sectional view of the<br />
home-made fixture. It consists of a top fixture made of Teflon<br />
to hold the etching solution; an o-ring to stable the silicon wafer;<br />
a Cu electrode; a bottom fixture to fix the silicon wafer and<br />
enable the back illumination through the =1 cm hole at its<br />
center. This vertical fixture enables the gaseous matter<br />
generated during etching process to escape from the pore inside<br />
to the wafer surface.<br />
Figure 1. Home-made fixture for the etching process<br />
(4) Mask<br />
The mask used in this study is shown in Figure 2. It is a <br />
=1 cm circle containing an array of holes with hole diameter<br />
and line pitch being 6 m and 10 m, respectively.<br />
11-13 <br />
May, 2011, Aix-en-Provence, France<br />
<br />
Figure 2. Mask used for patterning the pore array<br />
(5) Light source for backside illumination<br />
A 150 w halogen lamp is used as the light source for<br />
backside illumination. The light is guided to the bottom fixture<br />
by a 150 cm long optical fiber.<br />
2.2 Methods<br />
Figure 3 schematically illustrates the procedures of the<br />
proposed etching process for nanopore in silicon. It includes<br />
thin film deposition, photolithography, inductive coupled<br />
plasma (ICP) dry-etching, and end-point detecting<br />
anodic-etching. The process details are described below.<br />
Step (A): Deposit Si 3 N 4 as the hard mask<br />
The silicon wafer used is a 380 μm thick N-type<br />
wafer provided by the Wafer Works Corp. The Si 3 N 4 layer on<br />
both sides of the silicon wafer are deposited using the low<br />
pressure chemical vapor deposition (LPCVD) process such that<br />
the residual stress is reduced and a thicker film (10,000 Å) can<br />
be made. The process parameters are: temperature = 850 C,<br />
pressure = 180 mtorr, and reaction gases are NH 3 and SiH 2 Cl 2 .<br />
Step (B) & (C): Pattern the topside photoresist<br />
A etch window on the topside of the Si 3 N 4 deposited wafer<br />
is patterned by conventional photolithography process. Detail<br />
processes are listed below.<br />
i) Use the working mask as shown in Figure 2.<br />
ii) Spin-coat a 7 m thick positive photoresist (AZ-1518).<br />
Parameters for the spinning coating are: spinning speed of<br />
the 1 st stage= 500 rpm, spinning time for the first stage= 5<br />
sec, spinning speed of the 2 nd stage= 1500 rpm, spinning<br />
time for the 2 nd stage= 20 sec (Figure 3B).<br />
iii) Soft bake with temperature being set at 100 C for 2 min.<br />
vi) Expose (OAI 500 aligner) for 20 sec and develop (AZ-326<br />
MIF) for 15 sec.<br />
v) Hard bake at 120 C for 10 min.<br />
Step (D): Pattern the topside Si 3 N 4 film<br />
The ICP-RIE (Cirie-100) dry etching is adopted to<br />
transfer the pattern into the Si 3 N 4 film. The process parameters<br />
of the ICP-RIE are: reaction gas is CF 4 with flow rate being 45<br />
sccm, working pressure=5 mtorr, RF power=500 W, processing<br />
time=400 sec.<br />
Step (E): Remove the photoresist<br />
Step (F)-(I): Repeat processes (B)-(E) on the back surface with<br />
a 7×7 mm 2 working mask as the back etching window.<br />
Step (J): Process anisotropic wet etching on the backside silicon<br />
After removing the photoresist, the uncovered silicon<br />
surface is wet-etched by a 30% (w/w) KOH at 60C for 4 hours.<br />
Step (K): Deposit a 100 nm thick gold thin film as the end point<br />
detector for anodic etching.<br />
Step (L): Conduct anodic etching<br />
24
The anode and cathode are connected to the Cu electrode<br />
below the N-type silicon wafer and the Pt electrode in the<br />
etching solution to conduct anodic etching (Figure 4). When the<br />
pore array that is etched down from the tips of the inverted<br />
pyramids reaches the Au electrode, the current in the power line<br />
will increase rapidly (Figure 5). The anodic etching process can<br />
be terminated immediately.<br />
(A)<br />
(B)<br />
(D) (E) (F)<br />
(G) (H) (I)<br />
(J) (K) (L)<br />
Si Si 3 N 4 Photoresist Au<br />
Figure 3. Schematic illustration of the silicon pore array etching process<br />
Pt electrode<br />
(C)<br />
11-13 <br />
May, 2011, Aix-en-Provence, France<br />
<br />
III. RESULTS AND DISCUSSIONS<br />
3.1 Effect of HF concentration<br />
Several factors such as the applied voltage, concentration<br />
of HF, conductivity of etching solution, type of surfactant,<br />
intensity of illumination, and pH value affect the structure of a<br />
silicon porous array. Among them, the applied voltage and the<br />
concentration of HF are the major considerations. Therefore,<br />
only these two factors are selected as the process parameters for<br />
a more efficient investigation in this study. During the<br />
experiments, EtOH and DMSO are added to HF solution with<br />
concentration of 1M, 2M, and 4M respectively as the etching<br />
solutions. Various voltages were applied to the etching<br />
solutions for the anodic etching.<br />
(1) Etching results of the 1M HF<br />
Figure 6 shows the SEM images of the 1M HF etching results.<br />
In which, the insets are either the top view or the image of the<br />
45° incline. Table 1 tabulates the process parameters and the<br />
related results. For all experiments, the etching duration is 5 hr.<br />
The results in Table 1 indicate that the etching rates are similar<br />
except that of the 1.5 V. The SEM image in Figure 6(D) reveals<br />
that the process under 1.5 V potential can be categorized to<br />
electropolishing rather than anodic etching. In this condition, a<br />
silicon atom antecedently reacts with water molecule to become<br />
silicon dioxide. Silicon dioxide molecules are then dissolved by<br />
HF molecules. The chemical formulas for the reactions are<br />
shown in Equations (1)-(3). Electropolishing results in<br />
pore-widening on the pre-etched inverted pyramids; therefore,<br />
the originally designed 6 m pores are widened to 12 m pores<br />
with less depth.<br />
Etching solution<br />
O-ring<br />
Si + 4OH - +λh + → Si(OH) 4 + (4-λ)e - (1)<br />
Si(OH) 4 → SiO 2 +2H 2 O (2)<br />
SiO 2 + 6HF → H 2 SiF 6 + 2H 2 O (3)<br />
N-type Si<br />
Cu electrode<br />
Optical fiber<br />
Figure 4. Experimental apparatus<br />
Current (A)<br />
Figure 6. SEM images of the 1M HF etching results, (A)0.8V; (B)1V; (C)1.2V;<br />
(D)1.5V<br />
Figure 5. The current increase rapidly as the pore array reaches the Au electrode<br />
25
11-13 <br />
May, 2011, Aix-en-Provence, France<br />
Table 1. Etching results of the 1 M HF<br />
<br />
Voltage Channel depth Etching rate<br />
(V) (μm) (μm/hr)<br />
0.8 18.94 3.788<br />
1 20 4<br />
1.2 19.31 3.862<br />
1.5 14.06 2.812<br />
(2) Etching results of the 2M HF<br />
Figure 7 and Table 2 illustrate the etching results of the<br />
2M HF. As shown in Table 2, the applied voltage of 1V rather<br />
than the 1.5 V produced the fastest etching. It is presumed that<br />
the 1.5 V potential might reduce the effect of the electric field<br />
concentration, resulting in the increasing of the electric hole on<br />
the pore wall. The increasing etching reactions on the pore wall<br />
produced rugged wall surface. Since the etching reactions did<br />
not rivet on the pore tip, the etching rate is thus reduced (Figure<br />
7D).<br />
Figure 8. SEM images of the 4M HF etching results, (A)0.8V; (B)1V; (C)1.2V;<br />
(D)1.5V<br />
Table 3. Etching results of the 4 M HF<br />
Voltage<br />
(V)<br />
Channel depth<br />
(μm)<br />
Etching rate<br />
(μm/hr)<br />
0.8 94.62 18.924<br />
1 56.9 11.38<br />
1.2 22.765 4.553<br />
1.5 36.216 7.2432<br />
Figure 7. SEM images of the 2M HF etching results, (A)0.8V; (B)1V; (C)1.2V;<br />
(D)1.5V<br />
Table 2. Etching results of the 2 M HF<br />
Voltage<br />
(V)<br />
Channel depth<br />
(μm)<br />
Etching rate<br />
(μm/hr)<br />
0.8 29.6 5.92<br />
1 58.984 11.7968<br />
1.2 50.52 10.104<br />
1.5 49.743 9.9486<br />
(3) Etching results of the 4M HF<br />
The 4M HF etching results are shown in Figure 8 and<br />
Table 3. Figure 8 indicates that only the applied voltage of 0.8<br />
V can conduct acceptable etching. Since the concentration of<br />
HF is 4M, a larger applied voltage will draw forth far enough<br />
electric holes on the wafer surface to react with the fluoric ions.<br />
The etching process thus starts randomly from the wafer<br />
surface rather than form the tip of the inverted pyramids (Figure<br />
8B-8D).<br />
3.2 Effect of the protective Si 3 N 4 layer<br />
Electric field concentration is the basic principle of anodic<br />
etching. In the study, the spots of electric field concentration<br />
consist of the tips of the inverted pyramids and the four corners<br />
of etch pre-etched pore. The protective Si 3 N 4 layer further<br />
enhances this phenomenon. It is also interesting to investigate<br />
the influence of the protective Si 3 N 4 layer. The parameters for<br />
Figure 7(B) (HF=2M, applied voltage=1V) which had better<br />
etching performance are used for the etching without a Si 3 N 4<br />
layer.<br />
Figure 9 compares the results of the with Si 3 N 4 layer and the<br />
without Si 3 N 4 layer etchings. The top row and the bottom row<br />
illustrate the results of the without Si 3 N 4 layer and with Si 3 N 4<br />
layer etching, respectively. The etching time of the without<br />
Si 3 N 4 layer etching is 3 hr. The etching rate, which is estimated<br />
to be 11.61 m/hr, is close to that of the with Si 3 N 4 layer etching.<br />
For the without Si 3 N 4 layer etching, randomly etched cavities<br />
due to the reactions between fluoric acids and electron holes on<br />
the silicon surface can be observed. It can also be found that the<br />
wet etched squares are widened. For the Si 3 N 4 layer protective<br />
etching, symmetrical cannelures stretching from the four<br />
corners of etch wet etched square. It is presumed that the<br />
electric field concentration at the corners of a wet etched square<br />
at the initial stage of etching leads to the directional etchings of<br />
the cannelures. However, vertical microchannels having the<br />
size close to the wet etched square were fabricated. It reveals<br />
that a Si 3 N 4 layer is desired for a successful etching of porous<br />
array in silicon.<br />
26
11-13 May , 2011 , Aix-en-Provence, France<br />
the bottom part of the sample contains higher concentration of<br />
electric hole than the upper part. Hence, side etching is observed at the<br />
bottom part of a nanochannel. To reduce the side etching effect and<br />
ensure the dissolution reactions only occurring at the tip of the<br />
channel, the power of the halogen lamp needed to be weakened<br />
gradually along the etching process. In our later study, the<br />
power of the halogen lamp is gradually reduced to 10% of its<br />
maximum value. Figure 11 shows a SEM image of an etched<br />
channel under a light intensity attenuating process. A 61.4 nm<br />
nano-tip can be obtained.<br />
Figure 9. Effect of the protective Si 3 N 4 layer, top row: without Si 3 N 4 layer<br />
etching; bottom row: with Si 3 N 4 layer etching.<br />
3.3 Effect of the backside illumination<br />
Silicon atoms, electric holes, and fluorine ions are the<br />
three elements which involve in the etching process of silicon.<br />
Since the majority charge in N-type material is electron, back<br />
illumination is required to enhance the electron-hole pairs<br />
generation in the material. The generated electric holes will<br />
participate in the etching process, while the electrons will be<br />
conveyed from the anode through the power line to the cathode.<br />
A larger current in the power line indicates a larger amount of<br />
electric hole generated by the back illumination. Figure 10<br />
shows the dynamic polarization curves for the etching<br />
processes under back illuminations of various intensities. In<br />
which, the process of curve (C) is conducted under a stronger<br />
back illumination than that of the process of curve (B), while<br />
curve (A) denotes the process of no back illumination. The<br />
results reveal that a higher intensity of back illumination can<br />
generate a larger amount of electric hole to enhance the etching<br />
process.<br />
Figure 11. SEM image of an etched channel under a light intensity attenuating<br />
process<br />
IV. CONCLUSION<br />
In this study, a simple method for the fabrication of high<br />
aspect ratio silicon nanoporous arrays is developed. At the<br />
beginning, the photolithographic process is implemented to<br />
pattern micro-porous arrays on an N-type silicon wafer. The<br />
pre-etching by KOH is than conducted to produce an inverted<br />
pyramid array on the wafer, followed by an anodic etching<br />
process to further etch the pores down from the tip of each<br />
pyramid. The success of the proposed method can be attributed<br />
to two main reasons. (1) The home-made fixture to efficiently<br />
expel the etching generated air and promptly hold the back-side<br />
illumination light; (2) The using of a new etchant which<br />
consists of the hydrofluoric acid and the EtOH and EMSO<br />
mixed surfactant to effectively polish the pore surface and<br />
sharp the tips of the etched pores.<br />
ACKNOWLEDGEMENTS<br />
The authors would like to address their thanks to the National<br />
Science Council of Taiwan for their financial support of this<br />
work under grant NSC 97-2628-E-005-001-MY2.<br />
Figure 10. Dynamic polarization curves for the etching processes under back<br />
illuminations of various intensities<br />
Since the backside illumination is employed to produce<br />
electron-hole pairs in the N-type material, it can be presumed that<br />
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Dr. Gou-Jen Wang received the B.S. degree on 1981<br />
from National Taiwan University and the M.S. and<br />
Ph.D. degrees on 1986 and 1991 from the University<br />
of California, Los Angeles, all in Mechanical<br />
Engineering. Following graduation, he joined the<br />
Dowty Aerospace Los Angeles as a system engineer<br />
from 1991 to 1992. Dr. Wang joined the Mechanical<br />
Engineering Department at the National Chung-Hsing<br />
University, Taiwan on 1992 as an Associate Professor<br />
and has become a Professor on 1999. From<br />
2003-2006, he served as the Division Director of<br />
Curriculum of the Center of Nanoscience and Nanotechnology. Since 2007, he<br />
has been the joint Professor and Chairman of the Graduate Institute of<br />
Biomedical Engineering, National Chung-Hsing University, Taiwan. On 2008,<br />
he served as the Conference Chair of the Microfabrication, Integration and<br />
Packaging Conference (April/2008, Nice, France). From 2009, he is a<br />
Committee member of the Micro- and Nanosystem Division of the American<br />
Society of Mechanical Engineers. His research interests include MEMS,<br />
biomedical micro/nano devices, nano fabrication, and dye-sensitized solar<br />
cells.<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
28
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May 2011, Aix-en-Provence, France<br />
<br />
Fabrication Methods for the Manufacture of<br />
Sapphire Microparts<br />
David M. Allen, Roxana Redondo and Maximilien Dany<br />
Precision Engineering Centre, Cranfield University, Bedford MK43 0AL, UK<br />
Abstract- There is an increasing demand for microparts<br />
to be fabricated from an extremely hard-wearing,<br />
durable material such as sapphire but machining it to<br />
demanding specifications and tolerances poses<br />
considerable challenges. This paper describes<br />
experimental results obtained from laser machining and<br />
diamond machining of sapphire and concludes that, for<br />
optimum machining, a combination of these two<br />
techniques is required.<br />
I. INTRODUCTION<br />
The properties of sapphire, a form of alumina<br />
(Al 2 O 3 ), are attracting the attention of manufacturers<br />
for many different reasons. The material is the<br />
second hardest material known (9 on the Mohs<br />
hardness scale); second only to diamond (10 Mohs)<br />
and as such is “scratch-proof” and extremely durable<br />
with an exceptionally long service-life if used as a<br />
component. It is chemically inert and therefore<br />
resistant to attack by acidic and alkaline etchants. It<br />
also has exceptional optical properties being<br />
transparent in the infra-red, visible and ultra-violet<br />
regions of the electromagnetic spectrum from 170nm<br />
to 5500nm (see Fig. 1).<br />
It is worth noticing that various properties of sapphire<br />
such as hardness, dielectric constant, and thermal<br />
coefficient vary depending on the crystal orientation.<br />
For example, as shown in Table I, if the crystal’s<br />
orientation is perpendicular to the c-axis ( ┴ c-axis),<br />
the material is harder and a better insulator than if it<br />
is oriented parallel to the c-axis (║c-axis).<br />
However, the hardness of sapphire makes it a<br />
“difficult-to-machine” material. There is very little<br />
open literature, or even patents, on the methods of<br />
manufacturing sapphire parts from single crystal<br />
boules (ingots), although commercial companies<br />
obviously process single crystal sapphire by methods<br />
that are kept in-house as closely-guarded secrets. A<br />
paper published in 2010 shows a 500µm thick single<br />
crystal disk of sapphire that has been cut out by fine<br />
abrasive water jet machining [1]. However, it is<br />
acknowledged that the cut edge definition is not<br />
ideal, resulting in chipping as (quote) “small flakes<br />
were cut out from the edge”.<br />
Sapphire disks are frequently used in optical<br />
applications such as lenses or windows. To obtain the<br />
best optical quality, the most common crystalgrowing<br />
method is the Kyropolis method. The disks<br />
must then be cut and polished, which are two<br />
mechanical machining processes. A conventional<br />
approach to cut sapphire blocks involves the use of<br />
diamond abrasives bonded onto saw blades.<br />
Fig. 1. Transmission spectrum of a 2 mm thick sapphire window<br />
[2]<br />
However, one of the most recent tools developed to<br />
cut ceramic ingots or blocks, involves the use of<br />
wires coated with diamond abrasives. This technique<br />
is commonly applied in the form of multi-wire<br />
slicing, to cut very thin wafers of sapphire for the<br />
light emitting diodes (LED) industry. It is a fast<br />
process, though the loose grains may reduce the rate<br />
of material removal [3].<br />
The diamond abrasives used in wire cutting are<br />
bonded to a steel wire by nickel electroplating (Fig.<br />
2). According to the nomenclature of abrasives, the<br />
grit type is described using a letter and number<br />
system. The letter refers to the type of material,<br />
whilst the number refers to the average grain size of<br />
the abrasives, expressed as average diameter, or<br />
grains per unit area or volume. The grit type<br />
commonly used to wire-cut sapphire ranges between<br />
D 07 and D 91 grits, where D stands for diamond [3].<br />
29
Unfortunately, this fabrication method is slow and<br />
cannot readily be used to produce complex 2D and<br />
3D parts.<br />
TABLE I.<br />
PHYSICAL PROPERTIES OF SAPPHIRE [FROM 1]<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Property<br />
Value<br />
Density (kg/m 3 ) 3970<br />
Knoops Hardness (kg/mm 2 )<br />
1800 (║c-axis),<br />
2200 ( ┴ c-axis)<br />
Hardness in Mohs’ scale 9<br />
Young’s Modulus (GPa) 345<br />
Flexural Strength (kpsi) 100<br />
Compressive Strength (kpsi) 425<br />
Poisson’s Ratio 0.29 – 0.30<br />
Optical Transmission (nm) 170-5500<br />
Dielectric Constant<br />
11.53 (║c-axis),<br />
9.35 ( ┴ c-axis)<br />
Resistivity (ohms cm) 10·10 16<br />
Thermal Coefficient (1/°C)<br />
5.41·10 -6 (║c-axis),<br />
4.31·10 -6 ( ┴ c-axis)<br />
Melting Point (°C) 2050<br />
Boiling Point (°C) 2980<br />
Fig.2. SEM image of diamond abrasives on steel wire [3].<br />
One of the very few illustrations known of a sapphire<br />
micropart is a 0.25mm thick sapphire gear wheel<br />
made at Laser Zentrum Hannover, Germany (Fig. 3).<br />
It was used in a fluid sensor and was machined by<br />
multiple passes of higher harmonic 355nm<br />
wavelength radiation from a Nd:YAG laser with a<br />
high peak power intensity of 10 7 – 10 8 W/cm 2 .<br />
Successful machining was attributed to the short<br />
pulse length and superior beam quality [4]. However,<br />
a later publication acknowledged a problem<br />
associated with the effects of laser machining;<br />
namely “Among these undesired effects is the<br />
damaging of the rear side of thin wafers formed while<br />
surface structuring and drilling blind holes” [5].<br />
Fig. 3. A 0.25mm thick sapphire gear wheel made by multiple<br />
passes of 355nm wavelength laser pulses for high precision with<br />
no microcracking (Courtesy of A. Ostendorf and Laser Zentrum<br />
Hannover, Germany) [5]<br />
In an attempt to fabricate crack-free, extremely<br />
durable, optically-clear microcomponents, the authors<br />
have investigated and compared various methods of<br />
fabricating sapphire parts by both non-contact laser<br />
machining [6] and contact diamond machining [7].<br />
II. MATERIALS AND EQUIPMENT<br />
The manufacturing challenge is to fabricate<br />
microparts suitable for use in mechanical watches<br />
and instruments from sapphire discs 31mm in<br />
diameter and 1.2mm thick. Currently, scratch-proof<br />
watch “glasses” are made from sapphire and, very<br />
recently in January 2011, a new design of “seethrough”<br />
sapphire watch dial was advertised in the<br />
horological press. Apertures are stated to have been<br />
fabricated using a laser, but no technical details have<br />
been released, to the knowledge of the authors [8].<br />
In our research, several types of lasers have been<br />
used in order to machine the synthetic c-plane<br />
sapphire disks provided.<br />
As there is a vast range of laser beam machining<br />
systems available these days, collaborations with<br />
specialist laser facilities located in the UK and The<br />
Netherlands were instigated. The collaborating<br />
universities and companies were requested to drill<br />
circular holes of diameters ranging between 0.5 mm<br />
and 3.0 mm, and/or to machine a curve in the<br />
periphery of the disk. Due to the lack of experience in<br />
the machining of 1.2mm thick sapphire samples (such<br />
as those used throughout this project), the laser<br />
technology available for each of the lasers chosen<br />
was applied according to the interpretation of each of<br />
the specialist technicians involved in the machining<br />
processes.<br />
30
11-13 May 2011, Aix-en-Provence, France<br />
<br />
TABLE II.<br />
CHARACTERISTICS OF THE LASERS USED<br />
Type of Laser<br />
Used<br />
DPSS<br />
(University of Twente)<br />
Nd:YVO 4<br />
(MEC, Cardiff<br />
University)<br />
CVL<br />
(Oxford Lasers)<br />
Yb glass fibre<br />
(University of Cambridge)<br />
Repetition rate<br />
(kHz)<br />
400 50 10 2000<br />
Principal<br />
wavelengths (nm)<br />
343<br />
(uv)<br />
355<br />
(uv)<br />
511 (Green)<br />
578 (Yellow)<br />
1064 ± 10<br />
(ir)<br />
Pulse duration (s) 10·10 -12 < 12·10 -12 20·10 -9 0.7·10 -12<br />
Laser entrance side<br />
Laser exit side<br />
Fig. 4. Entrance (left) and exit (right) surfaces of sapphire disk after machining holes H1 (Φ = 2.5mm) and H2-H7 (Φ = 1.5mm) with the DPSS<br />
laser. Holes H1, H3 and H4 have been completely machined through the thickness of the disk after several passes. Holes H2, H5, H6 and H7<br />
have not been completely trepanned through. All holes at the exit side exhibit surface damage to some extent.<br />
Figure 5(a, left) showing a Nd:YVO 4 laser machined curve, 1.0 mm diameter hole and 21 slots on the laser entry side of the disk; (b, upper right)<br />
showing exit side of curve and slots with slight surface damage; (c, lower right) a sample showing exit side of hole and longer slots with similar<br />
surface damage.<br />
31
11-13 May 2011, Aix-en-Provence, France<br />
<br />
III. LASER MACHINING<br />
The details of the four different lasers used for<br />
micromachining sapphire are summarised in Table II.<br />
The diode-pumped solid state (DPSS) laser is a<br />
Trumpf TruMicro 5350 picosecond laser with<br />
maximum pulse energy of 50µJ, operating at 343nm;<br />
the most energetic of the laser wavelengths tested.<br />
Holes were trepanned as shown in Fig. 4. An<br />
additional observation that can be drawn from the<br />
figure is that the dimensions of the successfully<br />
drilled holes are different on both sides,<br />
demonstrating tapering. Hole H1 has a diameter of<br />
2.5 mm on the laser entrance side but only 2.1 mm on<br />
the exit side showing significant sidewall tapering.<br />
The most successful laser was the mode-locked<br />
Lumera Laser GmbH Nd:YVO 4 picosecond pulse<br />
duration laser operated at MEC in the uv range<br />
(355nm) as a third harmonic from the 1064nm<br />
wavelength. This laser has a maximum power of 2W<br />
and maximum pulse energy of 20µJ. Holes and a<br />
curved profile are shown in Fig.5. Less taper on the<br />
cut profiles was noted in comparison to the DPSS<br />
laser.<br />
The Oxford Lasers nanosecond pulse duration copper<br />
vapour laser (CVL) operates in the visible region of<br />
the electromagnetic spectrum with the intensity of the<br />
green wavelength (511nm) twice that of the yellow<br />
wavelength (578nm). Some machining was effected<br />
but the resolution was poor and fracturing at edges<br />
was prevalent in small holes of Φ = 0.1mm.<br />
However, it should be noted that even frequency<br />
doubled CVL radiation of 255nm wavelength has<br />
only been used successfully in the past to scribe<br />
sapphire to a maximum thickness of 90µm [9].<br />
The ytterbium-doped glass fibre laser has the highest<br />
repetition frequency and the longest wavelength (in<br />
the near ir) of the lasers investigated. The absorption<br />
of the 1064nm wavelength is therefore extremely low<br />
and the machining is ineffective over an acceptable<br />
maximum processing period of a few minutes.<br />
IV. DIAMOND MACHINING<br />
Typical sapphire components may require flats,<br />
chamfers and grooves to be machined and these<br />
features cannot be fabricated using laser technology.<br />
Diamond machining was therefore tested to<br />
determine whether it was a viable manufacturing<br />
technique. It should be noted that the surfaces<br />
produced needed to be extremely smooth and<br />
optically transparent to meet the aesthetic<br />
specifications in addition to those related to<br />
dimensions and tolerances. Diamond machining was<br />
carried out on a Kern Evo Machining Centre fitted<br />
with a prototype high speed Westwind air-bearing<br />
spindle capable of 350,000 rpm.<br />
Electroplated diamond pin tools (D76 and D126)<br />
were used for machining (Fig. 6). Phenolic resinbonded<br />
pin tools (D25 and D07) were utilised to<br />
reduce Sa but proved unsatisfactory due to high bond<br />
wear rate.<br />
Fig. 6. D76 electroplated diamond pin tool (Φ= 1mm)<br />
To machine a flat step, either the side or the tip of the<br />
diamond pin tool was used. In both cases, the<br />
sapphire disk was waxed onto a holder after an<br />
essential preheating to melt the wax. Then, the disk<br />
and the holder were mounted in the machine (on a<br />
pallet or in a vice); the surface of the disk being<br />
either horizontal (to use the tip of the tool) or vertical<br />
(to use the end of the tool).<br />
To machine a chamfer, the sapphire sample was<br />
waxed onto an aluminium set square holder. Firstly, a<br />
flat step was machined on the sample, then the holder<br />
was rotated by 45° and mounted on the pallet<br />
(without having to unwax the sapphire sample) and<br />
the chamfer was machined as shown in Fig.7.<br />
Fig. 7. Sapphire disk mounted on aluminium set square holder to<br />
machine a 45º chamfer using the tip of the tool [7].<br />
Machining parameters included; spindle rotation<br />
speed, feed rate, depth of cut, tool positioning used<br />
32
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
(machining with the side or with the tip), coolant<br />
and tool wear.<br />
The machining parameters were therefore chosen as<br />
follows: For the spindle rotation speed and the feed<br />
rate, the values chosen for the experimental tests were<br />
varied around the recommended values as shown in<br />
Table III.<br />
• For the depth of cut, most of the time a value of<br />
0.02 mm was chosen, even though a lower value<br />
was sometimes tested (especially for the<br />
finishing operations). Using smaller depth of cut<br />
would probably improve the surface roughness,<br />
but it would also increase the time of machining.<br />
• For the tool positioning, both the side and the tip<br />
were used.<br />
• Air and water-based coolants were used. White<br />
spirit was used once only as it was not<br />
environment-friendly.<br />
• For evaluating the influence of tool wear, some<br />
operations were performed twice, using the same<br />
machining parameters, first with a used tool, then<br />
with an unused tool.<br />
TABLE III.<br />
PARAMETERS RECOMMENDED BY TOOL MANUFACTURER<br />
Tool<br />
Spindle rotation Feed rate Depth of<br />
diameter<br />
speed (rpm) (mm/min) cut (mm)<br />
(mm)<br />
0.5 18,000 1.0 0.02<br />
1.0 40,000 1.0 0.02<br />
3.0 60,000 1.0 0.02<br />
3.5 60,000 1.0 0.02<br />
After grinding the sapphire samples, the surface<br />
quality was measured. The best way to analyse the<br />
surface quality is to measure the depth of sub-surface<br />
damage of the surface machined. However, the<br />
surface roughness of a machined area can be<br />
correlated with its sub-surface damage [10]. Since a<br />
lot of samples had to be measured, and because it is<br />
easier to measure a surface roughness than a depth of<br />
sub-surface damage, the surface average roughness<br />
Sa was measured to assess the surface quality of the<br />
machined areas.<br />
The machine used to measure the average roughness<br />
(Sa) was a Talysurf CCI 6000 (white light<br />
interferometer). The lowest value of Sa recorded was<br />
65nm. A magnification of x50 was chosen, which<br />
resulted in a 360µm x 360µm window analysis. The<br />
fractures and the dimensions of the features machined<br />
were measured using the SEM (Scanning Electron<br />
Microscope). This was also a good method to analyse<br />
the aesthetic appearance of the features.<br />
The results appeared to be completely different using<br />
the side or the tip of the tool. Indeed, a better surface<br />
roughness was usually achieved using the tip of the<br />
tool. However, machining with the side of the tool<br />
resulted in a homogenous result, while machining<br />
with the tip of the tool resulted in a nonhomogeneous<br />
result with more fractures. This is<br />
explained for two main reasons:<br />
• Firstly, when machining with the tip of the tool,<br />
the contact area between the tool and the<br />
workpiece is a disk, in which the cutting speed is<br />
not uniform. Indeed, the centre of the tool is not<br />
rotating, and the cutting speed increases with the<br />
radial position up to a maximum value on the<br />
edge of the tool. The middle part of the tool is<br />
therefore not cutting sapphire but rubbing on<br />
sapphire, which most probably chips off some<br />
material. When machining with the side of the<br />
tool, the contact area is theoretically a line where<br />
the cutting speed is uniform.<br />
• Secondly, when machining with the tip of the<br />
tool, since the contact area is a disk, machined<br />
material can get trapped between the tool and the<br />
workpiece, fracturing the sample. This problem<br />
does not occur when machining with the side of<br />
the tool, where the machined material can easily<br />
be removed from the working area.<br />
A comparison between the two machining methods is<br />
shown in Fig. 9. It can be seen that the side-machined<br />
surface has no fractures with a better aesthetic<br />
appearance than the tip-machined surface, but it also<br />
has a higher average roughness than the tip-machined<br />
surface. The average roughness of 810nm for the<br />
side-machined surface is mainly the result of the tool<br />
profile that has shaped the surface. Although the tipmachined<br />
surface has a better average roughness of<br />
690nm, it has a higher light scatter and a resultant<br />
lower aesthetic quality.<br />
Fig. 8. 0.5mm diameter hole diamond machined in sapphire [7].<br />
Using a 0.5mm diameter D76 pintool as a drill, some<br />
well-defined holes (Fig. 8) were also fabricated by<br />
drilling half-way through the disk, turning it over and<br />
registering a second machining operation with the<br />
first operation. This prevented edge fractures.<br />
33
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Fig. 9. (Top) SEM image and CCI analysis of surface F (machined with the side of the tool) with Sa = 810 nm;<br />
(Bottom) SEM picture and CCI analysis of surface G (machined with the tip of the tool) with Sa = 690 nm [7].<br />
V. CONCLUSIONS<br />
Rapid laser machining using uv radiation appears to<br />
be an acceptable method for roughing out microparts<br />
by profiling and perforating sapphire<br />
disks thicker<br />
than 1mm. The use of visible and near ir radiation<br />
with wavelengths between 400nm and<br />
2000nm does<br />
not seem effective for machining thick sapphire but it<br />
has been reported previously that 532nm radiation<br />
from a Trumpf laser was used to pattern the back of<br />
430µm thick sapphire dies [11]. Fabrication of vias<br />
appears to be possible but considerable cracking is<br />
apparent.<br />
Diamond machining is slow but can finish tapered<br />
edges formed by laser cutting and impart a sub-<br />
damage on<br />
micron surface finish by removing laser<br />
the beam exit side of sapphire disks. Further work on<br />
reducing Sa below 65nm will investigate the<br />
possibility of using resin-bonded tools with lower<br />
wear rates than those used in this research.<br />
The conclusion has therefore been reached that for<br />
efficient sapphire micropart production, uv lasers<br />
should be utilised for rapid roughing out of profiles<br />
and holes and the slower processs of diamond<br />
machining should be utilised for finishing the parts to<br />
an acceptable surface finish and aesthetic quality.<br />
However, it is highly probable that manual polishing<br />
could be a most effective final finishing<br />
process.<br />
ACKNOWLEDGMENTS<br />
The research was financed by the EC<br />
FP7 “Integmicro”<br />
collaborative project: CP-IP 214013-2; New<br />
production technologies of complex 3D<br />
microdevices<br />
through multiprocess integration of ultra precision<br />
engineering techniques. We wish to thank our FP7<br />
partners, Westwind, Kern and Swatch for their help<br />
and collaboration and to acknowledge the laser<br />
machining collaboration of Prof J. Meijer (University<br />
of Twente, The Netherlands) ); Dr P. Jefferies<br />
(Manufacturing Engineering Centre, Cardiff<br />
University, Wales); Dr A. Ferguson (Oxford Lasers<br />
Ltd., Didcot, England); Dr W. O’Neill (Centre for<br />
Industrial Photonics, Institute for Manufacturing,<br />
University of Cambridge, England) and, for the<br />
diamond machining of sapphire disks, Mr J. Hedge<br />
(Precision Engineering Centre, Cranfield University,<br />
England).<br />
REFERENCES<br />
[1] T. Aklint, P. Johander, K. Brinkfeldt, , C. Ojmertz and T. Ryd,<br />
Abrasive waterjet cutting for micro manufacture, Proc. 7 th<br />
International Conference on Multi-Material Micro Manufacture,<br />
Bourg en Bresse and Oyonnax, France, November 2010, 147-150.<br />
[2] J.S. Tydex Co., 2009.<br />
[3] Ceramic Industry, 2005.<br />
[4] J. Meijer et al, Laser machining by short and ultrashort pulses,<br />
state of the art and new opportunities in<br />
the age of the photons,<br />
CIRP Annals, 51/2, 531-550 (2002).<br />
[5] A. Ostendorf, T. Temme and K. Samm, Proc. 5 th euspen<br />
International Conference, Montpellier, France, May 2005, 719.<br />
[6] R. Redondo, Micromachining of Sapphire, MSc thesis,<br />
Cranfield University, UK, September 2009<br />
[7] M. Dany, Mechanical Micromachining of Sapphire, MSc<br />
thesis, Cranfield University, UK, September 2010<br />
[8] http://www.lacotedesmontres.com/No_ _8257.htm<br />
[9] D. Karnakis, E.K. Lilly, M.R.H. Knowles, E. Gu and M.D.<br />
Dawson, High throughput scribing for the t manufacture of LED<br />
components, Proc. Photonics West 2004: Lasers and applications<br />
in science and engineering. Proc SPIE 5366, 207; doi<br />
10.1117/12.531685<br />
[10] P.P. Hed and D.F. Edwards, Relationship between surface<br />
roughness and subsurface damage during<br />
grinding of optical glass<br />
with diamond tools, Applied Optics, 26, 4677-4680 (1987)<br />
[11] J. Vignes, F. Haring, S.S. Ahmad, K. Gerstner and A.<br />
Reinholz, Laser patterning and via drilling of sapphire wafers and<br />
die, Proc. 43 rd Int. Symp. on Microelectronics, IMAPS, North<br />
Carolina, USA, November 2010, 000513-000520.<br />
34
!<br />
11-13 May 2011, Aix-en-Provence, France<br />
Characterisation and Comparison of Water and<br />
Alcohol as Catalysts in Vapour Phase HF Etching of<br />
Silicon Oxide Films<br />
D. Drysdale 1 , T. O’Hara 2 , C. H. Wang 1<br />
1<br />
School of Engineering & Physical Sciences, Heriot-Watt University, Edinburgh, EH14 4AS, UK<br />
3<br />
memsstar Ltd, Scottish Microelectronics Centre, University of Edinburgh, Edinburgh, EH9 3JF, UK<br />
Abstract- The comparison of etch rates and selectivities for thin<br />
films of silicon dioxide and silicon nitride with respect to water<br />
and alcohol based (ethanol in this case) catalysts in a vapour phase<br />
HF etching process is discussed. Observation of etch rates for both<br />
PECVD Oxide and Nitride films are used to describe the<br />
behaviour of silicon dioxide etching. These behaviour<br />
characteristics can also be used to develop selectivity behaviours<br />
between the two films based on each of the catalysts. A number of<br />
factors are considered in the vapour phase etching process: the<br />
total gas flow for the etching process, process temperature and the<br />
etching pressure. The paper discusses the differences between<br />
both water and ethanol as process catalysts for the improvement<br />
of silicon dioxide etching selectivity with respect to silicon nitride.<br />
Results show that using water as a catalyst, a selectivity of up to<br />
40:1 can be achieved while with a direct comparison of the same<br />
etch process with ethanol, the highest achievable selectivity is 15:1.<br />
On the other hand, with comparable etch rates to that of the<br />
water catalyst process, the highest selectivity achieved was 10:1.<br />
I. INTRODUCTION<br />
The use of anhydrous HF vapour as an etchant has become<br />
commonplace within microelectromechanical systems<br />
(MEMS). It is typically used in the production of free-standing<br />
structures for a range of MEMS devices such as RF switches,<br />
accelerometers and microphones. From the initial etching<br />
technique using acid baths as described by Holmes and Snell<br />
[1] with further study by G. Van Barel et al. [2],[3] in<br />
understanding behaviour within wet processing to the recent<br />
work of vapour phase HF etching by Witvrouw et al. [4]. The<br />
use of HF for etching silicon dioxide has become a standard<br />
process technique and as its impact becomes more prominent<br />
with the growth of the MEMS industry in fabricating many<br />
devices. A critical process in MEMS production is the<br />
integration of the release process to the existing semiconductor<br />
fabrication processes.<br />
Many large-scale fabrication facilities still work solely with<br />
CMOS processes and materials thus developing new etching<br />
options that integrate well with these standardised methods<br />
reduces the difficulty of developing next generation MEMS<br />
devices. Many of today’s modern MEMS devices typically<br />
require one or more of five common CMOS materials:<br />
aluminium, silicon, polysilicon, silicon dioxide and silicon<br />
nitride. While HF etching is not a problem in terms of<br />
selectivity to the first three materials, problems arise in using<br />
an HF etch which is selective to silicon dioxide with respect to<br />
silicon nitride. The use of silicon dioxide as the sacrificial<br />
material of choice is for many reasons; typically used as<br />
passivation and insulating layers as well as dielectric layers for<br />
a device depending on the thickness of the film.<br />
The main focus of silicon dioxide in this study is its role as a<br />
sacrificial layer for the fabrication of MEMS devices such as<br />
microphones and RF MEMS switches. It is however, common<br />
to have silicon nitride layers and silicon dioxide layers stacked<br />
on top of each other with the nitride layer forming part of the<br />
structural or functional part of the device while the oxide acts<br />
as a sacrificial layer to be removed to realise the final freestanding<br />
structure.<br />
While a wet process can typically be used, this generates a<br />
widely experienced phenomenon called stiction [5],[6]. Stiction<br />
is the near permanent adhesion of two surfaces commonly due<br />
to electrostatic forces, hydrogen bonding and Van der Waals<br />
forces. Should the restoration force of a movable structure be<br />
less than the adhesion force being applied to it by an external<br />
source (such as a bead of moisture), it will adhere reducing<br />
yield of functional devices and causing high levels of device<br />
failures. Stiction commonly occurs due to moisture present in<br />
micron scale devices as scaling laws suggest that even a single<br />
droplet of moisture applies a strong enough force to hold a<br />
released structure. This is often seen as the other key reason in<br />
the push to vapour phase processing from wet processing<br />
commonly employed the world over. By etching in a vapour<br />
phase, moisture generation is reduced and stiction is therefore<br />
reduced which in turn creates a higher yield for devices. The<br />
equation for etching silicon dioxide is defined as:<br />
catalyst catalyst<br />
SiO2 + 4HF ! SiF4 + 2H2O (1)<br />
This equation requires a catalyst for the reaction to begin and<br />
is typically considered to be either water or alcohol. As can be<br />
seen by the reaction, water is generated as a by-product and<br />
while the presence of too much water or any moisture within<br />
the reaction chamber can damage product wafers, its presence<br />
is needed not only to initiate the reaction, but to maintain it. It<br />
is the ability not only to remove the excess water generated but<br />
to control the other process factors of the etch to keep a high<br />
and repeatable etch process that is the key to its success. By<br />
studying the behaviour of these two key materials, it is hoped<br />
that a better understanding of the etching behaviour can be<br />
achieved thus helping future designs for MEMS devices.<br />
<br />
35
!<br />
II.<br />
11-13 May 2011, Aix-en-Provence, France<br />
EXPERIMENTATION PROCEDURE<br />
III. EXPERIMENTAL SET-UP<br />
Films of silicon oxide and silicon nitride were fabricated for<br />
experimentation to help understand the etching behaviours<br />
during HF etching more clearly. As the study focused on oxide<br />
and nitride films, it was decided that the a practical choice<br />
would that of PECVD oxide and PECVD nitride which were<br />
deposited by an STS multiplex CVD tool. The films had an<br />
approximate thickness of 9300Å and fabricated on 150mm<br />
silicon wafers. Test chips were obtained by dicing the<br />
wafers into 2cm 2 die.<br />
The 2cm 2 die of oxide and nitride were placed side by side<br />
on a clean 6 inch silicon wafer in the process module handler.<br />
The die were then passed into the process module and<br />
processed in the chamber with a recipe based various process<br />
factors. These factors are: aHF etchant gas flow, nitrogen<br />
buffer gas flow, catalyst carrier gas flow, process temperature<br />
and process pressure. A standard “base” recipe was used for<br />
both catalysts. This recipe consisted of a total process gas flow<br />
of 350 standard cubic centimetres per minute (sccm); 150 sccm<br />
aHF etchant gas, 150 sccm nitrogen buffer, and 50 sccm of<br />
nitrogen to be used as a carrier gas for the catalyst bubbler. The<br />
bubbler itself was calibrated so that with a flow of 50 sccm<br />
nitrogen carrier gas, a flow of 50 sccm of catalyst was also<br />
introduced into the process module. The catalyst bubbler had a<br />
carrier gas flowing during the entirety of the etch process.<br />
The temperatures used in this study were a temperature of<br />
25! and a low temperature of 10!. This would provide a<br />
behaviour characteristic for the process at a standard<br />
temperature (25! ) and of a low temperature (10!) for the<br />
process based on the effects of temperature on etching as<br />
discussed previously by J. Anguita, F. Briones [7]. All recipes<br />
would use the same gas flows and etch times with only three<br />
variable factors: the process pressure, the process temperature<br />
and the catalyst carrier gas flow. The carrier gas flows would<br />
be varied between 25 sccm, 50 sccm and 75 sccm to analyse<br />
the effects of different catalyst flows. To maintain the same<br />
total gas flows throughout the each recipe, the nitrogen buffer<br />
gas flow would be varied accordingly along with the catalyst<br />
carrier gas from 175 sccm to 150 sccm and 125 sccm<br />
respectively as the carrier gas flows increased. The two<br />
pressure regimes being used would be a low and high pressure<br />
regime which would allow the analysis of a slow and fast etch<br />
process for both catalysts.<br />
While the first study used the same “base” recipe for both<br />
catalysts as stated above, a second comparable etch process<br />
was also developed for the study whereby the etch rate for the<br />
ethanol process was increased to match that of the water<br />
process. This was necessary in order to obtain a more<br />
conclusive study of etching behaviours for the ethanol process.<br />
The process pressure would be dictated by the pressure<br />
required to achieve a comparable etch rate for both catalysts<br />
with the standard “base” recipe.<br />
For each recipe, pairs of 2cm 2 oxide and nitride die were ran<br />
for 2 and 3 minutes separately. This allowed a a determination<br />
of etch rate for based on etch time. From these results, a value<br />
for selectivity to be obtained for the etch of oxide film against<br />
nitride for this time period.<br />
The experiments were carried out on a memsstar® aHF SVR<br />
platform. In the case of these experiments, along with the use<br />
of anhydrous HF (aHF), nitrogen is used as a buffer gas.<br />
Nitrogen is also used as the carrier gas for the bubbler system<br />
which is used as the source of the process catalyst. Inside the<br />
process module (PM), the aHF and catalyst are introduced via<br />
separate gas distribution ring systems which allow an even<br />
distribution of the reactant gasses. Inside the process module is<br />
a temperature controlled pedestal. This allows for the<br />
temperature at which the sample die sit at within the process<br />
module to be set during the etch process. A basic schematic for<br />
the hardware configuration is shown below:<br />
Fig. 1. Diagram of the experimental set-up.<br />
Fig. 2. Image of the memsstar® SVR platform.<br />
IV.<br />
RESULTS<br />
The experimental details described in section II and III were<br />
used to obtain the results presented in this section. Using fixed<br />
process gas flows and adjusting the catalyst carrier gas flows,<br />
etch process temperature and process pressure, a series of<br />
results were generated to show the behaviour of selectivity for<br />
the two catalysts used in these experiments. To make sure that<br />
the pressures used in the different pressure regimes allow a fair<br />
comparison of the etching behaviour, the total amount of oxide<br />
etched at low pressure with the defined “base” recipe for 3<br />
minutes was set at 1800 Å and at high pressure approximately<br />
<br />
36
!<br />
11-13 May 2011, Aix-en-Provence, France<br />
5700 Å. The etch rates were used as a reference for both the<br />
'!!!"<br />
standard process and the comparable processes developed later<br />
&#!!"<br />
for the ethanol catalyst. After each etching run, the oxide and<br />
&!!!"<br />
nitride chips were placed on a hot plate for 30 seconds at<br />
%#!!"<br />
200! to remove any residue that can be formed on a silicon<br />
%!!!"<br />
$#!!"<br />
nitride surface during HF etching as described in [2].<br />
A. PECVD Oxide and Nitride with Water Catalyst:<br />
At 25!, a low pressure of 8 Torr was used and for a high<br />
pressure process, 11T was satisfactory. This was the standard<br />
“base” recipe of 150 sccm aHF, with the remaining gas flow<br />
from the nitrogen buffer and catalyst flows dependant on the<br />
flow of the carrier gas. That is: 25 sccm and 175 sccm, 50 sccm<br />
and 150 sccm and 75 sccm and 125 sccm respectively for the<br />
carrier gas and nitrogen buffer gas flows. For etching at 10!,<br />
a low pressure of 3.5 Torr and a high pressure of 4.5 Torr were<br />
used. The gas flow of the carrier gas was again varied to study<br />
the behaviour of the catalysts.<br />
!"#$%&'()'*#+%,-.%<br />
$!!!"<br />
#!!"<br />
!"<br />
!" %!" '!" (!" )!"<br />
/'(()+(%0'1%2345%<br />
*+,"<br />
-./001./"<br />
2345/"<br />
*+,"-./01./"<br />
647.45/"<br />
849:"<br />
-./001./"<br />
2345/"<br />
849:"<br />
-./001./"<br />
647.45/"<br />
Fig. 6. Showing the etch variance between a 2 minute and 3 minute etch of<br />
Silicon Oxide to Silicon Nitride in a 10! etch regime at high and low<br />
pressures as a function of water catalyst carrier gas flow.<br />
B. PECVD Oxide and Nitride with Ethanol Catalyst:<br />
A direct comparison of the etch of silicon nitride and oxide<br />
could be conducted using the same gas flows as the water<br />
catalyst tests. At 25!, a high pressure of 11 Torr was used and<br />
a low pressure of 8 Torr was used. At 10!, the high and low<br />
pressures were 4.5 Torr and 3.5 Torr respectively.<br />
!"#"$%&'()*<br />
'%"<br />
'$"<br />
'#"<br />
'!"<br />
&"<br />
%"<br />
$"<br />
#"<br />
!"<br />
!" #!" $!" %!" &!"<br />
+,--'"-*.,/*0#12*3/$$45*<br />
()*"<br />
+,-../,-"<br />
0123"<br />
+,-../,-"<br />
Fig. 3. Selectivity of Silicon Oxide to Silicon Nitride in a 25! etch regime<br />
at high and low pressures as a function of water catalyst carrier gas flow.<br />
!"#$%&'()'*#+%,-.%<br />
&!!!"<br />
%#!!"<br />
%!!!"<br />
$#!!"<br />
$!!!"<br />
#!!"<br />
!"<br />
!" %!" '!" (!" )!"<br />
/'(()+(%0'1%2345%,1##6.%<br />
*+,"<br />
-./001./"<br />
2345/"<br />
*+,"<br />
-./001./"<br />
647.45/"<br />
849:"<br />
-./001./"<br />
2345/"<br />
849:"<br />
-./001./"<br />
647.45/"<br />
Fig. 4. Showing the etch variance between a 2 minute and 3 minute etch of<br />
Silicon Oxide to Silicon Nitride in a 25! etch regime at high and low<br />
pressures as a function of water catalyst carrier gas flow.<br />
!"#"$%&'()*<br />
'#"<br />
'!"<br />
&#"<br />
&!"<br />
%#"<br />
%!"<br />
$#"<br />
$!"<br />
#"<br />
!"<br />
!" %!" '!" (!" )!"<br />
+,--'"-*.,/*0#12*3/$$45*<br />
*+,"<br />
-./001./"<br />
2345"<br />
-./001./"<br />
Fig. 5. Selectivity of Silicon Oxide to Silicon Nitride in a 10! etch regime<br />
at high and low pressures as a function of water catalyst carrier gas flow.<br />
!"#"$%&'()*<br />
"%$<br />
&#$<br />
&%$<br />
#$<br />
%$<br />
!#$<br />
%$ "%$ '%$ (%$ )%$<br />
!&%$<br />
!&#$<br />
!"%$<br />
!"#$<br />
+,--'"-*.,/*0#12*3/$$45*<br />
*+,$<br />
-./001./$<br />
2345$<br />
-./001./$<br />
Fig. 7. Selectivity of Silicon Oxide to Silicon Nitride in a 25! etch regime<br />
at high and low pressures as a function of ethanol catalyst carrier gas flow.<br />
!"#$%&'()'*#+%,-.%<br />
&"#$<br />
&##$<br />
%"#$<br />
%##$<br />
"#$<br />
#$<br />
#$ &#$ '#$ (#$ )#$<br />
!"#$<br />
/'(()+(%0'1%2345%,1##6.%<br />
*+,$<br />
-./001./$<br />
2345/$<br />
*+,$<br />
-./001./$<br />
647.45/$<br />
849:$<br />
-./001./$<br />
2345/$<br />
849:$<br />
-./001./$<br />
647.45/$<br />
Fig. 8. Showing the etch variance between a 2 minute and 3 minute etch of<br />
Silicon Oxide to Silicon Nitride in a 25! etch regime at high and low<br />
pressures as a function of ethanol catalyst carrier gas flow.<br />
!"#"$%&'()*<br />
&#$<br />
"%$<br />
"#$<br />
%$<br />
#$<br />
#$ &#$ '#$ (#$ )#$<br />
!%$<br />
!"#$<br />
+,--'"-*.,/*0#12*3/$$45*<br />
*+,$-./001./$<br />
2345$-./001./$<br />
Fig. 9. Selectivity of Silicon Oxide to Silicon Nitride in a 10! etch regime<br />
at high and low pressures as a function of the catalyst carrier gas flow.<br />
<br />
37
!<br />
!"#$%&'()'*#+%,-.%<br />
(##$<br />
'#$<br />
&#$<br />
%#$<br />
"#$<br />
#$<br />
!"#$<br />
#$ "#$ %#$ &#$ '#$<br />
/'(()+(%0'1%2345%<br />
)*+$<br />
,-.//0-.$<br />
1234.$<br />
)*+$<br />
,-.//0-.$<br />
536-34.$<br />
7389$<br />
,-.//0-.$<br />
1234.$<br />
7389$<br />
,-.//0-.$<br />
536-34.$<br />
Fig. 10. Showing the etch variance between a 2 minute and 3 minute etch of<br />
Silicon Oxide to Silicon Nitride in a 10! etch regime at high and low<br />
pressures as a function of the catalyst carrier gas flow.<br />
C. PECVD Oxide and Nitride with Ethanol Catalyst<br />
(Comparable Etch Rate Process Comparison):<br />
A discussed in section II, a process to produce a comparable<br />
etch rate to the water catalyst process was developed to analyse<br />
the effects of the selectivity for the ethanol catalyst process.<br />
This was carried out as there were no clear behaviour<br />
characteristics from the direct recipe comparison of ethanol<br />
from the previous section. For a low pressure regime, the<br />
criteria was that with the “base” recipe (150 sccm aHF, 150<br />
sccm nitrogen buffer and 50 sccm catalyst carrier gas etched<br />
and etching time of 3 minutes), the total amount of etched<br />
oxide should equal in the region of 1800Å. For a high pressure<br />
regime, the total etched oxide for the process should equal<br />
approximately 5700Å. For the 25! process, a low pressure of<br />
24 Torr was used and a high pressure of 35 Torr was required.<br />
At 10!, a low pressure of 12 Torr provided the right value for<br />
total etched oxide and at high pressure, a pressure of 28 Torr<br />
was used.<br />
!"#"$%&'()*<br />
'#"<br />
'!"<br />
&"<br />
%"<br />
$"<br />
#"<br />
!"<br />
!" #!" $!" %!" &!"<br />
+,--'"-*.,/*0#12*3/$$45*<br />
()*"<br />
+,-../,-"<br />
0123"<br />
+,-../,-"<br />
Fig. 11. Selectivity of Silicon Oxide to Silicon Nitride in a 25! etch regime<br />
at high and low pressures as a function of ethanol catalyst carrier gas flow.<br />
%#!!"<br />
11-13 May 2011, Aix-en-Provence, France<br />
!"#"$%&'()*<br />
'#"<br />
'!"<br />
&"<br />
%"<br />
$"<br />
#"<br />
!"<br />
!" #!" $!" %!" &!"<br />
+,--'"-*.,/*0#12*3/$$45*<br />
()*"<br />
+,-../,-"<br />
0123"<br />
+,-../,-"<br />
Fig. 13. Selectivity of Silicon Oxide to Silicon Nitride in a 10! etch regime<br />
at high and low pressures as a function of ethanol catalyst carrier gas flow.<br />
!"#$%&'()'*#+%,-.%<br />
#!!!"<br />
'&!!"<br />
'%!!"<br />
'$!!"<br />
'#!!"<br />
'!!!"<br />
&!!"<br />
%!!"<br />
$!!"<br />
#!!"<br />
!"<br />
!" #!" $!" %!" &!"<br />
/'(()+(%0'1%2345%,1##6.%<br />
()*"<br />
+,-../,-"<br />
0123-"<br />
()*"<br />
+,-../,-"<br />
425,23-"<br />
6278"<br />
+,-../,-"<br />
0123-"<br />
6278"<br />
+,-../,-"<br />
425,23-"<br />
Fig. 14. Showing the etch variance between a 2 minute and 3 minute etch of<br />
Silicon Oxide to Silicon Nitride in a 10! etch regime at high and low<br />
pressures as a function of ethanol catalyst carrier gas flow.<br />
TABLE 1<br />
SUMMARY OF SELECTIVITIES WITH DIRECT COMPARISON OF ETCH PROCESS<br />
Pressure/<br />
Carrier Gas<br />
Flow<br />
Water, 25!<br />
Ethanol,<br />
25!<br />
8T/25sccm 2.8 2.2<br />
8T/50sccm 7.1 2.5<br />
8T/75sccm 5.6 2.2<br />
11T/<br />
25sccm<br />
11T/<br />
50sccm<br />
11T/<br />
75sccm<br />
12.5 1.5<br />
13.9 -21.6<br />
14 15.3<br />
Pressure/<br />
Carrier Gas<br />
Flow<br />
3.5T/<br />
25sccm<br />
3.5T/<br />
50sccm<br />
3.5T/<br />
75sccm<br />
4.5T/<br />
25sccm<br />
4.5T/<br />
50sccm<br />
4.5T/<br />
75sccm<br />
Water, 10!<br />
Ethanol,<br />
10!<br />
17.6 -8.8<br />
17.4 6.1<br />
21.5 3<br />
33.9 5.6<br />
37.8 -5<br />
39.5 14<br />
TABLE 2<br />
SUMMARY OF SELECTIVITIES OF WATER CATALYST PROCESS AND COMPARABLE<br />
ETCH ETHANOL PROCESS AT 25!<br />
Pressure/<br />
Carrier Gas<br />
Flow<br />
Water, 25!<br />
8T/25sccm 2.8<br />
8T/50sccm 7.1<br />
Pressure/<br />
Carrier Gas<br />
Flow<br />
24T/<br />
25sccm<br />
24T/<br />
50sccm<br />
Ethanol,<br />
25!<br />
10<br />
6.1<br />
!"#$%&'()'*#+%,-.%<br />
%!!!"<br />
$#!!"<br />
$!!!"<br />
#!!"<br />
!"<br />
!" %!" &!" '!" (!"<br />
/'(()+(%0'1%2345%,1##6.%<br />
)*+"<br />
,-.//0-."<br />
1234."<br />
)*+"<br />
,-.//0-."<br />
536-34."<br />
7389"<br />
,-.//0-."<br />
1234."<br />
7389"<br />
,-.//0-."<br />
536-34."<br />
Fig. 12. Showing the etch variance between a 2 minute and 3 minute etch of<br />
Silicon Oxide to Silicon Nitride in a 25! etch regime at high and low<br />
pressures as a function of ethanol catalyst carrier gas flow.<br />
8T/75sccm 5.6<br />
11T/<br />
25sccm<br />
11T/<br />
50sccm<br />
11T/<br />
75sccm<br />
12.5<br />
13.9<br />
14<br />
24T/<br />
75sccm<br />
35T/<br />
25sccm<br />
35T/<br />
50sccm<br />
35T/<br />
75sccm<br />
6.7<br />
7.9<br />
8<br />
5.5<br />
<br />
38
!<br />
TABLE 3<br />
SUMMARY OF SELECTIVITIES OF WATER CATALYST PROCESS AND COMPARABLE<br />
ETCH ETHANOL PROCESS AT 10!<br />
Pressure/<br />
Carrier Gas<br />
Flow<br />
3.5T/<br />
25sccm<br />
3.5T/<br />
50sccm<br />
3.5T/<br />
75sccm<br />
4.5T/<br />
25sccm<br />
4.5T/<br />
50sccm<br />
4.5T/<br />
75sccm<br />
Water, 10!<br />
17.6<br />
17.4<br />
21.5<br />
33.9<br />
37.8<br />
39.5<br />
Pressure/<br />
Carrier Gas<br />
Flow<br />
12T/<br />
25sccm<br />
12T/<br />
50sccm<br />
12T/<br />
75sccm<br />
28T/<br />
25sccm<br />
28T/<br />
50sccm<br />
28T/<br />
75sccm<br />
Ethanol,<br />
10!<br />
9.6<br />
9.8<br />
9.1<br />
7.1<br />
6.4<br />
6.2<br />
11-13 May 2011, Aix-en-Provence, France<br />
the oxide and nitride films. The etch rate of the oxide chips was<br />
therefor much more substantial at higher pressures.<br />
Changing the flow of the catalysts was another factor which<br />
was changed throughout these experiments. From these<br />
experiments, an etching behaviour could be determined for this<br />
process factor. The effect of increasing the flow during the<br />
process was seen to help increase the etch rate of both the<br />
oxide and nitride films, but not nearly to the extent that was<br />
expected. It was also clear from the results with the water<br />
catalyst that changing the catalyst carrier gas flows at low<br />
temperatures had a much more linear affect to the etch rate than<br />
at a higher temperature. It is not obvious from these<br />
experiments as to whether this effect is tied solely to the<br />
increase in catalyst to the etch of if it is tied also to the lower<br />
processing temperature which inherently provides a much<br />
faster etch as discussed previously.<br />
V. DISCUSSION<br />
A. PECVD Oxide and PECVD Nitride with Water Catalyst:<br />
The etching of silicon oxide and nitride films with a water<br />
catalyst seen to provide the greatest overall selectivities during<br />
this study. As is commonly reported, the etch rate of silicon<br />
oxide films at both 25! and 10! is faster. This is typically<br />
due to the oxide films have a higher moisture content than their<br />
nitride counterparts. The effect is much more pronounced<br />
during 10! etching process also as at lower temperatures there<br />
tends to be a higher moisture content within the process<br />
module than at 25! as the higher temperature tends to drive<br />
out any additional moisture which resides on or in the films.<br />
This is why a slower etch is generally observed for oxide and<br />
nitride films at higher temperatures. The key observation here<br />
is that while the nitride etches faster with an increase in<br />
temperature at both 10! and 25!, the amount of nitride<br />
actually etched at lower temperatures is markedly less at lower<br />
temperatures. This suggests that with an increase in available<br />
water vapour in the chamber because of the lower temperature,<br />
the oxide is then preferentially etched over the nitride film.<br />
While the oxide film may have etched to approximately equal<br />
amounts in both low temperature and high temperature<br />
conditions, nitride etching can be lessened because the water<br />
vapour will tend to etch the oxide more readily than the nitride.<br />
This much more preferential etching of oxide over nitride at<br />
low temperatures is why a much higher selectivity between the<br />
two films is observed here.<br />
With respect to process pressure, it is also accepted that<br />
higher pressures produce a faster etch process than lower<br />
pressures. This is associated with the increase in residence time<br />
of the etchant gasses. To maintain a higher pressure within the<br />
process module, the chamber pumping line is more highly<br />
restricted meaning that the etch gasses have a longer period of<br />
time in which to react with the surface of the structures. With a<br />
greater level of reaction, a higher level of resultant by product<br />
is created of which water vapour is one. By creating more<br />
water vapour in the chamber which is not pumping as strongly<br />
as at lower pressures, this additional water vapour produced<br />
will add to the etch process producing a faster etch rate for both<br />
B. PECVD Oxide and PECVD Nitride with Ethanol Catalyst<br />
(Direct Water Catalyst Process Comparison):<br />
Comparing the water process directly with the ethanol<br />
etching process using the same process parameters allows a<br />
look into the main differences of the etch. What was observed<br />
were results that did not typically fit known behaviours for<br />
silicon oxide etching.<br />
Firstly, the highest selectivities were observed at higher<br />
temperatures which is unlike the results seen with water as a<br />
catalyst. This may however, not be correct to presume as there<br />
was very little etching took place in the ethanol etch processes<br />
with which to build an accurate picture. What we can say from<br />
this etching process is that using ethanol in a direct comparison<br />
of processes cannot provide as fast an oxide etch as that with a<br />
water catalyst but is still capable of providing a reasonable etch<br />
rate to nitride films. It was also seen that there were times<br />
when measurements from a 2 minute etch were higher than that<br />
of the 3 minute etch. This provided a negative selectivity as can<br />
be seen in Fig. 7 and Fig. 9. The error for the film etch<br />
measurements may also be in the error of the nanospec<br />
microscope used for measuring the films as the amounts being<br />
etched were found to be at the lower limits for accuracy that it<br />
can deal with.<br />
With respect to the etch pressures, in both the low and high<br />
temperature processes, it is difficult suggest which process<br />
provided the better selectivity as a whole (excluding of course<br />
the negative selectivities as already accounted for). In low<br />
pressure regimes, it was seen that as the carrier gas flow for the<br />
catalyst was increased, selectivity dropped off. This suggests<br />
that higher levels of ethanol in the chamber etch nitride films<br />
more readily than that of the oxide films which is counter to the<br />
focus of this study. In high pressure regimes at both low and<br />
high temperatures, the highest selectivities are found as<br />
expected when there is a higher carrier gas flow for the catalyst<br />
however the highest selectivity is found at 25!, this would<br />
suggest that the effects of the water vapour from the etch of the<br />
oxide film is not as influential as it is with a water based<br />
catalyst process. For the low temperature processes, the<br />
selectivities were higher at a high pressure as expected but only<br />
with a value of 14:1 versus the water catalyst process<br />
selectivity of 40:1. The low temperature process provided one<br />
of the highest selectivities of 15:1 but as stated earlier, it is hard<br />
<br />
39
!<br />
to be completely certain as to the accuracy of this value as the<br />
film measurements were of such a small value that it was in the<br />
lower accuracy regions of the microscope. The variation of<br />
carrier gas flows did not improve the selectivity as observed<br />
with the water catalyst process. It in fact caused a decrease in<br />
selectivity during low pressure regimes at both temperature<br />
ranges as stated earlier. At high pressure regimes, the increase<br />
behaved more as expected and did provide higher selectivities<br />
at their highest flow values.<br />
C. PECVD Oxide and PECVD Nitride with Ethanol Catalyst<br />
(Comparable Etch Rate Process Comparison):<br />
Comparing the etch rates and selectivities of a comparable<br />
etch against the water catalyst process provides a clearer<br />
picture than the direct comparison experiments. It can be seen<br />
that the highest selectivities are obtained interestingly at low<br />
pressures where a higher level of the etch by products are being<br />
pumped away and as such have a lower residence time for the<br />
etch gasses.<br />
Looking at the behaviour of the etch with respect to the<br />
pressures, it can be seen that while higher pressures cause a<br />
higher etch level, the overall values of etched oxide decrease as<br />
the etch nitride levels increase when the carrier gas flow is<br />
increased throughout all the tests carried out in this regime.<br />
This would show some agreement with the direct ethanol<br />
comparison results which suggest that as a higher level of<br />
ethanol is introduced into the process module, nitride films<br />
begin to etch at a faster rate than that of oxide films. High<br />
temperature processes provided inconsistent selectivity values<br />
where the low temperature processes did provide a more<br />
consistent level of selectivity. At 25!, the selectivities were all<br />
reasonably close in value which suggests that an ethanol<br />
catalyst either does not etch the oxide as quickly as a water<br />
catalyst or that the nitride films etch equally as fast regardless<br />
of the pressure such that when the pressure increases, the<br />
nitride etch increases at an equivalent rate as the oxide making<br />
it difficult to achieve an improvement in selectivity.<br />
A key point that should be discussed with regards to the<br />
comparable etch processes is that the pressures required to<br />
provide comparable etch rates were very high (as high as 35<br />
Torr at 25!), which would allow a very high residence time<br />
for the etch gasses within the process module. This in turn<br />
would allow the water vapour produced from the oxide film<br />
etching to reside in the chamber for a long period of time. One<br />
must ask the question of how much of the etching is in fact due<br />
to the ethanol catalyst and how much is from the water<br />
produced from the reaction in the process module.<br />
11-13 May 2011, Aix-en-Provence, France<br />
process achieved a selectivity of 10:1. This initial study is an<br />
investigation solely into the effects of the catalysts on vapour<br />
phase HF etching of sacrificial materials for MEMS<br />
manufacturing. A more detailed investigation of the etching<br />
method is being undertaken and the results will be published in<br />
the future.<br />
ACKNOWLEDGMENT<br />
The authors wish to acknowledge the Engineering and<br />
Physical Sciences Research Council (EPSRC) for their<br />
financial support through the Engineering Doctorate<br />
programme. They are also grateful to the Scottish<br />
Microelectronic Centre for their assistance in the fabrication of<br />
the samples for this study.<br />
REFERENCES<br />
[1] P.J. Holmes and J.E. Snell, “A vapour etching technique for the<br />
photolithography of silicon dioxide,” Microelectronics Reliability, vol. 5,<br />
issue 4, Pages 337-341, November 1966.<br />
[2] Gregory Van Barel, Luc Mertens, Ward De Ceuninck and Ann Witvrouw.<br />
“Apparent and steady-state etch rates in thin film etching and underetching<br />
of microstructures: I. Modelling,” Journal of Micromechanics and<br />
Microengineering, Volume 20, Number 5, 2010.<br />
[3] Gregory Van Barel1,2,4, Bert Du Bois1, Rita Van Hoof1, Jef De Wachter3,<br />
Ward De Ceuninck2 and Ann Witvrouw1, “Apparent and steady-state etch<br />
rates in thin film etching and under-etching of microstructures: II.<br />
Characterization,” Journal of Micromechanics and Microengineering,<br />
Volume 20, Number 5, 2010.<br />
[4] B. Du Bois, G. Vereecke, A. Witvrouw, P. De Moor, C. Van Hoof, A. De<br />
Caussemaeker, A. Verbist, “A comparison between wet HF etching and<br />
vapor HF etching for sacrificial oxide removal,” Proc. SPIE vol. 4174,<br />
pages 130-141, Micromachining and Microfabrication Process Technology<br />
VI, September 2000.<br />
[5] Chang-Jin Kim, John Y. Kim, Balaji Sridharan, “Comparative evaluation<br />
of drying techniques for surface micromachining,” Sensors and Actuators<br />
A 64, pp. 17-26, 1998.<br />
[6] E. Forsén, Z.J Davis, M. Dong, S.G. Nilsson, L. Montelius and A. Boisen,<br />
“Dry release of suspended nanostructures,” Microelectronic Engineering<br />
73-74, pp. 487-490, 2004<br />
[7] J. Anguita, F. Briones, “HF/H2O vapor etching of SiO2 sacrificial layers<br />
for large-area surface-micromachined membranes,” Sensors and Actuators<br />
A 64, pp./ 247-251, 1998.<br />
VI.<br />
CONCLUSIONS & FUTURE WORK<br />
A comparative study of water and ethanol as catalysts for an<br />
anhydrous Hydrogen Fluoride vapour phase etching of silicon<br />
dioxide and silicon nitride has been described. The results<br />
show that with the regimes defined in this work that the best<br />
selectivity can be achieved using water as the etch catalyst at<br />
low temperature and high pressure conditions. A selectivity of<br />
40:1 has been obtained. Under comparable etching conditions<br />
the best selectivity for ethanol as a process catalyst is 15:1.<br />
While comparable etch rates to that of the water catalyst<br />
<br />
40
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Agile MEMS packaging<br />
for mass customized MEMS products<br />
Jens G. Kaufmann, David Flynn, Keith Brown, Marc P.Y. Desmulliez<br />
Heriot-Watt University<br />
School of Engineering and Physical Sciences<br />
Edinburgh, EH14 4AS, Scotland, UK<br />
jens@jens-kaufmann.net<br />
Abstract - The market for micro electromechanical systems<br />
(MEMS) is becoming increasingly competitive especially for<br />
small and medium sized enterprises. The progressing trend in<br />
mass manufacturing leads to very few competitors that dominate<br />
the MEMS market. This competitive environment results in<br />
cheap devices with limited variation, therefore opportunities at<br />
the tail end of the market are usually neglected. The MEMS<br />
research community predominantly focuses on new device types<br />
and manufacturing methods with a high impact factor and those<br />
are naturally settled in the mass market segments.<br />
This paper introduces an approach for research and<br />
development that aims to open the opportunities within the long<br />
tail of the market. It consists mostly of a combination of new<br />
design approaches for MEMS and packaging, software tools and<br />
digital manufacturing technologies to ensure that solutions<br />
developed are ones that can be adapted to different customer<br />
requirement on a level that is perhaps uncommon in the MEMS<br />
community.<br />
I. INTRODUCTION<br />
The high investment costs involved with producing MEMS<br />
have in the past proven to put an emphasis on devices and<br />
products for the mass market [1]. Even products that can have<br />
revolutionary properties for some applications can take<br />
decades to arrive in the commercial markets. This is often due<br />
to a limit sized market and high development and<br />
manufacturing cost of MEMS and therefore often rejected<br />
until the market potential becomes large enough to justify the<br />
common mass manufacturing approach [2].<br />
The obsolescence of older aircrafts as well as the enormous<br />
investments required to replace them, has stimulated an<br />
increasing demand for innovative service concepts in the<br />
avionics industry [3]. One of the key areas to keep an aircraft<br />
flying is the condition of the electrical systems. However<br />
faults in the wiring harnesses are highly difficult to detect [4].<br />
To overcome these issues, a new sensor approach was<br />
developed based on a distributed sensor network on board the<br />
aircraft itself. [5] To make this technology available to the<br />
ageing aircraft market it was necessary to incorporate the<br />
specific requirements of the market. The large variety of<br />
geometry, material combination and signals employed for<br />
aircraft wiring over the last 70 years, not only between models<br />
of aircraft but also amogst planes of the same model, makes it<br />
impossible to develop a one-fits-all solution hence mass<br />
manufacturing of the sensor product is not a viable<br />
proposition.<br />
II. MASS CUSTOMISATION OF MEMS PACKAGING<br />
The ability to manufacture responsively is one of the principal<br />
requirements for commercial MEMS products. The<br />
automotive industry was the first to successfully optimize the<br />
production of their MEMS devices, followed by the consumer<br />
electronics and telecommunications industry. To achieve these<br />
commercial volumes, the types of devices were limited.<br />
New forms of MEMS devices have problems presenting an<br />
attractive business opportunity, if they have a limited range of<br />
applications. The resulting niche of potential customers is<br />
often too small to justify large-scale productions that come<br />
with the current methods of MEMS manufacturing. [1]<br />
Even though this seems to be a problem that is difficult to<br />
overcome, packaging as a support technology allows, in many<br />
cases, the properties of a MEMS device to be extended into a<br />
previously inaccessible application.<br />
As with MEMS themselves the packaging technology is<br />
mainly derived from the mass manufacturing methods from<br />
electronics manufacturing and consequently suffers from<br />
similar constraints. As a result of this, change has a high<br />
impact on the entire product lifecycle and neither the<br />
conventional design, manufacturing nor test tools have the<br />
capacity to deal with an agile design and cost competitive<br />
product requirement.<br />
In this paper the Health and Usage Monitoring MicroSystems<br />
(HUMMS) that are shown below demonstrate the vairous<br />
strategies that were employed in the different stages of its<br />
product life cycle to avoid impact of change originated from<br />
the customer and allow selling of MEMS into the long tail<br />
market of maintenance of ageing aircrafts. [5]<br />
III. DESIGN CONCEPTS<br />
Overview<br />
The way the product was conceptualized is based on four<br />
different concepts: Axiomatic design (AD), Generative design<br />
(GD), Rapid manufacturing (RM), which allow for an<br />
integrated design and manufacturing approach, and Automated<br />
Testing. AD is the overarching design framework, GD is what<br />
41
models the framework and the associated (rapid)<br />
manufacturing in software. The testing enables the quality<br />
management of the system and the model that the designer has<br />
of the processes involved.<br />
Axiomatic Design<br />
An Axiom driven Design process allows the business to<br />
establish interactions between user requirements, design<br />
solutions and manufacturing processes. Impacts of change can<br />
be retraced top down as well as bottom up [6]. As can be seen<br />
in Figure 1, Axiomatic Design divides the problem space into<br />
four domains:<br />
• Requirement domain<br />
• Functional domain<br />
• Physical domain<br />
• Process domain<br />
Within each domain there are independent nodes that are<br />
always caused by another node that is not on the same level.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
established by the Axiomatic Design. Using this concept, user<br />
requirements are no longer static but dynamic with the<br />
changing variables. The capabilities of generative design also<br />
called explicit history design are closer to a programming<br />
environment rather then to a classical Computer Aided Design<br />
(CAD). Parametric CAD today already allows for change<br />
management but not of the magnitude required. Being<br />
optimized for the workflow of a certain user, product<br />
designers, architects, mechanical or electrical engineers,<br />
commercial Computer Aided Design packages are to limited<br />
for all tasks required [7].<br />
Generative Design, as employed by many modern architects,<br />
requires much more work to create design files but allows for<br />
quasi real-time interaction with input parameters, even over<br />
the World Wide Web. Some samples can be seen in Figure 2.<br />
Figure 2: Computationally generated current sensors.<br />
The tool chosen to implement the GD approach was the<br />
grasshopper3D plug-in within Rhino3D, version 4. This was<br />
later linked to a web service that allows for data entering at the<br />
customer's side.<br />
Figure 1: Domains of Axiomatic Design<br />
For example, a customer wants to monitor a certain wire. That<br />
customer requirement, R1, spawns some functional<br />
requirements F1 and F2. Let’s assume F1 is: “the product must<br />
be mounted to the wire”. And F2 is “product must be sensing<br />
the wire performance”. To do so F2 spans Pa3 in the physical<br />
domain, which is a “current sensor”. The Pa3 can’t function<br />
without some support. So Pa3 causes a new functional<br />
requirement, F3, support infrastructure and so on.<br />
The way a design problem is seen in AD, as a mathematical<br />
graph with no circular dependency, makes it a perfect match<br />
for Generative Design (GD) techniques. The trick is to be able<br />
to quantify the relationships between the nodes and to make<br />
sure that no dependencies are circular. In this case study, the<br />
design was complex and derived using data gathering form<br />
potential customers in the aircraft industry.<br />
Rapid Manufacturing<br />
Given the mesoscale of the packaging intended for the<br />
applications there is a large variety of rapid manufacturing,<br />
also called Digital Manufacturing processes, that can deliver<br />
the required properties inclusive of a versatile fabrication<br />
attributes. In the case study presented, classical 3D printing<br />
and fused deposition modeling were employed to manufacture<br />
the structural 3D parts of the system. The difference in build<br />
quality can be seen in Figure number 3.<br />
Generative Design<br />
The Generative Design approach was chosen to build a suite<br />
of dedicated design tools that implement the requirements<br />
42
11-13 May 2011, Aix-en-Provence, France<br />
<br />
! "<br />
#<br />
%<br />
$<br />
&<br />
Figure 4: System concept of the onboard units:<br />
! Aircraft wire<br />
" Current sensor<br />
# 3D core part<br />
$ Flex circuit<br />
% Batteries<br />
& Main controller and<br />
environmental sensor board<br />
Figure 3: Fused Deposition Modeling (left), solvent based 3D<br />
printing (right)<br />
A CNC CO 2 laser cutter was also used to pattern the flex<br />
circuits that form the enclosure, the electrical connections and<br />
mount the system onto the wire that needs to be monitored.<br />
Automated testing<br />
The testing is implemented on two levels, the component level<br />
which is done mostly by the supplier of the individual parts as<br />
well as an automated 3D machine for the 3D parts. The second<br />
level is the functional test of the entire system. This allows, in<br />
conjunction with the axiomatic design model, for testing of the<br />
entire unit, not just the parts but also the processes and the<br />
validity of the design model itself.<br />
IV. IMPLEMENTATION<br />
The HUMMS design<br />
The product is designed for retrofitting on existing wiring on<br />
board of an aircraft. The system had to be mounted to the wire<br />
to enable the user to gather in flight health data. An intrusive<br />
way like splicing into the wire as well as exchanging part like<br />
the connector was not an option for the customers questioned.<br />
The initial design idea was a multi-sensor tag that would<br />
monitor the environmental conditions as well as monitoring<br />
the signal of the wire, the latter through the use of a micro coil<br />
as shown in Figure 4.<br />
After several design iterations, the prototypes shown below in<br />
Figure 6, which is a combination of RM part and self selfadhesive<br />
flex PCB, were chosen and implemented in<br />
grasshopper.<br />
The grasshopper3D implementation<br />
The implementation of the design idea is done in 3 layers<br />
depending on their interdependencies. Primary parts are parts<br />
that can be directly derived from customer requirements and<br />
have prerequisites that are influenced by other parts and of the<br />
shelf components. Secondary parts take the information from<br />
the requirements and then design components as shown in<br />
Figure 5.. It controls the configuration files for the host and<br />
embedded software. The final layer is the generation of the<br />
manufacturing files.<br />
Figure 5: Visualization of the real dimensions in the desired structure<br />
of a low fidelity print.<br />
3D freeform modeling<br />
Even though the 3D Systems rapid prototyping system is not<br />
able to produce aircraft certified parts it is sufficient to check<br />
43
11-13 May 2011, Aix-en Provence, France<br />
<br />
the validity of the process in a prototyping stage. But our<br />
industrial partners have made certifiable parts on machines<br />
from EOS in Germany and Fused Deposition Modeling<br />
machines form Stratasys.<br />
The 3D printing approach chosen generates a 3D structure by<br />
layering a mineral powder in a container in the z-axis<br />
moveable platform. It fixes the powder in place by ink jetting<br />
a binder on the surfaces of the desired cross section and so<br />
rendering it solid. The repetition of the process after a few<br />
layers of powder is applied to generate the 3D part, as show in<br />
Figure 6.<br />
Figure 7: assembled Humms onboard unite<br />
Figure 6: 3D part variations based on the different current sensors.<br />
After printing, the part needs to be freed from the excessive<br />
powder and fixed with one of several agents that can vary<br />
based on the desired properties of the part. In the parts<br />
demonstrated cyanoacrylate was used which gives the part a<br />
very strong structural integrity but is not as heavy as a part<br />
infiltrated with an epoxy resin.<br />
Laser structuring of flexible printed circuit boards.<br />
To form the flex PCB a copper coated (9!m) polyimide film is<br />
coated with anacrylic black paint. This paint is removed with<br />
the CO 2 laser and then etched. After the etch, the rest of the<br />
resist is removed in an acetone bath. Then the adhesive film is<br />
patterned and applied to the flex PCB. The advantage of this<br />
method, compared to electroforming for example, is again the<br />
flexibility and speed of the process. By using a spray on resist<br />
the shape of the circuit is unimportant and by using the laser<br />
we can shape a circuit that is within the capabilities of the<br />
machine.<br />
Assembled system<br />
All components are assembled by hand as pictured in Figure 7<br />
and, with the protective film of the PCB completely removed,<br />
the system can be mounted to the wire.<br />
V. RESULTS<br />
The work conducted demonstrates that ”batch-of-one” MEMS<br />
applications can be viable from a business perspective if<br />
modern design methodologies are combined with rapid<br />
manufacturing within the MEMS domain. The work has also<br />
shown that 24h iteration from customer requirements to a<br />
manufactured and tested product is possible.<br />
The test of the system showed (Figure 7) that it the system<br />
could deliver an even better result then a macroscale current<br />
pickup from Pearson’s reference coil in the range from 1-<br />
10MHZ.<br />
Figure 7: Oscilloscope display when sending a 2 MHz sine wave<br />
along the WUT. Blue trace is the micro Rogowski sensor; orange<br />
trace is the Pearson reference coil.<br />
The systems environmental sensors on the main rigid circuit<br />
board also allow for temperature (-200º-200ºC) acceleration (-<br />
5g – +5g) in x, y and z and humidity pickup (10% - 99.8%<br />
rH). The employed sensors transmitting their data over a wired<br />
connection to a host system that was compiled for MacOS,<br />
Windows and Linux systems and can forward the data to any<br />
remote iOS device as show in Figure 8.<br />
44
11-13 May 2011, Aix-en-Provence, France<br />
<br />
ACKNOWLEDGMENT<br />
The development team also would like to acknowledge the<br />
support and information made available by Ultra Electronics-<br />
Electrics and the Ministry of Defense.<br />
Without the continuous feedback from the commercial as well<br />
as the technical divisions within Ultra and the service men and<br />
women from the MoD that grounded the project in reality, the<br />
project would not have succeeded.<br />
Figure 8: Remote environmental monitoring interface host software<br />
and the iPhone and iPad clients.<br />
VI.<br />
FUTURE WORK<br />
The system will use several service providers to manufacture<br />
the parts with different RM processes and materials to analyze<br />
the impact of those on the capabilities of the united and so<br />
refine the model and widen its application to new products.<br />
Furthermore other potential application areas will be explored.<br />
Together with a small team the author is in the process of<br />
forming EnvironMEMS, a design house that delivers rapidly<br />
developed environmental monitoring based on the technology<br />
above.<br />
REFERENCES<br />
[1] Senturia, “Perspectives on MEMS Past and Future:<br />
the Torturous Pathway from the Bright Ideas to Real<br />
Products.”, Digest Tech - Papers Transducers ’03<br />
Conference, 2003<br />
[2] Anderson, “The Long Tail”, Wired US, October 2004<br />
[3] Furse and Haupt – “Down to the wire [aircraft<br />
wiring].”, IEEE Spectrum (2001) vol. 38 (2) pp. 34-<br />
39<br />
[4] Sullivan and Slenski “Managing Electrical<br />
Connection Systems and Wire Integrity on Legacy<br />
Aerospace Vehicles” Proceedings of the FAA<br />
Principal Inspectors and Engineers Workshop, 2001<br />
[5] Moffat et al – “MEMS Sensors and High Frequency<br />
Test Techniques for Prognostic Health Management<br />
of Aircraft Wiring.” <strong>Online</strong>-http://www.patentdfmm.org/site/Restricted/CDROM2005/WP7/D7.9B<br />
CF_Report_1.pdf , Workshop 2005<br />
[6] Kim. “AXIOMATIC DESIGN OF MULTI-SCALE<br />
SYSTEMS.” Proceedings of ICAD2004 The Third<br />
International Conference on Axiomatic Design<br />
(2004) pp. 1-5<br />
[7] McCormack et al. “Generative design: a paradigm for<br />
design research.” Proceedings of Future ground,<br />
Design Research Society, Melbourne (2004)<br />
45
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Wafer-Level Glass-Caps for Advanced Optical Applications<br />
Juergen Leib, Oliver Gyenge, Ulli Hansen, Simon Maus, Karin Hauck * , Kai Zoschke * , Michael Toepper *<br />
MSG Lithoglas AG<br />
Gustav-Meyer-Allee 25, 13355 Berlin, Germany<br />
juergen.leib@lithoglas.de<br />
* Fraunhofer Institute for Reliability and Microintegration, Berlin<br />
Gustav-Meyer-Allee 25, 13355 Berlin, Germany<br />
Abstract<br />
A novel process flow to manufacture miniaturized optical<br />
windows on wafer-level is presented. Those windows can be<br />
used for miniaturized optical products like high-brightness<br />
LEDs (HB-LED) and digital projection (DLP) as well as more<br />
complex optical data-communication, since integrated optical<br />
functions can be implemented with low tolerances.<br />
We explain the fabrication of cap-wafers having a shallow<br />
cavity with a depth of typically 10 µm used in photo sensors<br />
and a unique manufacturing process for cap-wafers with a<br />
deep cavity of e.g. 300 µm used in LED packaging.<br />
Those cap-wafers are used in wafer-level integration of<br />
advanced, miniaturized optical products. We discuss two<br />
options for wafer bonding i.e. bonding using adhesive as well<br />
as anodic bonding.<br />
As an example on product level a miniaturized photo<br />
sensor package, a pressure sensor package as well as a LED<br />
package is discussed.<br />
Wafer-Level Capping of Optical and MEMS Devices<br />
Wafer level packaging of optical devices is becoming<br />
more and more mainstream [1]. Beside the overall deciding<br />
advantage of costs per die and system yield, performance and<br />
reliability targets can be matched today for a wide variety of<br />
applications.<br />
Especially for optical devices it is very important to seal<br />
the (particle) sensitive areas of the chip as early as possible in<br />
the assembly process in order to avoid yield loss due to<br />
particle contamination. This issue is well-known from image<br />
sensor modules [2] and was one of the main drivers for the<br />
early adoption of wafer-level-packaging for camera module<br />
packaging. In the effort to seal off the optical devices as early<br />
as possible in the packaging process, we have shown that the<br />
overall assembly yield can be increased dramatically by<br />
introducing a wafer-level-capping process prior to<br />
conventional Chip-On-Board (COB) packaging and at the<br />
same time reducing the system costs.<br />
Depending on the method of micro structuring of the cap<br />
wafer, the individual caps may provide a cavity for the<br />
encapsulated devices. These cavities being obvious and wellknown<br />
for MEMS are also required for optical applications<br />
like MOEMS or image sensors – e.g. if these have micro<br />
lenses on the optical area of a camera chip. However, for<br />
optical chips the cavity, the glass cap and their relevant<br />
surfaces in particular, are contributing to the overall optical<br />
performance of the device. Therefore their tolerances and<br />
quality have to meet stringent optical requirements.<br />
As an example the wafer level package of miniaturized<br />
photodiodes used for high density optical storage (Fig. 1) is<br />
discussed, demanding special attention on advanced optical<br />
performance, UV stability and high reliability. Since intensive<br />
blue laser light (405 nm) is used, a glass window package is<br />
considered to perfectly meet the needs for long term stability<br />
and performance.<br />
Fig. 1a) Miniaturized glass cavity windows on product wafer<br />
Fig. 1b) Typical glass cavity window after dicing bonded<br />
to a device wafer. A 10 µm high Lithoglas glass rim on the<br />
cover glass acts as the bond frame with a total width of 100<br />
µm forming an optical cavity over the functional area of the<br />
device.<br />
Having the devices protected very early in the assembly<br />
process, conventional COB assembly – including wirebonding<br />
and molding – can be performed on standard<br />
equipment in a cost competitive environment with high<br />
component yields being achieved. Beside significantly lower<br />
overall costs, the wafer level capping process does offer low<br />
profile and small dimensions of the final package as well as<br />
compliance with standard wafer design rules. This way it is<br />
outperforming other packaging approaches [3].<br />
Novel Fabrication Process for Optical Cap-Wafers<br />
46
The fabrication of optical Cap-Wafers, i.e. the fabrication of<br />
cavity windows with high optical performance (Fig. 1b) on<br />
wafer-level processing, poses an extra challenge for<br />
manufacturing.<br />
Conventional methods forming a cavity are commonly using<br />
subtractive processes such as plasma etching, wet etching or<br />
even more coarse processes like sandblasting and ultrasound<br />
milling. These processes remove material from a wafer<br />
substrate (e.g. a polished glass wafer) in unmasked areas thus<br />
forming a cavity; whereas the material which is protected by a<br />
masking technology remains and forms bond frames or<br />
similar structures. It should be noted, that processing takes<br />
place in the bottom of the cavity,<br />
In case of an optical cavity this surface represents the optical<br />
active surface and is, especially with shallow cavities, very<br />
close to the focal plane of an optical sensor. Furthermore any<br />
optical surface finish in the cavity area such as highly<br />
polished surfaces, antireflective / filter coatings or optical<br />
elements like gratings and apertures are damaged by the<br />
cavity formation using subtractive processes.<br />
These drawbacks can be avoided by using additive techniques.<br />
In this case bond frames or similar structures are formed<br />
on top of a high quality optical wafer by depositing material<br />
whereas the optical area in the bottom of the cavity shall<br />
remain intact.<br />
Most commonly photoactive polymers are used to form the<br />
frames like BCB or SU-8. These materials are typically<br />
applied by spin- or spray-on, structured using lithography<br />
where the photosensitive material is selectively exposed and<br />
removed in a development step thus forming the bond frames<br />
very precisely. This method is the mainstream solution for<br />
low-end product, where the disadvantages of polymers like<br />
limited reliability, moisture uptake, oxygen or moisture<br />
transmission or degradation at elevated temperatures play a<br />
minor role.<br />
In the effort to meet reliability requirements for advanced<br />
products stencil printing of frit-glass or solder glasses is used.<br />
However these glass-powder based materials generate<br />
particles during processing and the resolution of the lateral<br />
dimensions as well as the height of the frames structures is<br />
limited by the minimal grain size of the glass-powder, the<br />
stencil printing process and the need for a reflow process for<br />
post-processing the glass-slurries.<br />
In order to overcome the limitations of todays mainstream<br />
solutions, we propose a novel method to fabricate precise and<br />
hermetic Cap-Wafers with high optical quality:<br />
Shallow Cavities using Additive Microstructuring of Glass<br />
– Lithoglas process<br />
An advanced, additive microstructuring process of glass<br />
(Lithoglas process) is used to fabricate a “glass-only” cap<br />
wafer providing utmost optical performance at a very low<br />
level of defects. The novel deposition and microstructuring of<br />
glass allows the formation of thin films of dense borosilicate<br />
glass on a broad range of substrates at substrate temperatures<br />
below 100 °C [3, 4].<br />
The deposition of the glass is done by a plasma-assisted e-<br />
beam evaporation process. It is a high rate deposition process<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
with deposition rates of several hundred nm/min and at the<br />
same time low substrates temperature. With the high<br />
deposition rate typical film thicknesses of 3 – 20 µm can be<br />
achieved easily and production can be run as a cost effective<br />
batch process. Thicker layers as thick as 100 µm have been<br />
done in R&D – this is only possible due to the outstanding<br />
control of stress in the deposited layers.<br />
The glass layer can be microstructured by lift-off (Fig. 2a).<br />
Since the deposition process is working at low temperatures<br />
standard photo resists can be used for masking.<br />
Process Step 1 – Lithography: As a first step photo resists<br />
is deposited by spin-on. It is exposed by mask aligner or<br />
stepper and developed. The photo resist carries the negative<br />
image of the target structures in the glass layer.<br />
Process Step 2 – Glass Deposition: The borosilicate glass<br />
layer is deposited by plasma-assisted e- beam deposition on<br />
the full wafer. The substrate temperature stays below 100 °C<br />
during this process.<br />
Process Step 3 – Lift-Off: The glass microstructures are<br />
developed by dissolving the photo resist mask and with this<br />
removing the glass on top of it. A structured glass layer<br />
remains at the locations which were not covered by the resist;<br />
areas covered by photo resist were protected trough out the<br />
process and reveal their original surface finish after lift-off.<br />
Depending on the layer thickness an aspect ratio of 1.8 can<br />
be achieved yielding 1.3 µm fine structures in a 3 µm glass<br />
layer by using a mask aligner for lithography (Fig 2b).<br />
Glass, especially borosilicate glass, is well known for its<br />
chemical inert behavior and stability. It is temperature stable<br />
and hardly dissolves in most acids, bases and solvents. It is<br />
the close-to-perfect material for semiconductor packaging in<br />
respect to its electrical, chemical and physical properties. The<br />
use of Lithoglas microstructured thin-film glass as passivation<br />
and functional layer and its unique advantages is described<br />
elsewhere [5]. In this paper we focus on use of Lithoglas as<br />
bond frame.<br />
Fig. 2.a) Lithoglas Process Flow – Microstructuring of<br />
Glass by Lift-Off Borosilicate glass (blue) can be deposited at<br />
low temperatures of below 100°C on a variety of substrates<br />
(grey) by using plasma-assisted e-beam deposition. This<br />
permits the microstructuring of the glass thin-film by lift-off<br />
using standard photo resists (red)<br />
47
11-13<br />
<br />
May 2011, Aix-en-Provence, France<br />
<br />
glass frame<br />
adhesive<br />
Fig. 2.b) Microstructured Glass on Silicon with smallest<br />
Feature being 1.3 µm with an Aspect Ratio of 1.8. The<br />
deposited glass is structured by photo resist lift-off.<br />
It should be noted, that also larger areas, as frequently<br />
used as optical cavity windows for image sensors and optical<br />
MEMS, can be lifted without residues. Throughout the whole<br />
process sensitive areas where no glass shall be deposited are<br />
covered and protected by the lift-off photo resist. As<br />
mentioned earlier, this is especially useful for MOEMS and<br />
optical sensors, when e.g. microstructured glass bond frames<br />
are formed on a glass cap wafer using an antireflective or<br />
other coating in order enhance the optical performance of the<br />
final device.<br />
At the same time very low variations on critical<br />
dimensions like width and height of the bond frame on a<br />
wafer and wafer-to-wafer are achieved. The use of the<br />
microstructured glass as bond frames provides a solid and<br />
dense embodiment of a cavity. The rigid frame guaranties a<br />
well-defined height of the cavity throughout processing as<br />
well as in the final product and can be controlled at tolerances<br />
less than 1 µm in mass production. This is especially<br />
important due to the critical optical design required for the<br />
advanced optical application.<br />
Quasi-Hermetic Wafer-Level-Integration of Shallow<br />
Cavities using ultra-thin Adhesive Bonding<br />
Furthermore, the dense frame acts as an efficient diffusion<br />
barrier for humidity and has – being a bulk glass frame – an<br />
extremely small moisture uptake. In combination with the thin<br />
bond line it provides a quasi-hermetic cavity. In the specific<br />
product described (Fig. 3) a cavity with a height of 10.8 µm is<br />
provided by using a glass frame with a height of 10.3 µm and<br />
the thin adhesive layer being typically in the range of 0.5 µm<br />
– depending on the topography of the device wafer. The final<br />
products passes JEDEC MSL Level 1.<br />
As of today the average wafer capping yield is in the high<br />
nineties, whereas on champion wafers 100 % yield can be<br />
achieved. Please refer to Ref. [3] for more details on<br />
reliability data on production level.<br />
´<br />
Fig. 3.a) Typical glass cavity window after dicing bonded<br />
to a dummy wafer. An 80 µm glass rim acts as the bond frame<br />
with a total width of 100 µm. In order to achieve these tight<br />
design rules the amount of the bond adhesive as well as its<br />
bleeding during the bonding process must be well controlled<br />
Fig. 3.b) Cross-Cut through the bond interface revealing<br />
the thin adhesive bond line of the µCapping process and its<br />
tide control of glue bleeding.<br />
Hermetic Wafer-Level-Integration of Shallow Cavities<br />
using Anodic Bonding<br />
Due to the nature of the additive deposited Lithoglas being<br />
a borosilicate glass, the bond frames can be used directly as<br />
bond interface, when using an anodic bond process.<br />
Anodic bonding is used for a wide range of applications,<br />
especially in MEMS industry, where reliable, hermetic sealing<br />
of silicon devices is required. Anodic bonded devices are well<br />
known for their high mechanical and chemical stability [4].<br />
For mainstream applications a borosilicate glass substrate,<br />
such as Pyrex 7740, Schott 8330 (also called Tempax) or<br />
Borofloat is bonded to substrates – today predominately<br />
silicon - by applying an external voltage at elevated<br />
temperatures. This process was first reported as field assisted<br />
bonding in 1969 by Wallis et.al [5]. The sealing of two silicon<br />
surfaces by deposition of a glass thin-film was first reported in<br />
1972 by Brooks et.al. who produced piezo-resistive pressure<br />
sensors [6].<br />
The use of a glass thin-film is advantageous compared to<br />
using a glass wafer especially when two silicon wafers need to<br />
be bonded to form a wafer stack or a package. Bulk glass<br />
wafers with a thickness of several hundred micrometres do<br />
not only limit the miniaturization of the final devices – unless<br />
48
the glass is polished down in a costly process after being<br />
bonded to the first substrate – they also require higher external<br />
bond voltages to drive the alkali ion migration in the glass<br />
during the bond process [7].<br />
However, using sputtering as deposition method for glass<br />
thin-films has a number of disadvantages as well, such as low<br />
deposition rates, high substrate temperatures, high intrinsic<br />
film stress and the tendency to form pin-holes. On a<br />
commercial basis it is very costly to form hermetic glass thinfilms<br />
with a thickness of several microns by using sputtering.<br />
Multiple efforts have been undertaken to establish more<br />
cost-effective deposition methods for borosilicate glass [8,9]<br />
such as e-beam deposition, which did not find their way into<br />
mainstream applications in wafer-level packaging yet. This is<br />
the mission of Lithoglas applying the unique plasma-assisted<br />
e-beam evaporation process.<br />
Anodic Bonding of Lithoglas on Borofloat to Silicon<br />
As an example we first demonstrate the anodic bond of<br />
Lithoglas bond frames with a height of 3 µm deposited on a<br />
6” Borofloat 33 wafer. The frame design is similar to that<br />
used for adhesive bond (Fig. 4).<br />
<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Borofloat 33 substrate are very similar to those of bulk<br />
Borofloat. A typical bond process is shown in Fig 5.<br />
Fig. 5) Record of anodic bonding parameters of 3 µm high<br />
Lithoglas frames on 500 µm Borofloat 33 to a silicon wafer<br />
are very similar to bulk Borofloat. As a standard the bond<br />
voltage is increased in 3 steps (400 V, 600 V and 800 V). The<br />
bond current show the characteristic spikes when stepping the<br />
voltage. A total charge of 432 mC is transferred during<br />
bonding of a net bond area of about 1100 mm². Bonding<br />
temperature is 390 °C.<br />
Fig. 4.a) Lithoglas bond frames with a height of 3 µm<br />
form a cavity on top of a 500 µm thick Borofloat 33 wafer.<br />
The Lithoglas rim acts as the bond frame with a total width of<br />
100 µm. The Lithoglas Cap-Wafer is anodic bonded to a<br />
silicon substrate.<br />
Fig. 4.b) Anodic bonding of Lithoglas frames to a silicon<br />
wafer achieves good quality bonds with high yields – visual<br />
inspection from the glass side.<br />
It shall be noted, that the process parameters for anodic<br />
bonding of 3 µm high Lithoglas frames on 500 µm thick<br />
Anodic Bonding of Silicon to Silicon using Lithoglas<br />
As mentioned earlier the use of a Lithoglas interface is<br />
most beneficial when two (or more) silicon wafers shall be<br />
hermetically sealed by anodic bonding. We have<br />
demonstrated that a 3 µm thick Lithoglas layer is sufficient to<br />
anodic bond two silicon wafers without any special pre- or<br />
post-treatment of the silicon wafers, however the bond<br />
voltages was adapted:<br />
For anodic bonding of bulk glass to silicon voltages of<br />
several hundred volts are used to drive the sodium ions in<br />
bulk borosilicate glass at elevated temperatures. Typical bond<br />
conditions are in the range of 500 to 1000 V as bond voltage<br />
at a temperature 300 to 400 °C to anodic bond a 500 µm thick<br />
Pyrex glass to silicon. Reducing the glass thickness to several<br />
microns it is obvious, that the bond voltage can be reduced<br />
proportionally yielding a similar electrical field for the fieldassisted<br />
anodic bond. In case the borosilicate film has a high<br />
dielectric breakdown voltage the electrical field can be<br />
increased significantly allowing anodic bonding at lower<br />
temperatures and at the same time lower external bond<br />
voltages.<br />
The borosilicate thin-films processed by plasma assisted e-<br />
beam evaporation exhibit a typical specific breakdown<br />
voltage of better than 240 V/µm at room temperature. This<br />
high value allows for relatively high electrical fields for<br />
anodic bonding allowing both bonding at lower temperatures<br />
as well as faster bonding processes. Compared to standard<br />
processes for bonding bulk glass – where typical electrical<br />
fields of 1 – 2 V/µm are applied – it was shown that using a<br />
Lithoglas layer a field strength of 20 V/µm are achieved<br />
applying e.g. 60 V as maximum bond voltage over an 3 µm<br />
borosilicate thin-film.<br />
A comparison of anodic bonding parameters of 3 µm high<br />
Lithoglas frames on 500 µm Borofloat 33 (#1) with 3 µm high<br />
Lithoglas on silicon (#3) while bonding to a silicon wafer is<br />
49
given in figure 6. The bond voltage is reduced from 400 V<br />
(#1) to 40 V (#3) while the bonding temperature was kept at<br />
390 °C for both cases. The bond current and the transferred<br />
charge for anodic bond of the borosilicate thin-film show the<br />
expected characteristic behavior, though the bond voltage is<br />
reduced to 40 V. Due to the small thickness of the bondinterface<br />
an increased electrical current of can be observed<br />
through the borosilicate thin-film in the steady-state. However<br />
the sheet-resistance is as high as 3 GΩ/□ at 390 °C.<br />
Fig. 6) Comparison of anodic bonding parameters of 3 µm<br />
high Lithoglas frames on 500 µm Borofloat 33 (#1) with 3 µm<br />
high Lithoglas on silicon (#3) while bonding to a silicon<br />
wafer. The bond voltage is reduced from 400 V (#1) to as low<br />
as 40 V (#3). Bonding temperature is 390 °C.<br />
As mentioned above, the deposited borosilicate thin-films<br />
can be microstructured by photo-resist lift-off. This allows the<br />
fabrication of well-defined anodic bondable areas on devices<br />
wafers. This technique was used to seal a silicon pressure<br />
sensor as shown in figure 7.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
Deep Cavities using novel “Cavity-by-Grind” method<br />
The Lithoglas process is very cost effective for the<br />
formation of thin anodic bond interfaces or shallow cavities of<br />
some ten micrometres. For the formation of deep cavities of<br />
up to some hundred micrometres hybrid materials such as<br />
Silicon-Glass Cap-Wafers are commonly used.<br />
Principally there are two ways used for manufacturing<br />
today. Firstly, bonding of a pre-structured spacer substrate<br />
(e.g. a silicon wafer) having cavity holes to the cover substrate<br />
(e.g. a glass wafer). However this process is limited to thick<br />
spacers of some hundred micrometres in order to avoid<br />
breakage during handling and bond.<br />
The second method is to form the cavity in the spacer<br />
substrate after bonding it to the cover substrate. In this case<br />
typically a silicon wafer is bonded to glass substrate and then<br />
the silicon is structured by wet- or plasma-etching using the<br />
glass wafer as etch stop. Though it allows the formation of<br />
cavities with a large scope of sizes, depths and shapes, it<br />
requires processing on the optical surfaces as discussed<br />
earlier.<br />
In order to avoid the drawbacks of conventional<br />
processing of those hybrid cap wafers, we propose a new<br />
process flow for their manufacturing, which is very stable and<br />
high yielding and is especially suitable for optical applications<br />
(Fig. 8).<br />
Fig. 8) Novel “Cavity-by-Grind” process for the<br />
manufacturing of hybrid cavity wafers with high optical<br />
performance.<br />
Fig. 7) Silicon pressure sensor devices on wafer with a<br />
3 µm thick Lithoglas anodic bond-frame around the central<br />
structure. The deposited bond-frame hermetically seals the<br />
conduction leads on the sensor devices. The image was taken<br />
prior anodic bonding of the device wafer to a silicon cap<br />
wafer.<br />
Sealing of conduction leads by deposition of the<br />
borosilicate glass on top is an unique feature of the Lithoglas<br />
process, thus enabling cost-effective, hermetic feed-throughs.<br />
As a first process step blind indentations are formed into a<br />
thick spacer substrate. These indentations are slightly deeper<br />
than the final cavities. With this the spacer wafers allows for<br />
automated handling without any carrier systems even for<br />
shallow cavities of some ten microns. We use standard (100)-<br />
silicon wafers as spacer wafers and use wet-etching for the<br />
formation of the blind cavities.<br />
As a second step the thick spacer is bonded to a cover<br />
substrate thus forming the cavities. The bonding can be done<br />
by conventional bonding techniques like anodic, adhesive or<br />
eutectic bonding, but also direct bonding is feasible due to the<br />
high quality of the surfaces. For our standard product, we use<br />
anodic bonding to bond the silicon spacer to a borosilicate<br />
cover glass.<br />
50
11-13 May 2011, Aix-en-Provence, France<br />
<br />
As a last step the excessive material is removed from the <br />
backside of the spacer substrate thus opening the cavities and<br />
defining their final depth. We thin down the silicon to the<br />
final thickness by conventional back-grinding and with this<br />
open the cavities.<br />
Fig. 9) Example of packaging of opto-devices using<br />
hybrid cap wafers.<br />
Due to the high quality of cap wafers incl. their low warp<br />
and geometrical tolerances, they can be used of wafer-level<br />
integration of devices, but also for housing of pre-assembled<br />
opto-devices on PCB as shown in figure 9. Today this is still<br />
the most common configuration.<br />
Lithoglas µCap used in Chip-On-Board Packaging<br />
After the capping of the devices on wafer level, the<br />
individual chips can be assembled in conventional ways with<br />
only minor modifications to the standard processes e.g.<br />
yielding COB components (Fig. 10). Since the devices are<br />
pre-packaged, the subsequent assembly steps can be<br />
performed in a cost effective, high through-put setup<br />
significantly reducing the requirements for clean room<br />
facilities and in-line inspection efforts even for image sensors<br />
and other optical applications.<br />
Fig. 10) Wafer-Level Capped Chips used in standard<br />
Chip-on-Board Assembly Flow;<br />
The final components achieve superior performance and<br />
reliability, e.g.:<br />
• There is no adhesive layer in optical path, which might<br />
cause changes in optical characteristics due to degradation of<br />
polymer under intensive illumination or due to delamination<br />
of the polymer layer.<br />
• The optical window is precisely positioned with tight<br />
control on x, y, z as well as tilt and rotation. The tight<br />
tolerances are proven on production level and even apply to<br />
very small windows such as 750 x 750 µm size or smaller,<br />
which are difficult to handle otherwise.<br />
• The outstanding control on glue bleeding in case of<br />
adhesive bonding and the small minimal width of the bond<br />
frames (typ. 100 µm) allows for advanced, miniaturized<br />
design.<br />
• The bond interface of the Lithoglas µCap is robust and<br />
positioned in the inner of the package, being additionally<br />
sealed and protected by the COB molding material.<br />
The package shown in figure 11 is used for optical pick-up<br />
for a 405 nm application achieving high yields greater than<br />
95% for the wafer capping process and passing JEDEC Level<br />
1 as molded COB component.<br />
Fig. 11) Final COB Devices with optical cavity window.<br />
51
Conclusions<br />
Microstructured borosilicate thin-films are used to form<br />
bond frames and cavities on optical wafers. Direct anodic<br />
bonding of these structures has been demonstrated for glassto-silicon<br />
and silicon-to-silicon wafers. The use of the<br />
Lithoglas structures as bond layers opens up the opportunity<br />
to anodic bond two silicon device wafers using only one<br />
anodic bond step. Furthermore the bond layer can be easily<br />
structured by lift-off and can hermetically seal underlying<br />
electrical leads forming a hermetic feed-through. Due to the<br />
outstanding dielectric properties of the borosilicate thin-films<br />
advanced bond parameters can be used enabling a fast anodic<br />
bond process and with this lowering the costs of hermetic<br />
wafer-level integration of silicon devices.<br />
A novel “Cavity-by-Grind” process was introduced, which<br />
allows the cost-effective manufacturing of optical cap-wafers<br />
with deeper cavities (some ten to some hundred micrometres)<br />
with high yields and high optical performance.<br />
The use of capped device for conventional COBpackaging<br />
was discussed.<br />
Acknowledgments<br />
MSG Lithoglas AG wishes to thank Prof. Klaus-Dieter<br />
Lang, Oswin Ehrmann and their team of Fraunhofer Institute<br />
for Reliability and Microintegration in Berlin for their<br />
ongoing support.<br />
Parts of the work was part-funded by the European<br />
Regional Development Fund (ERDF) and the government of<br />
Berlin, Germany as well as by the Federal Ministry of<br />
Education and Research, Germany.<br />
<br />
11-13 May 2011, Aix-en-Provence, France<br />
References<br />
1. Toepper, M., P. Garrou, The Wafer Level Packaging<br />
Evolution”, - Semiconductor , - - International, , Reed Elsevier<br />
Inc, Oct. 2004, p. SP-13.<br />
2. Chowdry, A. “Camera Module Assembly and Test<br />
Challenges”, Semiconductor International, Reed Elsevier<br />
Inc, Feb. 2006, p. SP- 4.<br />
3. Hansen, U., S. Maus, J. Leib, M. Toepper, “Novel<br />
Hermetic and Low Cost Glass-Capping Technology for<br />
Wafer-Level-Packaging of Optical Devices”, Proceedings<br />
of Conference ESTC 2010, Berlin, Sept 2010, Paper 170<br />
4. M. Esashi. Encapsulated micromechanical sensors.<br />
Microsystem Technol., 1:2–9, 1994.<br />
5. G. Wallis and D. I. Pomerantz. Field assisted glass-metal<br />
sealing. J.Appl. Phys., 40(10):3946–3949, 1969.<br />
6. A. D. Brooks, R. P. Donovan, and C. A. Hardesty. Lowtemperature<br />
electrostatic silicon-to-silicon seals using<br />
sputtered borosilicate glass. J. Electrochem. Soc.,<br />
119(4):545–46, 1972.<br />
7. B. Schmidt, P. Nitzsche, S. Grigull, U. Kreissig, B.<br />
Thomas, K. Herzog, and K. Lange. In situ investigation of<br />
ion drift processes in glass during anodic bonding. Sensors<br />
and Act. A, 67(1-3):191–198, 1998.<br />
8. Anders Hanneborg, Martin Nese, and Per Øhlckers.<br />
Silicon-to-silicon anodic bonding with a borosilicate glass<br />
layer. J. Micromech. Microeng., 1:139–144, 1991.<br />
9. Experimental analysis on the anodic bonding with an<br />
evaporated glass layer; Woo-Beom Choi et al 1997 J.<br />
Micromech. Microeng. 7 316-322.<br />
10. Mund, D., J. Leib, M. Toepper, Novel Hermetic Wafer-<br />
Level-Packaging Technology Using Low-Temperature<br />
Passivation, Proceedings of 55th ECTC Conference,<br />
Orlando, June. 2005, pp. 562-565.<br />
11. Hansen, U., S. Maus, J. Leib, M. Toepper, Robust and<br />
Hermetic Borosilicate Glass Coatings by e-Beam<br />
Evaporation, Proceedings of the Eurosensors XXIII<br />
conference, Lausanne, September 2009.<br />
52
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Reduced Order Modelling of MEMS Dynamics<br />
Stefano Mariani 1 , Saeed Eftekhar Azam 1 , Aldo Ghisi 1 , Alberto Corigliano 1 , Barbara Simoni 2<br />
1 Politecnico di Milano, Dipartimento di Ingegneria Strutturale<br />
Piazza Leonardo da Vinci 32, 20133 - Milano (ITALY)<br />
2<br />
STMicroelectronics, MSH Division<br />
Via Tolomeo 1, 20010 Cornaredo (ITALY)<br />
Abstract- The dynamics of a uniaxial micro-accelerometer<br />
subjected to accidental drop events is studied by means of a<br />
reduced order model. A two degrees of freedom model is built,<br />
which carefully reproduces the MEMS response under high<br />
acceleration events. The results of the reduced order model are<br />
compared to those obtained with a three-dimensional finite<br />
element model, in terms of accuracy of the results and<br />
simulation speed-up<br />
I. INTRODUCTION<br />
During accidental drop events, polysilicon MEMS<br />
sensors are often exposed to high-g loadings because of<br />
their extremely small mass and, therefore, inertia [1-3].<br />
This fact can cause mechanical failure due to cracking in<br />
high stressed regions. In a series of papers [4-8] we<br />
recently proposed a numerical approach to accurately link<br />
the features of the shock-inducing cause (like, e.g. a drop)<br />
to the effects at the MEMS level. Because of the several<br />
length-scales affecting the dynamics of the whole device<br />
when subjected to such shocks, a multi-scale frame was<br />
adopted. We then showed that macro-scale (at device level)<br />
and meso-scale (at sensor level) analyses can be routinely<br />
investigated making use of commercial finite element<br />
codes, since the features of the polycrystalline film<br />
constituting the movable parts of the MEMS have a<br />
marginal impact. A different situation characterizes microscale<br />
(at polysilicon film level) analyses, which turn out to<br />
be extremely complicated and time demanding, in case high<br />
accuracy of the results is mandatory.<br />
A possible way to drastically reduce the computing time<br />
is to make use of reduced order models for the whole<br />
MEMS sensor, built in an accurate and micro-mechanically<br />
informed way. Reduced models would allow to avoid<br />
running analyses at the micro-scale, keeping a similar<br />
accuracy in the results. This issue was partially addressed<br />
in previous works [9-10].<br />
In the present work we go further on in the use of reduced<br />
models built on the basis of purely mechanical<br />
considerations, routed by the investigated details of the<br />
MEMS dynamics. A simple two degrees of freedom<br />
reduced model is built for a commercial micro<br />
accelerometer (Fig. 1) which measures the acceleration<br />
orthogonal to the device substrate. The dynamics of<br />
accidental drop events characterized by two acceleration<br />
levels, is numerically studied by means of the reduce model<br />
and compared with the outcome of a fully 3D finite element<br />
model.<br />
More sophisticated reduced order modelling techniques<br />
could be used, e.g. based on the proper orthogonal<br />
decomposition (POD) [11] and compared with the approach<br />
presented in this paper; this issue will be the subject of<br />
forthcoming works.<br />
II. MODELLING IMPACTS IN MEMS ACCELEROMETERS<br />
The accelerometer was assumed to be subjected to a low-g<br />
acceleration input directed along the Z-axis (see Fig. 2),<br />
whose maximum is about 90 g, and to a high-g input (see<br />
Fig. 3), with a maximum acceleration peak of about 5,500 g.<br />
As for the boundary conditions applied to the model, the<br />
accelerometer is anchored at its center with two slender<br />
suspension springs. As a reference solution a finite element<br />
model of the accelerometer, featuring 34,000 nodes and<br />
26,000 elements, i.e. about 100,000 degrees of freedom (dof)<br />
has been considered. In parallel, a reduced, two-dof model<br />
has been envisaged: if we impose a rigid behavior for the<br />
plate, only dof reproducing the vibration modes #1 and #5<br />
from Fig. 6 need to be considered. The spectral content of<br />
the two input accelerations, as shown in Fig. 4-5 through<br />
their energy spectral density, confirms that the relevant,<br />
excited dynamics pertains the relative rotation of the plate<br />
−=Δ<br />
θθθ<br />
and its relative translation −=Δ<br />
www<br />
at the<br />
spring clamped end, as shown in Fig. 1.<br />
The motion of the device in this two-dof framework can<br />
be described as follows:<br />
− aMuKu<br />
=++ (1)<br />
where the vector u=[Δw Δθ ] T collects the two dof,<br />
superposed dots indicate time derivative. If we assume h to be<br />
the plate thickness, L the plate length in the x direction, ρ the<br />
polysilicon mass density (reduced to take into account for the<br />
holed plate), l the spring length, and V the plate volume, then<br />
the mass matrix collects the plate translational mass<br />
53
ww<br />
is rotating<br />
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May 2011, Aix-en-Provence, France<br />
<br />
= ρ dyh and its inertial contribution LM<br />
when the plate<br />
∫<br />
V<br />
θθ<br />
∫<br />
= ρ<br />
V<br />
2<br />
dyyh and the coupling term.<br />
LM<br />
The stiffness matrix is diagonal, more precisely K ww =2 k f ,<br />
3<br />
K θθ =2 k t , since f = x /1 l is the flexural I stiffness 2 and Ek<br />
= tx / lIGk<br />
the torsional stiffness zt<br />
for one spring and I and I t<br />
being the bending and torsional moments of inertia of the<br />
spring cross section.<br />
The damping matrix is defined according to the quality factor<br />
Q: in our calculations we set: d ww =d θθ =Q 2 t mk , d wθ =d θw =0.<br />
A direct time integration scheme has been followed to solve<br />
(1) and a penalty coefficient amplifies the diagonal terms of<br />
the stiffness matrix when the displacements at the plate corner<br />
overcome the lower or the upper gap, therefore reproducing<br />
the plate contact with the die or the cap, which are assumed as<br />
rigid walls.<br />
Fig. 3. High-g acceleration input.<br />
z<br />
y<br />
x<br />
w<br />
θ<br />
w<br />
θ<br />
Fig. 1. Considered uniaxial MEMS accelerometer.<br />
Fig. 4. Nondimensional energy spectral density of the low-g input.<br />
Fig. 2. Low-g acceleration input.<br />
Fig. 5. Nondimensional energy spectral density of the high-g input.<br />
54
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May 2011, Aix-en-Provence, France<br />
<br />
input case, assuming the same damping coefficient as in the<br />
previous damped case. The relative displacements at two<br />
opposite (with respect to the spring axis) corner nodes for the<br />
FE and the two-dof models are depicted in Fig. 11 and, again,<br />
they well agree. It is worthwhile to point out that when one<br />
Mode 1: 1,527 Hz<br />
Mode 2: 10,047 Hz corner enters into contact with the die, since the plate is rigid<br />
the opposite node is limited in its movement as well. This<br />
confirms that for the low-g case the motion due to the<br />
excitation is mainly rotational around the spring axis, i.e.<br />
according to the degree of freedom Δθ . Again, the stresses<br />
Mode 3: 17,737 Hz<br />
Mode 4: 23,665 Hz<br />
are underestimated by the two-dof model (Fig. 12), but their<br />
absolute value is lower than the previous damped case when<br />
no contact is considered. This means that, due to the small<br />
lower gap, the die works as a stopper for the moving plate.<br />
Finally, a calculation has been carried out for the high-g<br />
acceleration input case, where both damping and contact have<br />
Mode 5: 28,945 Hz<br />
Mode 6: 164,090 Hz<br />
to be considered. As in the previous low-g case with contact,<br />
we present in Fig. 13 the relative displacements for both the<br />
opposite corner nodes. Again, the dynamics of the sensor is<br />
well captured by the two-dof model; in this case, however,<br />
when one node enters into contact with the die, the opposite<br />
corner evidences a larger displacement in the other direction.<br />
Mode 7: 259,820 Hz<br />
Mode 8: 359,310 Hz<br />
This behavior is possible only if a contribution of the<br />
translational degree of freedom Δw is also present; in other<br />
words, the plate, still rigid, translates along the Z-axis and<br />
rotates along the spring axis during high-g excitation. For the<br />
high-g case, the stresses calculated by the two-dof model are<br />
even more underestimated than the previous cases, as shown<br />
Mode 9: 526,660 Hz<br />
Mode 10: 847,540 Hz in Fig. 14.<br />
In conclusion, the good performance of the two-dof model for<br />
Fig. 6. Lowest vibration modes of the MEMS sensor.<br />
what concerns the sensor dynamics is undermined by the lack<br />
of accuracy for the stresses. Other methods like the one<br />
recalled in the following Section IV could possibly solve this<br />
drawback.<br />
III. RESULTS<br />
The advantages of the reduced order modelling in terms of<br />
computational burden are evident: in our examples the FE<br />
calculations required about 5 hours versus a few seconds for<br />
the two-dof model. In order to appreciate, instead, the quality<br />
of the reduced order model approximation Figs. 7-14 compare<br />
the FE solution with the simplified, two-dof approach, in four<br />
different cases.<br />
First, for the low-g acceleration input case, an undamped and<br />
damped response has been considered when the contact is<br />
neglected for the moving plate. In Fig. 7 and Fig. 9 the<br />
relative displacements along the Z-axis u Z at one corner node<br />
of the plate are compared: in both the cases the sensor<br />
dynamics is correctly captured (thus confirming the hypothesis<br />
of a rigid plate), an almost negligible difference is visible only<br />
in the peak of the damped case. In the Figs. 8 and 10 the<br />
envelope of maximum principal stresses at the spring clamped<br />
end section, calculated as in [9], are shown. The two-dof<br />
stresses appear clearly underestimated with respect to the FE<br />
solution; this discrepancy is due to the stress concentration<br />
effect because of the rounded corners between the plate and<br />
the spring end; this effect is captured by a refined FE mesh,<br />
but is not reproduced by the two-dof model.<br />
As a third case, we allow for contact in the low-g acceleration<br />
Fig. 7. Time history of the relative vertical displacement at the plate corner<br />
node for the low-g undamped case, neglecting contact.<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
55
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May 2011, Aix-en-Provence, France<br />
<br />
Fig. 8. Time history of the envelope of the maximum principal stress at the<br />
spring clamped end section<br />
for the low-g undamped case, neglecting contact.<br />
Fig. 11. Time history of the relative vertical displacement at the plate corner<br />
node for the low-g damped case, allowing for contact.<br />
Fig. 9. Time history of the relative vertical displacement at the plate corner<br />
node for low-g damped case, neglecting contact.<br />
Fig. 12. Time history of the envelope of the maximum principal stress at the<br />
spring clamped end section for the low-g damped case, allowing for contact.<br />
Fig. 10. Time history of the envelope of the maximum principal stress at the<br />
spring clamped end section for the low-g damped case, neglecting contact.<br />
Fig. 13. Time history of the relative vertical displacement at the plate corner<br />
node for the high-g damped case, allowing for contact.<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
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56
Fig. 14. Time history of the envelope of the maximum principal stress at the<br />
spring clamped end section for the high-g case.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
PCA is to identify the dependence structure behind a<br />
multivariate stochastic observation in order to obtain a<br />
compact description. The central idea of the PCA is to reduce<br />
the dimensionality of a data set, while retaining as much as<br />
possible the variation present in the data set.<br />
m<br />
Consider the aforementioned random variable x ∈ R ,<br />
y , y … , y are first, second,… and m th principal<br />
1 2<br />
m<br />
components respectively.<br />
Let the first principal component y 1 be a linear combination<br />
of each element of the original random vector:<br />
m<br />
T<br />
T<br />
= ( )<br />
1 ∑ x = x, α = , ,...,<br />
i1 i 1 1 11 21 m 1<br />
i = 1<br />
y<br />
The variance of y is then:<br />
1<br />
2<br />
y1<br />
α α α α α (6)<br />
S = α Ζ α<br />
(7)<br />
T<br />
1 x 1<br />
where Ζ x<br />
is the covariance of the random variable x .<br />
IV.<br />
PROPER ORTHOGONAL DECOMPOSITION<br />
The proper orthogonal decomposition (POD) is a stochastic<br />
method used to assemble the model-specific optimal linear<br />
subspace from an ensemble of system observations. Owing to<br />
the stochastic nature of the subspace calculations, the POD is<br />
also suited for nonlinear phenomena.<br />
The main idea is to find a set of ordered orthonormal bases<br />
and express samples optimally using the selected first l basis<br />
vectors.<br />
m<br />
Consider a random vector x ∈ R for which a set of arbitrary<br />
m<br />
orthonormal bases denoted by φ i<br />
spans its vector space R :<br />
{ φ i<br />
}, i + 1, 2,..., m and let us assume that the original random<br />
variable be written as a linear combination of the introduced<br />
bases, where the coefficients of this combination are denoted<br />
by y :<br />
i<br />
m<br />
x = y φ = Φy, y = φ x, i = 1, 2,...., m (2)<br />
∑<br />
i = 1<br />
i i i i<br />
( )<br />
( , ,..., ), [ , ,..., ]<br />
y = y y y Φ = φ φ φ (3)<br />
1 2 m<br />
1 2<br />
Mathematically speaking, the objective of the POD is to find a<br />
set of basis vectors that satisfies the following extreme value<br />
problem:<br />
2<br />
2<br />
min ε l = E x−x<br />
l<br />
(4)<br />
φ<br />
i<br />
( )<br />
( ) ( )<br />
s. t. φ T<br />
φ = δ , i, j = 1, 2,..., m<br />
x<br />
i j ij<br />
l<br />
∑<br />
( ) = φ ,( ≤ )<br />
l y l m (5)<br />
i = 1<br />
i<br />
i<br />
POD has been extensively used in different engineering fields,<br />
under different names; three main versions of the technique<br />
are the following:<br />
- Principal Component Analysis (PCA)<br />
- Karhunen Loéve Decomposition (KLD)<br />
- Singular Value Decomposition (SVD)<br />
Here we introduce PCA but it can be proved that the three<br />
methods are the same in essence [12]. The purpose of the<br />
m<br />
Maximum of<br />
S would not be achieved for a finite value<br />
2<br />
y<br />
of α , so a constraint have to be exerted:<br />
1<br />
T<br />
T<br />
max α Ζα , st . . α α = 1<br />
α<br />
1<br />
(<br />
1 x 1) (<br />
1 1)<br />
Introducing the Lagrangian multiplier λ it gives:<br />
1<br />
( α, ) = α Ζα +<br />
1 1 1 1 1( 1−α α<br />
1 1)<br />
T<br />
T<br />
L λ λ (9)<br />
x<br />
After differentiating it will give:<br />
∂L<br />
( α , λ )<br />
1 1<br />
= 2( Ζ −λI)<br />
α ⇒ Ζα = λ α (10)<br />
x 1 1 x 1 1 1<br />
∂α<br />
1<br />
where λ and α<br />
1<br />
1<br />
are the eigenvalue and the corresponding<br />
eigenvector of the covariance matrix Ζ x<br />
, respectively.<br />
Applying the same procedures, the objective function to be<br />
maximized in order to extract the principal components of a<br />
random variable reads:<br />
m<br />
⎛ T ⎞<br />
T<br />
max ⎜∑ α Ζα⎟, st ..( α α)<br />
= δ (11)<br />
i x i i j ij<br />
αi<br />
⎝ i = 1 ⎠<br />
and the approximation error due to representing the random<br />
l<br />
variable by the first l principal components x ≈ ∑ y α<br />
i i<br />
would be:<br />
2<br />
ε<br />
2<br />
( )<br />
( l) = E x−x( l)<br />
(<br />
i<br />
)<br />
m<br />
m<br />
2 2<br />
∑ E y ∑S<br />
y i<br />
i = l + 1 i = l + 1<br />
= =<br />
(12)<br />
where, α , i = 1, 2,..., l are the eigenvectors of Ζ<br />
i x<br />
.<br />
In order to find the principal components, one needs the<br />
covariance matrix of the random variable; however, since in<br />
practical problems it is usually impossible to find the<br />
covariance matrix, it is a common practice to use a correlation<br />
matrix as an acceptable approximation of the random variable<br />
covariance matrix. To approximate with the desired fidelity<br />
the covariance matrix one needs an appropriately chosen<br />
i =1<br />
(8)<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
57
ensemble of the random variable samples. In the jargon of<br />
MOR, such a seed of the samples is called a matrix of<br />
snapshots, where each snapshot is the state of the system in a<br />
time instant (see Fig 15).<br />
Fig. 15. Constituting matrix of snapshots<br />
The covariance of a data set allocated in a snapshot matrix<br />
X is calculated as:<br />
⎛ 1 ⎞<br />
Σ = lim ⎜Σ <br />
T<br />
= XX ⎟<br />
(13)<br />
x<br />
x<br />
n →∞⎝<br />
n ⎠<br />
and it is usually approximated by a finite number of samples,<br />
thus becoming<br />
1<br />
Σ ≈ ( X−X)( X−X)<br />
T<br />
(14)<br />
x<br />
n<br />
Once the bases are found, one can project the space of the<br />
random variable onto the new space comprising of first<br />
(hopefully) few principal components, and decrease the order<br />
of the states.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
“Two-scale vs three-scale FE analyses of shock-induced failure in<br />
polysilicon MEMS,” Proceedings of the 11 th International<br />
Conference on Thermal, Mechanical and Multiphysics Simulation<br />
and Experiments in Micro-Electronics and Micro-Systems<br />
(EuroSimE 2010), Bordeaux, France, April 2010.<br />
[9] A. Ghisi, S. Kalicinski, S. Mariani, I. De Wolf, and A. Corigliano,<br />
“Polysilicon MEMS accelerometers exposed to shocks: numericalexperimental<br />
investigation,” Journal of Micromechanics and<br />
Microengineering, vol. 19, 035023, 2009.<br />
[10] S. Mariani, A. Ghisi, R. Martini, A. Corigliano, B. Simoni.<br />
Analysis of shock-induced polysilicon MEMS failure: a multiscale<br />
finite element approach. DTIP 2010, Symposium on Design, Test,<br />
Integration & Packaging of MEMS/MOEMS, Seville (Spagna), 5-7<br />
May 2010.<br />
[11] R. A. Białecki, A. J. Kassab, A. Fic. Proper orthogonal<br />
decomposition and modal analysis for acceleration of transient<br />
FEM thermal analysis. International Journal for Numerical<br />
Methods in Engineering, vol. 62, pp. 774–797, 2005.<br />
[12] H. M. Hilber, T. J. R. Hughes and R. L. Taylor. Improved<br />
numerical dissipation for time integration algorithms instructural<br />
dynamics Earthq. Eng. Struct. Dyn. Vol. 5 283–92, 1977.<br />
ACKNOWLEDGMENT<br />
Financial support to this work has been provided by MIUR through<br />
PRIN08 grant Mechanics of microstructured materials: multi-scale<br />
identification, optimization and active control (grant #2008KNHF9Y), and<br />
by Cariplo Foundation through grant 2009: Surface interactions in micro<br />
and nano devices. The work has been also supported by Regione<br />
Lombardia and CILEA Consortium through the 2010 LISA Initiative<br />
(Laboratory for Interdisciplinary Advanced Simulation), grant: M 2 -MEMS.<br />
REFERENCES<br />
[1] G. Li, and F. Shemansky, “Drop test and analysis on micro<br />
machined structures,” Sensors and Actuators A, vol. 85, pp. 280–<br />
286, 2000.<br />
[2] V. Srikar, and S. Senturia, “The reliability of<br />
microelectromechanical systems (MEMS) in shock environments,”<br />
Journal of Microelectromechanical Systems, vol. 11, pp. 206–214,<br />
2000.<br />
[3] E. Suhir, “Is the maximum acceleration an adequate criterion of the<br />
dynamic strength of a structural element in an electronic product?”<br />
IEEE Transactions on Components, Packaging and Manifacturing<br />
Technology, vol. 20, pp. 513–517, 1997.<br />
[4] S. Mariani, A. Ghisi, A. Corigliano, and S. Zerbini, “Multi-scale<br />
analysis of MEMS sensors subject to drop impacts,” Sensors, vol.<br />
7, pp. 1817-1833, 2007.<br />
[5] S. Mariani, A. Ghisi, A. Corigliano, and S. Zerbini, “Modeling<br />
impact-induced failure of polysilicon MEMS: a multi-scale<br />
approach,” Sensors, vol. 9, pp. 556-567, 2009.<br />
[6] S. Mariani, A. Ghisi, F. Fachin, F.Cacchione, A. Corigliano, and S.<br />
Zerbini, “A three-scale FE approach to reliability analysis of<br />
MEMS sensors subject to impacts,” Meccanica, vol. 43, pp. 469-<br />
483, 2008.<br />
[7] S. Mariani, A. Ghisi, F. Fachin, F. Cacchione, A. Corigliano, and<br />
S. Zerbini, “Multi-scale analysis of polysilicon MEMS sensors<br />
subject to accidental drops: effect of packaging,” Microelectronics<br />
Reliability, vol. 49, pp. 340–349, 2009.<br />
[8] S. Mariani, A. Ghisi, R. Martini, A. Corigliano, and B. Simoni,<br />
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May 2011, Aix-en-Provence, France<br />
<br />
A Model for Two-Dimensional Arrays of Cantilevers<br />
in the Dynamic Regime<br />
Hui HUI 1, 2 , Michel LENCZNER 2<br />
1 School of Mechatronic Northwestern Polytechnical University,<br />
127, Youyi Xilu,<br />
710072 Xi’an Shaanxi, China<br />
2 FEMTO-ST, Département Temps-Fréquences Université de Franche-Comté,<br />
26, Chemin de l’Epitaphe,<br />
Abstract- We present a model for two-dimensional arrays of<br />
micro-cantilevers in elasto-dynamical regime. It has been<br />
derived by a two-scale approximation method related to<br />
strongly heterogeneous system. We also report validation<br />
results regarding its modal structure compared with the one of<br />
a direct Finite Element Model (FEM).<br />
I. INTRODUCTION<br />
Since its invention [1], the Atomic Force Microscope<br />
(AFM) has open new directions for a number of operations at<br />
the nanoscale with an impact in various sciences and<br />
technologies. A number of research laboratories are now<br />
developing large arrays of AFM that can achieve the same<br />
kind of task in parallel. The most advanced system is the<br />
Millipede from IBM [2] for data storage, but again, a number<br />
of new architectures are emerging see [3], [4], [5], [6], [7].<br />
The main limitations of the AFM devices are their low<br />
speed of operation and their low reliability, this is even more<br />
true for arrays. Thus, the modeling of single AFM and their<br />
model-based control are more and more studied, see M.<br />
Napoli et al. [8], S.M. Salapaka et al. [9], M. Sitti [10] for<br />
instance. Regarding arrays, only the group of B. Bamieh, see<br />
[8] and the reference therein, has published a model of<br />
coupled cantilever arrays. It takes into account an<br />
electrostatic coupling, and its derivation is<br />
phenomenological. It turns out that numerical simulations<br />
must handle the full array, leading to a prohibitive<br />
computational time for the time scale of a designer.<br />
The goal of this paper is to present a simplified model for<br />
the elastic behavior of large two-dimensional cantilever<br />
arrays as these appearing in AFM arrays, as depicted in<br />
Figure 1. It extends the paper [11] by taking into account the<br />
dynamical regime instead of the static regime, and it applies<br />
to two-dimensional arrays instead of one-dimensional arrays.<br />
A detailed paper has been submitted for publication, it<br />
includes all necessary information about the model and its<br />
derivation. The corresponding model for one-dimensional<br />
arrays in dynamic regime was announced in the letter [12].<br />
Our model is mainly based on a homogenization technique<br />
dedicated to strongly heterogeneous materials or systems<br />
expressed in the framework of two-scale convergence (or<br />
approximation) as introduced in M. Lenczner [13], [14] or in<br />
25030 Besançon Cedex, France<br />
D. Cioranesco, A. Damlamian and G. Griso [15]. In a<br />
preliminary step, its derivation also make use of the<br />
asymptotic method related to thin structures of P.G. Ciarlet<br />
[16] and of P. Destuynder [17]. We emphasize that the choice<br />
of a method for the modeling of the periodic array is not<br />
straightforward, in particular, a standard homogenization<br />
method is not relevant. On this point of view, a particular<br />
feature of a cantilever array is that the local mechanical<br />
displacements of the moving parts may be of the same order<br />
of magnitude as the displacements of the common support.<br />
Another point is that the lowest local eigenfrequencies of the<br />
moving parts are also in the same order of magnitude as those<br />
of the common supporting base. These features may be usual<br />
in many microsystems arrays but not in continuum<br />
mechanics for which the homogenization methods were<br />
developed. So the usual homogenization methods, do not<br />
yield interesting models even with introduction of correctors.<br />
We review the main features of our simplified model. The<br />
array is comprised of cantilevers clamped in a common base,<br />
and possibly being equipped with tips. We assume that the<br />
base is much stiffer than the cantilevers. This is expressed by<br />
saying that their stiffness have different asymptotic<br />
behaviors. The resulting model is composed of two evolution<br />
equations, one for the macroscopic behavior, related to the<br />
supporting base, and the other one at the microscopic level,<br />
which takes into account the cantilever dynamics. As<br />
required, their time scales are in the same range of magnitude<br />
and so is their mechanical displacements. We further assume<br />
that the tip is perfectly rigid, this is a commonly accepted<br />
assumption. All these assumptions yield our model with<br />
which we have carried numerical simulations and<br />
validations.<br />
The paper is organized as follows. In Section II, we start<br />
by describing the array geometry, and then shortly introduce<br />
the two-scale approximation method. Then we introduce our<br />
model in Section III. In Section IV, we discuss the<br />
eigenmodes of the model and its validation is detailed in<br />
Section V.<br />
II. THE TWO-SCALE APPROXIMATION<br />
We consider a two-dimensional array of cantilevers, see<br />
Figure 1 (a) with cell represented in Figure 1 (b).<br />
59
It is comprised of bases crossing the array in which<br />
cantilevers are clamped. The bases are connected both in the<br />
x -direction and in the x 2 -direction, so they constitute a<br />
single common support clamped on its external boundary.<br />
Cantilevers may be equipped with a rigid tip, as in AFMs.<br />
The whole array can be viewed as a periodic repetition of a<br />
same cell, in the two directions and x 2 , see also Figure 2<br />
(a) for a two-dimensional view We suppose that the numbers<br />
of rows and columns of the array are sufficiently large,<br />
namely larger or equal to 10. The simplified model will be an<br />
approximation of the three-dimensional elasticity model in<br />
the sense of small values of , the ratio of the cell size to<br />
array size i.e.<br />
To build it, we shall make use of the so-called two-scale<br />
approximation that we briefly introduce. Each point<br />
x x ,x 2 ,x of the three-dimensional space is decomposed<br />
as<br />
x x y,<br />
where x represents the coordinates of the center of the cell<br />
containing the point of x,<br />
(a)<br />
(b)<br />
Fig. 1. (a) Array of cantilevers and (b) A single cell<br />
Fig. 2. A two-dimensional view of (a) an array and (b) a single cell<br />
,and y - x-x<br />
is the dilated relative location of with respect to Points<br />
with coordinates vary in the unique so-called reference<br />
cell, that is obtained through a translation and the dilatation<br />
-<br />
of any current cell, see Figure 2 (b) for a two-dimensional<br />
<br />
<br />
view of the reference cell.<br />
Now, considering a distributed field , we introduce its<br />
two-scale transform<br />
u x,y u x y ,<br />
defined for any x x ,x 2 belonging to the two-dimensional<br />
filled section of the cell, centered at x x ,x 2 ,x , and for<br />
any y y ,y 2 ,y varying over the reference cell. We<br />
emphasize that through this construction varies in a filled<br />
rectangle covering the full array, that we refer to as . By<br />
construction, the two-scale transform is constant, with<br />
respect to its first variable x, over each cell. Since it depends<br />
on the ratio then it may be approximated by the<br />
asymptotic field, denoted by u , obtained when<br />
approaches (mathematically) 0:<br />
u u<br />
The approximation is called the two-scale<br />
approximation of u We mention that as a consequence of<br />
the asymptotic process, the partial function x u x, is<br />
continuous to the contrary of x u x, .<br />
Now, we observe that u x,y is a two-scale field, and<br />
therefore cannot be directly used as an approximation of the<br />
field u x in the real array of cantilevers. So, an inverse<br />
two-scale transform must be applied to u . However, since<br />
x u x, is continuous, u does not belong to the range of<br />
the two-scale transform. Hence we introduce an<br />
approximated inverse for the two-scale transform,<br />
v x,y v x ,<br />
in the sense<br />
(1)<br />
u u and v v ,<br />
for sufficiently regular functions u x and v x,y For x<br />
belonging to a cell centered at x , we introduce separate<br />
definitions of x v x in two parts The first one applies to x<br />
belonging to a cantilever,<br />
v x v , x x x,<br />
it is a mean in x over the cell. The other is for in the base,<br />
v x v x, x x<br />
Once an approximate inverse two-scale transform is<br />
defined, we retain u as our approximation of in the<br />
physical system. In the paper, we apply this technique to the<br />
mechanical displacements in the array, and we derive the<br />
equations governing the resulting two-scale field u<br />
III. MODEL STATEMENT<br />
Our models are formulated from the Kirchhoff-Love thin<br />
plate model of the whole structure, and we will always<br />
assume that the ratio of cantilever thickness h C to base<br />
thickness is small. More precisely, we will assume that<br />
h C<br />
,<br />
h<br />
(2)<br />
because it is the appropriate choice yielding the following<br />
non-degenerated model coupling cantilevers and base in an<br />
appropriate manner. Applying the two-scale approximation<br />
technique to the third component of the vector of mechanical<br />
displacement fields yields u t,x,y where represents the<br />
time variable and is treated as a parameter. In the following,<br />
we detail the equations governing u .<br />
From the asymptotic analysis, we find that u is<br />
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11-13 May 2011, Aix-en-Provence, France<br />
<br />
independent of y everywhere. In the model, we consider part Y R about the junction C,R<br />
in the direction y 2<br />
. Last, the<br />
cantilevers made of an isotropic material and then variations external base boundary being clamped in a fixed support<br />
of y u t,x,y are neglected. So their motions are governed<br />
u and xu n x (8)<br />
by a classical Euler-Bernoulli beam equation in the<br />
on the boundary.<br />
microscopic space variable y 2<br />
,<br />
m C<br />
tt<br />
2 u r C y 2 y 2<br />
u F C , (3)<br />
IV. EIGENMODES OF THE MODEL<br />
with m C l C m C their linear mass density, There is an infinite number of eigenvalues and<br />
r C - l CE C I C<br />
l<br />
Y C - 2 r C eigenvectors x,y<br />
their linear stiffness coefficient,<br />
2<br />
associated to the model. For<br />
convenience, we parameterize them by two independent<br />
indices, i and j , both varying in an infinite countable<br />
set. The first index refers to an infinite set of eigenvalues i<br />
and eigenvectors i<br />
x of a problem posed in the base. The<br />
eigenvalues i i<br />
constitutes a sequence of positive<br />
number increasing towards infinity. At each such eigenvalue<br />
corresponds another eigenvalue problem posed in a<br />
and F C l C F C their load per unit length. Here<br />
m C , Y ,E C ,I C , ,F C and l C are the linear mass in cantilever, the<br />
in-plane section area of the reference cell, the cantilever<br />
elastic modulus, the second moment of cantilever section, the<br />
Poisson’s ratio, a load per unit length in the antilever and the<br />
scaled cantilever width l C l C in the reference cell.<br />
This model holds for all x x ,x 2 , and therefore<br />
represents motions of an infinite number of cantilevers<br />
parameterized by x<br />
For varying along the base, the function y u t,x,y is<br />
constant and the displacement u t,x is governed by a<br />
Kirchhoff-Love plate equation<br />
ttu div x div x R x T x u<br />
l C r C y2 y 2 y 2<br />
u<br />
jun tion f , (4)<br />
with<br />
Y<br />
Y<br />
dy in R r<br />
Y<br />
r<br />
and<br />
dy are respectively its effective surface mass,<br />
its homogenized stiffness tensor, and its effective load per<br />
unit surface, where Y per unit area in the base and Y<br />
subdomain of Y . The term r C y2 y 2 y 2<br />
u<br />
jun tion is a<br />
distributed load originating from shear forces exerted by<br />
cantilevers on the base at base-cantilever junctions.<br />
At base-cantilever junctions, a cantilever is clamped in the<br />
base, so<br />
u antilever u ase<br />
2<br />
(5)<br />
and y 2 y 2<br />
u and y2 y 2 y 2<br />
u ,<br />
because yu in the base. Other cantilever ends may be<br />
free with equations,<br />
2<br />
y 2 y 2<br />
u and y2 y 2 y 2<br />
u , (6)<br />
or may be equipped with a rigid part (usually a tip in AFM),<br />
then<br />
at a junction between an elastic part and a rigid part. Here, J R<br />
is a matrix of moments and F R is comprised of effective<br />
forces and moments. For the model, this equation was<br />
restated as a boundary condition (6) at C,R<br />
where<br />
F R C<br />
dy l C G<br />
J R J J<br />
and F<br />
J J R C,R<br />
Y R<br />
2<br />
with J k YR<br />
Y R<br />
F R<br />
y 2<br />
y 2 C,R dy G 2 R<br />
(7)<br />
y 2<br />
-y 2 C,R<br />
k<br />
dy2 being a k th moment of the rigid<br />
cantilever, which has also a countable infinity of solutions<br />
denoted by C C<br />
ij and y<br />
ij 2<br />
. The index of i being fixed, the<br />
sequence ij C is a positive sequence increasing towards<br />
j<br />
infinity. On the other side, when the index is fixed, the<br />
sequence ij C , ij<br />
C<br />
i<br />
is an infinite sequence converging to an<br />
eigenelement associated to a clamped-free cantilever. We<br />
can show that the eigenvectors ij<br />
x,y 2<br />
are the product of a<br />
mode in the base by a mode in a cantilever<br />
i x C<br />
y<br />
ij 2<br />
Now we report observations made on eigenmode<br />
computations. We consider a silicon array of cells<br />
with relatively small, so that to make the results more<br />
readable. For larger , the results are qualitatively similar, so<br />
we chose , or with base dimensions: left base:<br />
m m m ; right base: m m m ; top<br />
base: m m m ; bottom base:<br />
m m m ; and cantilever dimensions<br />
2 m m 2 m for one cell. We have carried out our<br />
numerical study on both cases, with or without tips. But we<br />
limit the following comparisons to cantilevers without tips,<br />
because configuration including tips yields similar results.<br />
We restrict our attention to a finite number n<br />
2 of<br />
eigenvalues i in the base. Computing the eigenvalues ,<br />
we observe that they are grouped in bunches of size n<br />
accumulated around a clamped-free cantilever eigenvalues.<br />
A number of eigenvalues are isolated far from the bunches. It<br />
is remarkable that the eigenelements in a same bunch share a<br />
same cantilever mode shape, (close to a clamped-free<br />
cantilever mode) even if they correspond to different indices<br />
j That is why, these modes will be called "cantilever<br />
modes". Isolated eigenelements also share a common<br />
cantilever shape, which looks like a first clamped-free<br />
cantilever mode shape excepted that the clamped side is<br />
shifted far from zero. The induced global mode is then<br />
dominated by base deformations and therefore will be called<br />
"base modes". Densities of square root of eigenvalues are<br />
reported in Figure 3 for , and respectively. This<br />
figure shows three bunches for 2 cells as well as the<br />
isolated modes that remains unchanged.<br />
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11-13 May 2011, Aix-en-Provence, France<br />
V.<br />
<br />
MODEL VALIDATION WITH A DIRECT FEM 4.<br />
Our validation consists in comparing the results of our<br />
model with results obtained using a direct FEM for the<br />
three-dimensional elasticity system. We have carried our<br />
computation with a small number due to long computing<br />
time of FEM simulations. We compare the modal structure of<br />
our model to the modal structure of a FEM in<br />
three-dimensional elasticity. The eigenvalues of the<br />
three-dimensional elasticity equations also constitute an<br />
increasing positive sequence that accumulates at infinity. As<br />
for the two-scale model, its density distribution exhibits a<br />
number of concentration points and also some isolated<br />
values. Bunch sizes are still equal to the number 2 of<br />
cantilevers for low eigenvalues (log ), see Figure<br />
3 representing eigenmode distributions for , and<br />
Fig.4. Eigenmode density distributions for finite element model and<br />
for the two-scale model<br />
We remark that a number of eigenvalues in the FEM<br />
spectrum have not their counterparts in the two-scale model<br />
spectrum. We have checked that the missing elements<br />
correspond to the modes which have membrane<br />
displacement in some local cells and torsion in the<br />
cantilevers. These cases are not modeled in the current<br />
simple two-scale model. We also compare the eigenmodes<br />
and especially those belonging to bunches of eigenvalues. By<br />
visual inspection, we observe that the deformed shape of<br />
cantilevers from FEM model and from our model are similar<br />
for identical eigenvalues, see Figure 5.<br />
Fig. 3. Eigenmode Density Distributions for Finite Element Model<br />
Extrapolating this observation, we derive that when the<br />
number of cantilevers increases bunch size increases<br />
proportionally. Since the two-scale model is an<br />
approximation in the sense of an infinitely large number of<br />
cantilevers, this explains why the two-scale model exhibits<br />
mode concentration with infinite number of elements. This<br />
remark provides guidelines for operating mode selection in<br />
the two-scale model. In order to determine an approximation<br />
of the spectrum for an -cantilevers array, we suggest to<br />
operate a truncation of the mode list so that to retain a simple<br />
infinity of eigenvalues ij i , ,<br />
2 and j<br />
We stress on the<br />
fact that<br />
2 - eigenvalue bunches are generally not<br />
corresponding to a single column of the truncated matrix ij .<br />
We consider a silicon array of 10 by 10 cantilevers, see<br />
Figure 1 (a). Computing the eigenvalues associated to the<br />
two-scale model, we observe that they are grouped in<br />
bunches of size 100 accumulated around each clamped-free<br />
cantilever eigenvalue. A number of eigenvalues are isolated<br />
far from these bunches. We compare the modal structure of<br />
our model with the one of a FEM based on three-dimensional<br />
elasticity system for the same configuration. Densities of<br />
square root of eigenvalues in logarithm are reported in Figure<br />
(a)<br />
(c)<br />
(d)<br />
Fig. 5. (a) First base mode in two-scale model (b) First mode in FEM<br />
model (c) First cantilever mode in two-scale model (d) Matched mode<br />
in FEM model<br />
In a future work, we will develop a numerical test, as in the<br />
paper [18] related to one-dimensional arrays of cantilevers,<br />
so that to eliminate modes corresponding to physical effects<br />
not modeled by our model. It will be applied on transverse<br />
displacement only. We will also conduct FEM calculations<br />
for larger (more than 10) on a more powerful computing<br />
system in order to complete the convergence analysis of the<br />
solution to the FEM towards the solution of our model.<br />
In order to compare the distribution of the spectrum for a<br />
-cantilever array, we operate a truncation of mode list. It<br />
corresponds to the range [ 6 of log in Figure 4.<br />
(b)<br />
62
We have reported that how base modes alternate with<br />
cantilever modes both in our model and in the FEM model,<br />
see Figure 6 (a). The relative errors between both<br />
eigenvalues sequences are represented in Figure 6 (b). Note<br />
that errors are far from being uniform among eigenvalues. In<br />
fact, the main error source resides in a poor precision of the<br />
beam model for representing base deformations in some<br />
particular deformation modes.<br />
Fig. 6. Eigenmode density distributions for finite element model and<br />
for the two-scale model<br />
VI. CONCLUSION<br />
A two-scale model for two-dimensional cantilever arrays<br />
in dynamic regime has been derived based on a theory of<br />
strongly heterogeneous homogenization where the<br />
cantilevers play the role of soft parts. We conclude to a<br />
globally good agreement with the three-dimensional<br />
elasticity model based on eigenvalue density and mode shape<br />
comparisons. The validation of the model demonstrates that<br />
the two-scale model was sufficiently light to apply to<br />
two-dimensional AFM arrays. More comparisons with FEM<br />
results are still needed for large arrays.<br />
ACKNOWLEDGMENT<br />
This work is partially supported by the European<br />
Territorial Cooperation Programme INTERREG IV A<br />
France-Switzerland 2007-2013. The Computations have<br />
been performed on the super computer facilities of the<br />
Mésocentre de calcul de Franche-Comté.<br />
<br />
<br />
REFERENCES<br />
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and C Quate, “ x 2D FM antilever arrays a first step towards a<br />
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[4] Y.-S. Kim, H.-J. Nam, S.-M. Cho, J.-W. Hong, D.-C. Kim, and J.<br />
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[5] G.-W. Hsieh, C.-H. Tsai, W.-C. Lin, C.-C. Liang, and Y.-W. Lee,<br />
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[6] J.- D Green and G U ee, “ tomi for e mi ros opy with<br />
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[7] Z Yang, X i, Y Wang, H ao, and M iu, “Mi ro antilever<br />
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Microelectronics Journal, vol. 35, no. 5, pp. 479 – 483, 2004.<br />
[8] M. Napoli, W. Zhang, K. Turner, and B. Bamieh,<br />
“Chara terization of ele trostati ally oupled mi ro antilevers,”<br />
Journal of Microelectromechanical Systems, vol. 14, no. 2, pp. 295<br />
– 304, 2005.<br />
[9] S. M Salapaka, T De, and Se astian, “ ro ust ontrol ased<br />
solution to the sample-profile estimation problem in fast atomic<br />
for e mi ros opy,” Internat. J. Robust Nonlinear Control, vol. 15,<br />
no. 16, pp. 821–837, 2005.<br />
[10] M Sitti, “ tomi for e mi roscope probe based controlled<br />
pushing for nanotri ologi al hara terization,” IEEE/ASME<br />
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AFMs cantilever in the static ase,” Mathematical and Computer<br />
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091908, 2007.<br />
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<br />
Sensitivity Analysis and Adaptive Multi-Point Multi-<br />
Moment Model Order Reduction in MEMS Design<br />
Andreas Köhler, Sven Reitz, Peter Schneider<br />
Fraunhofer Institute for Integrated Circuits,<br />
Division Design Automation, Zeunerstraße 38,<br />
D-01069 Dresden, Germany<br />
Abstract- We present a model order reduction algorithm for<br />
linear time-invariant descriptor systems of arbitrary derivative<br />
order that incorporates sensitivity analysis for network<br />
parameters in respect to design parameters. It is based on<br />
implicit moment matching via rational Krylov subspace methods<br />
with adaptive choice of expansion points and number of moments<br />
based on an error indicator. Additionally, we demonstrate how<br />
parametric reduced order models can be obtained at nearly no<br />
extra costs, such that parameter studies are extremely<br />
accelerated. The finite element model of a yaw rate sensor<br />
MEMS device has been chosen as a numerical example, but our<br />
method is also applicable to systems arising in modeling and<br />
simulation of electromagnetics, electrical circuits, machine tools,<br />
heat conduction and other phenomena.<br />
I. INTRODUCTION<br />
For finite element models of micro- or nanoscale devices,<br />
the state space dimension easily reaches magnitudes of<br />
10 4 …10 7 . As a result, time and frequency domain simulations<br />
become computationally expensive or even impossible,<br />
especially when a coupling of several large scale subsystems<br />
is required for system level simulation. For this reason, model<br />
order reduction (MOR) emerged to an essential method for<br />
model generation for MEMS components.<br />
Typical MOR methods are balanced truncation and moment<br />
matching based on rational Krylov subspace methods [1].<br />
While balanced truncation provides a global error bound that<br />
allows easy control of the frequency domain error of the<br />
reduced order model (ROM), the computational costs of<br />
render this method inapplicable for >10 4 .<br />
In contrast, moment matching based on rational Krylov<br />
subspace methods is able to exploit the sparsity of the system<br />
matrices which yields computational costs of only .<br />
Therefore, these methods have proven as a cost efficient way<br />
to reduce the state space dimension of large scale dynamical<br />
systems during the last decades [1-6]. The lack of a global<br />
error bound does not carry weight in our applications, where<br />
the signals have a limited bandwidth and local convergence is<br />
sufficient.<br />
The biggest challenge for the integration of Krylov subspace<br />
based moment matching methods into Electronic Design<br />
Automation (<strong>EDA</strong>) software is the difficulty to choose the<br />
parameters that determine the approximation error of the<br />
ROM: the set of expansion points and the number of moments<br />
to be matched per expansion point. To overcome this, we<br />
developed an adaptive rational Krylov subspace based MOR<br />
algorithm called AMPXT [7] providing a push button solution<br />
for generating accurate reduced order models for complex<br />
structures. The algorithm automatically selects expansion<br />
points and the number of moments to be matched. Originally,<br />
AMPXT has been developed for large scale finite element<br />
models from 3D electromagnetic simulation, but the method is<br />
also applicable to other problems such as mechanical systems,<br />
heat transfer, etc. Furthermore, the adaptive strategy from [7]<br />
has been improved in this paper, such that less expansion<br />
points are involved, which reduces the overall computational<br />
costs.<br />
The second contribution of this paper is an extension of<br />
MOR for rapid network parameter sensitivity computations.<br />
This development was motivated by a growing demand for<br />
tools incorporating manufacturing tolerances during<br />
simulation and design for process variation, yield analysis, and<br />
reliability studies. Also, for design optimization tasks the<br />
influence of material and geometrical properties on the<br />
transfer characteristics of the device has to be analyzed.<br />
Finally we will demonstrate how the proposed method can<br />
be used as a simple tool for parametric MOR. This is<br />
extremely beneficial e.g. for parameter studies, where a<br />
speedup factor of 150…900 was obtained for the numerical<br />
example presented in section IV.<br />
II. PROBLEM STATEMENT<br />
Our starting point is a description of the model as a linear<br />
time-invariant descriptor system. For better readability we use<br />
a frequency domain formulation:<br />
A time domain formulation can be easily obtained via<br />
inverse Laplace transform and our proposed methodology is<br />
still applicable.<br />
The system matrices are defined as polynomials in the<br />
complex frequency with constant matrix-valued<br />
coefficients:<br />
(1)<br />
(2)<br />
(3)<br />
(4)<br />
(5)<br />
64
We assume that<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
does not become singular for more<br />
(7)<br />
than a finite set of points and<br />
and . This is the most general form of systems that<br />
(8)<br />
can be treated with the proposed MOR algorithm. Spatial Fast and accurate computation of these quantities is useful<br />
discretizations of partial differential equations like the heat for optimization loops, where the gradient of the objective<br />
equation, Maxwell’s equations, or mechanical systems as well function would have to be approximated by finite differences<br />
as RLC circuit equations fit into this framework.<br />
otherwise.<br />
The transfer function of system (1) is defined as<br />
Another objective is the acceleration of parameter studies<br />
(6)<br />
with the help of parameter dependent reduced order models,<br />
which approximate the parameter dependency of the transfer<br />
The polynomial degree of is called order of the function of the original system (1).<br />
system. But when speaking of model order reduction, we<br />
III. METHODOLOGY<br />
usually mean a reduction of the state space dimension of<br />
the system. This is a common misconception to be found in A. Model Order Reduction<br />
literature. It can be justified by the fact that every -th order The MOR method used in our methodology is a projection<br />
system can be equivalently transformed into a first order based approach. It constructs (bi-)orthonormal projection<br />
system with identical transfer function but an increased state matrices<br />
such that the reduced system<br />
space dimension of . In this sense, the terms order and<br />
state space dimension relate to each other.<br />
For systems resulting from finite element discretization, the<br />
state space dimension of the system corresponds to the total<br />
number of degrees of freedom (DOF). In mechanics, mostly<br />
second order systems arise and , and resemble the<br />
stiffness, damping, and mass matrix of the model.<br />
A frequent property of the system (1) is reciprocity, which<br />
is defined as the symmetry of the transfer function for all<br />
where is defined. This is especially the case when<br />
is symmetric for all and if a scaling function<br />
exists such that .<br />
Such systems will be called symmetric from now on.<br />
First of all, we are interested in the frequency response of<br />
the system which is obtained by evaluating for a discrete<br />
set of frequency points , ,<br />
. From (6) it follows that solutions of different<br />
linear systems of equations have to be computed, which is<br />
usually very time consuming. Time domain simulation of the<br />
system may be even impossible for large state space<br />
dimensions.<br />
The second problem to be solved is the incorporation of<br />
design parameters or manufacturing tolerances, such that the<br />
system matrices as well as the solution vector and the<br />
transfer function become parameter dependent. In the<br />
sequel we assume that only depends on a parameter<br />
. In the majority of cases the input and output matrices<br />
and respectively are incidence matrices that pick<br />
certain nodes of the model for excitation or measurement and<br />
therefore do not depend on parameters. The feed through<br />
matrix may be parameter dependent as well, but this case<br />
is omitted for the sake of simplicity as the MOR algorithm is<br />
not affected by the presence of a feed through matrix. Finally,<br />
an extension to multiple parameters is straight forward and<br />
will be demonstrated in section IV.<br />
Furthermore, we are interested in the first order sensitivities<br />
of the transfer function and the output respectively<br />
w.r.t the parameter at a given nominal value , which<br />
are given as<br />
matrices are obtained from<br />
(9)<br />
(10)<br />
(11)<br />
(12)<br />
Obviously, the corresponding reduced order model (ROM)<br />
has the same number of inputs and outputs, and . For<br />
behavior modeling and system level simulation this means that<br />
the full order model (FOM) can be seamlessly replaced by the<br />
ROM. The transfer function of the ROM is defined<br />
analogously to (6). Additionally, the -th order structure of<br />
(1) will be preserved, which prevents the costly alternative of<br />
reducing an equivalent first order system with increased state<br />
space dimension. Finally, the orthonormality of the projection<br />
matrix preserves symmetry and definite properties of the<br />
system matrices, such that passivity and stability are preserved<br />
for the ROM as well.<br />
A detailed description of the method to construct<br />
is beyond the scope of this paper, but we will give a brief<br />
outline of the essentials. First of all, we utilize multi-point<br />
moment matching. That means that a certain amount of the<br />
Taylor coefficients of the transfer function of the reduced<br />
order model for certain expansion points is equal to those of<br />
the full order transfer function. This is not to be<br />
misinterpreted as Taylor approximation in terms of a truncated<br />
Taylor series, but more like the approximation of a rational<br />
function with high numerator and denominator degrees by<br />
another rational function with lower degrees. The precise<br />
mathematical term is Padé approximation. If the system is not<br />
symmetrical and only a single projection matrix<br />
is<br />
used, we speak about Padé-type approximation. The link<br />
between moment matching and Krylov subspaces is<br />
thoroughly studied in [2]. A common synonym for multipoint<br />
moment matching is rational interpolation, indicating<br />
that the ROM interpolates the transfer function at selected<br />
frequency points up to certain derivative orders. Basically, if<br />
both -th Krylov subspaces associated with an expansion<br />
point are contained in the column spaces of and<br />
65
espectively, the resulting ROM will match at least the<br />
first moments w.r.t. , that is<br />
The resulting state space dimension of the ROM is equal<br />
to the number of columns of the projection matrices.<br />
In our case, the method of choice for the computation of the<br />
columns of the projection matrices is the Well-Conditioned<br />
Asymptotic Waveform Evaluation (WCAWE) algorithm [5].<br />
It is an efficient extension of the well-known Arnoldi method<br />
to higher order systems which prevents the otherwise costly<br />
transformation of the -th order system (1) to first order,<br />
which in turn would increase the number of rows of the<br />
projection matrix by a factor of , such that the<br />
orthonormalization process would be slowed down<br />
significantly. This is especially remarkable in view of the fact,<br />
that only a fraction of the rows of this extended projection<br />
matrix is needed for the final projection matrix, that is –<br />
depending on the specific implementation of the algorithm –<br />
the set of the first or the last rows of the extended<br />
projection matrix.<br />
For a given expansion point , we implicitly employ<br />
two parallel runs of WCAWE – one for the construction of<br />
related to and , another one for and<br />
related to the construction of . In both cases, the<br />
same factorization of<br />
is used to solve the<br />
corresponding systems of linear equations. Alternatively, one<br />
could use an extended version of the (unsymmetrical) Lanczos<br />
method, but in literature such a method does not exist for<br />
higher order multiple-input, multiple-output (MIMO) systems<br />
yet. Moreover, in the context of finite element models or<br />
electrical circuits most systems are symmetric systems in<br />
practice. Thus, the column spaces of and are<br />
identical and a single run of WCAWE is sufficient.<br />
B. Adaptive Moment Matching<br />
For the application of rational Krylov subspace MOR<br />
methods, expansion points and the number of moments per<br />
expansion point have to be selected manually by experienced<br />
users. Increasing the number of expansion points or the<br />
number of moments may or may not reduce this error, but in<br />
any case it will increase the state space dimension of the<br />
ROM. Furthermore, unless iterative methods are used, each<br />
expansion point involves a computationally expensive<br />
factorization of such that the number of expansion points<br />
has a dominant impact on the computation time needed for<br />
generating the ROM. Finally, the optimal choice of these<br />
parameters requires a-priori knowledge of system<br />
characteristics that could only be made available with high<br />
computational effort, comparable to costs of simulating the<br />
FOM.<br />
Targeting a push button solution, we developed a heuristic<br />
algorithm AMPXT [7] which can be viewed as an extension of<br />
[3] to higher order systems. The only parameter to be chosen<br />
by the end user is a frequency range of interest, i.e. an interval<br />
where the frequency response of the FOM<br />
should be approximated by the ROM up to a certain error<br />
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May 2011, Aix-en-Provence, France<br />
<br />
threshold. In contrast to other methods, AMPXT adds<br />
expansion points as they are needed and only keeps a single<br />
matrix factorization in memory at a time.<br />
(13) So, how do we heuristically determine expansion points?<br />
According to [3], real expansion points will provide more<br />
global convergence of the ROMs transfer function while<br />
expansion points on the imaginary axis provide local<br />
convergence. Depending on the invertibility of , we<br />
either choose as the initial expansion point, which<br />
would assure an exact match of the steady state or DC<br />
solution. Otherwise, we start with a real expansion point close<br />
to zero, such that is invertible. Further expansion<br />
are added as needed and they are plain imaginary, i.e.<br />
with .<br />
For each expansion point, the following error indicator is<br />
points<br />
successively evaluated in order to determine the number of<br />
columns being added to the projection matrix:<br />
(14)<br />
denotes the transfer function of the ROM resulting<br />
from the previous extension of the projection matrix by an<br />
WCAWE iteration and relates to the current ROM.<br />
The initial value for is set to the identity matrix and<br />
the norm used in (14) is the matrix infinity norm<br />
evaluated for a certain frequency .<br />
This error indicator resembles the successive relative<br />
change of the ROM’s transfer function. If there is little<br />
change, extending the ROM any further will not provide much<br />
more accuracy. To be more specific, in each iteration of<br />
AMPXT is evaluated on a discrete set of equally or<br />
logarithmically spaced frequency points:<br />
(15)<br />
If becomes small enough for all , the<br />
model is considered as converged within the frequency range<br />
of interest.<br />
Another important quantity is the relative error of the<br />
ROM’s transfer function defined as:<br />
(16)<br />
The computational costs for this quantity are equivalent to<br />
the evaluation of the full order transfer function. Therefore,<br />
cannot be computed for all . Unfortunately, an<br />
error bound that is both – sharp enough and cheaply to<br />
compute – is not known yet.<br />
This points us directly to the question on how to set a good<br />
threshold for and how to decide if adding another<br />
expansion point would be beneficial. As the exact relative<br />
error can be easily computed for the already selected<br />
expansion points by reusing the matrix factorization of<br />
, we set the convergence threshold to a value<br />
slightly above the maximum error:<br />
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May 2011, Aix-en-Provence, France<br />
(17)<br />
<br />
The whole strategy is clearly heuristic. The error indicator<br />
must not be misinterpreted as an error bound and the<br />
This may seem draconic at a first glance, as this quantity has algorithm may fail if the set of monitored frequency points<br />
a magnitude of 10 -6 …10 -12 in practice due to numerical round is not dense enough. Therefore, in step 4 we successively<br />
off related to the condition number of . Furthermore, refine in the subinterval of the currently selected<br />
one would expect that increasing should allow less candidate expansion point if the number of contained<br />
accuracy in favor of a more compact ROM being faster to frequency points is too small.<br />
simulate. But while the state space dimension of the ROM<br />
decreases indeed, its accuracy may be worse than expected,<br />
C. Efficient Computation of Sensitivities<br />
e.g. the error indicator may report a maximum of 0.1%, This section describes how the quantities from (7) and (8)<br />
but the real error according to (16) may exceed even 100% for can be computed with basically no extra costs compared to the<br />
some .<br />
evaluation of the transfer function and how MOR can be<br />
The complete strategy for building up the projection matrix incorporated to speed things up further.<br />
in an automated way can be sketched as follows:<br />
We start with the following abbreviations, where<br />
denotes the nominal value of the parameter :<br />
1. Compute the factorization of for the initial<br />
expansion point .<br />
2. Extend the projection matrices with WCAWE until either a<br />
minimum number of moments as been matched for or<br />
until the ROM has converged for at least half of the<br />
frequency points .<br />
3. Terminate if for all .<br />
4. Choose a new expansion point such that<br />
for all .<br />
5. Compute the factorization of .<br />
6. If => discard and either go to step<br />
4, if the total number of discarded expansion points does<br />
not exceed a certain threshold, otherwise the new<br />
expansion point does not seem to improve the model, so<br />
terminate.<br />
7. Extend the projection matrix with WCAWE until either a<br />
maximum number of moments is matched for or if<br />
for all .<br />
8. Terminate if for all ,<br />
otherwise go to step 4.<br />
There are several improvements over [7]. First of all, step 2<br />
puts emphasis on the initial expansion point. This is rooted in<br />
the fact, that in many applications a single real expansion<br />
point is sufficient to cover the whole frequency range of<br />
interest. As a result, fewer expansion points are needed in<br />
practice which saves computational costs. In order to improve<br />
the effect of additional expansion points, we assure in step 4,<br />
that new expansion points are placed not too close to the<br />
previous ones. Furthermore, we put a restriction on the<br />
maximum number of moments matched in step 7. This<br />
prevents stagnation of the convergence commonly found for<br />
plain imaginary expansion points in practice. Otherwise, the<br />
ROM could grow unnecessarily larger.<br />
It should also be noted that in the case of plain imaginary<br />
expansion points the resulting ROM system matrices can be<br />
forced to be real by the implicit use of complex conjugated<br />
expansion points [2]. This way, we avoid the extra costs of<br />
involving complex arithmetic for the orthonormalization and<br />
projection according to (9)-(11).<br />
(18)<br />
(19)<br />
Now, partial derivation of both equations of system (1)<br />
w.r.t. , application of the chain rule, and substitution into (6)<br />
yields an explicit expression for the transfer function<br />
sensitivity:<br />
(20)<br />
Hence, the factorization of the system matrix<br />
needed for the evaluation of the transfer function (6) for the<br />
nominal system at a given frequency can be reused.<br />
As a result, the extra costs for computing the sensitivity of the<br />
transfer function consist only of additional forward<br />
backward substitutions and a few matrix-vector products.<br />
These extra costs are negligible compared to the factorization<br />
of and can be reduced even further for the common<br />
case of symmetric systems.<br />
The evaluation of (8) is trivial as soon as has been<br />
computed. But significant acceleration of the computation of<br />
(20) can be obtained by the incorporation of model order<br />
reduction. For this purpose we assume that the system matrix<br />
for the nominal system<br />
is given as a matrix<br />
polynomial as in (2). We then construct the projection<br />
matrices for the nominal system with help of<br />
AMPXT. It is now possible to substitute the full order system<br />
matrices by the ROM’s system matrices as defined in (9)-(12):<br />
with<br />
(21)<br />
(22)<br />
The most expensive step in evaluating the transfer function<br />
sensitivity is the factorization of which has now been<br />
reduced to an problem. If is given as a matrix<br />
polynomial as well, the evaluation of (21) becomes even more<br />
efficient.<br />
It can be shown that this method of approximating the<br />
transfer function sensitivity by the sensitivity of the ROM<br />
transfer function is equivalent to the method proposed in [8].<br />
But [8] involves explicit moment matching, which is known to<br />
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be numerically less stable than implicit moment matching with<br />
<br />
Now, the sensitivity of w.r.t. can be explicitly<br />
help of the projection approach as described in section A. computed by deriving (23) and both – the system matrix for<br />
Furthermore, the error for the transfer function sensitivity is the nominal system and its sensitivity for<br />
– are<br />
proportional to the error of the state vector of the nominal represented as matrix polynomials w.r.t. analogously to (2).<br />
system [10]. This guarantees the convergence of<br />
Hence, the MOR method described in sections A-C is<br />
towards for arbitrary parameters. As a consequence, a applicable.<br />
single run of AMPXT for the nominal system is sufficient to<br />
E. Parametric Model Order Reduction<br />
allow rapid computation of sensitivities of the transfer<br />
function for a whole series of different parameters . The main use of fast sensitivity computations is accelerating<br />
At this point, we did not put any assumption on the structure optimization loops. Of course, one could estimate the<br />
of the sensitivity of the system matrix . This will be variation of the transfer function for a parameter sweep. But<br />
addressed in the next section.<br />
as we are only able to efficiently compute first order<br />
sensitivities, the parameter dependence of the frequency<br />
D. System Interpolation<br />
response cannot be accurately computed for a broad parameter<br />
The analytical dependence of the system matrices on<br />
geometrical or material parameters is not always available in<br />
practice. Most tools like finite element simulators behave like<br />
black boxes and thus only provide system matrices for fixed<br />
parameter values.<br />
For linear material properties like the Young’s modulus in<br />
mechanics or permittivity and permeability in electromagnetics,<br />
the dependency of on can be easily<br />
reconstructed such that the sensitivity of the system matrix<br />
can be explicitly computed.<br />
For geometrical parameters it has been shown that the<br />
parameter dependency can be obtained explicitly as well [9].<br />
But this would require manual extension of the source code of<br />
a finite element simulator because at the time of writing, no<br />
commercially available tool is able to provide .<br />
This is why we decided to use polynomial interpolants of<br />
the system matrices based on a series of system matrices for<br />
fixed parameter values from the neighborhood of the nominal<br />
value . The parameter dependent system matrix is then<br />
represented as a multivariate polynomial<br />
(23)<br />
with and for at least one<br />
.<br />
To be more precise, we generate an initial set of systems<br />
for equally spaced parameter values within the neighborhood<br />
of the nominal value . Starting with 2 interpolation<br />
points, we successively add more from the desired<br />
neighborhood of the nominal value using Chebychev<br />
spacing, until the following relative error for the interpolated<br />
system matrices is below a given threshold:<br />
(24)<br />
In (24),<br />
denotes the system matrix related to<br />
as defined in (2) and is a fixed parameter value from the set<br />
of systems. corresponds to the interpolant as defined<br />
in (23) and is evaluated for all and all where<br />
.<br />
range with this method. This is why we will focus on<br />
parameter dependent ROMs in this section.<br />
The topic of parametric MOR is very complex and a<br />
multitude of methods has been proposed to provide parameter<br />
dependent ROMs, see [11]-[14] and references therein. While<br />
[11] is the most general one, providing simultaneous multipoint<br />
multiple-moment matching w.r.t. both – the complex<br />
frequency and multiple parameters – it matches<br />
an equal number of moments for the complex frequency and<br />
the parameters, which results in larger ROMs, the more<br />
parameters are used. Therefore, we restrict to multi-point<br />
multiple moment matching w.r.t. to and match only the<br />
zeroth and the first moment w.r.t. the parameters. In the next<br />
section we will show that resulting parametric ROMs still<br />
capture the parameter dependency of the transfer function in a<br />
satisfactory way.<br />
Given the projection matrices<br />
for the nominal<br />
system, the reduced system matrices can be obtained<br />
simply by projection of (23) analogously to (9). This method<br />
is denoted with PMORnom from now on and instead of<br />
computing the sensitivity w.r.t. the nominal value via<br />
(21) and (22), the sensitivity of the parametric ROM can be<br />
computed directly by derivation of the reduced order system<br />
matrix w.r.t. and substitution in to (21).<br />
An alternative method called PMORinterp in the sequel is to<br />
generate ROMs for each of the previously generated<br />
FOMs for fixed parameter values and then generate a<br />
parametric ROM via polynomial interpolation. Obviously,<br />
PMORinterp takes more computation time because<br />
projection matrices<br />
have to be computed. But<br />
these extra costs lead to better accuracy for a broad parameter<br />
range as will be demonstrated in the next section. Another<br />
advantage over PMORnom is the fact that the initial FOMs are<br />
not required to have the same state space dimension. Hence,<br />
different discretizations or finite element meshes can be used<br />
for the initial FOMs. In order to assure that the states of the<br />
different ROMs used for interpolation match the same<br />
physical quantities, we apply a state transform prior to the<br />
interpolation step as proposed in [12].<br />
68
Y<br />
<br />
X<br />
<br />
Fig. 1. Scanning electron microscope image (left) and ANSYS ® finite<br />
element model (right) of a yaw rate sensor by courtesy of Robert Bosch<br />
GmbH. The nodes selected for excitation are denoted by (1) and (2).<br />
IV. NUMERICAL EXAMPLES<br />
This section demonstrates the proposed methods for<br />
adaptive MOR, fast sensitivity computations, and parametric<br />
MOR for the mechanical finite element model of a yaw rate<br />
sensor [6] as depicted in Figure 1.<br />
The full order ANSYS ® model is made of 18,508 nodes<br />
connected by 3 rd order BEAM188 elements and the total<br />
number of degrees of freedom (DOF) is 177,996. The model<br />
has 4 inputs and 4 outputs as described in Table I.<br />
The system matrices were exported to MATLAB ® with an inhouse<br />
tool and all computations were done on a laptop<br />
computer with 2.4GHz Intel Core2Duo ® and 8GB of RAM<br />
running Windows XP x64. For speed improvement compared<br />
to MATLAB‘s built in solver, a MEX-interface to the<br />
PARDISO sparse matrix solver [15], [16] was used for the<br />
computation of the matrix factorizations needed for the<br />
WCAWE iterations and for the evaluation of the transfer<br />
function of the FOM.<br />
We selected two parameters of the sensor’s model for<br />
sensitivity analysis: the Young’s modulus of the beam<br />
elements with a nominal value of<br />
GPa and the<br />
thickness of the four suspension beams in the center of the<br />
sensor with a nominal value of<br />
m. Afterwards,<br />
we generated parameter dependent FOMs with the<br />
interpolation method described in section III.D. The error<br />
threshold for the interpolation of the system matrices<br />
according to (23) and (24) was set to . Expectedly, for<br />
this results in a polynomial degree of 1, because the system<br />
matrices depend linearly on . In contrast, a polynomial<br />
degree of 4 is needed for such that 5 FOMs are involved.<br />
The topology of the finite element mesh was kept constant for<br />
TABLE I<br />
INPUT AND OUTPUT DESCRIPTION OF YAW RATE SENSOR MODEL<br />
Input 1: Y-Force at node 1 Output 1: Y-Displacement of node 1<br />
Input 2: Z-Force at node 1 Output 2: Z-Displacement of node 1<br />
Input 3: X-Force at node 2 Output 3: X-Displacement of node 2<br />
Input 4: Z-Force at node 2 Output 4: Z-Displacement of node 2<br />
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<br />
TABLE II<br />
RUNTIMES FOR THE SENSITIVITY COMPUTATIONS<br />
Evaluation of transfer function of nominal FOM for 500<br />
2203.5s<br />
frequency points<br />
Extra costs for evaluating FOM transfer function sensitivity<br />
599.4s<br />
w.r.t. via (20) for 500 frequency points<br />
Extra costs for evaluating FOM transfer function sensitivity<br />
237.9s<br />
w.r.t. via (20) for 500 frequency points<br />
Total costs for FOM evaluation 3040.8s<br />
<br />
Generation of projection matrix and ROM for nominal FOM 68.8s<br />
Evaluation of transfer function of nominal ROM for 500<br />
frequency points<br />
0.6s<br />
Extra costs for generating parametric ROM w.r.t. 1.6s<br />
Extra costs for evaluating ROM transfer function sensitivity<br />
w.r.t. for 500 frequency points<br />
0.5s<br />
Extra costs for generating parametric ROM w.r.t. 2.4s<br />
Extra costs for computing ROM transfer function sensitivity<br />
w.r.t. for 500 frequency points<br />
1.0s<br />
Total costs for ROM generation and evaluation 74.9s<br />
all parameter variations.<br />
For reference, we also computed the FOM’s transfer<br />
function and its sensitivities w.r.t. the nominal values of and<br />
at 500 logarithmically spaced frequency points between<br />
and Hz. This enabled us to compute the exact relative<br />
errors of the ROMs according to (16). The frequency of the<br />
model is almost constant between and Hz, so we omitted<br />
this range to improve readability of the plots. Finally, we the<br />
projection matrix for the nominal FOM needed for fast<br />
transfer function and sensitivity computations and for<br />
generating the parametric ROMs w.r.t. and with method<br />
PMORnom. As the FOM is symmetric, a single projection<br />
matrix was sufficient.<br />
The runtimes for the particular model generation and<br />
evaluation steps are summarized in Table II. Compared to ,<br />
the computation of the FOM transfer function sensitivities for<br />
took more than twice as much of the time. This was caused<br />
by the different number of non-zero elements in the matrix<br />
coefficients of and respectively. The former has<br />
4,258,012 non-zero elements, because applies to all beam<br />
elements, while the latter has 36,320 non-zero elements due to<br />
a more local influence of the parameter on the finite element<br />
model. Therefore, the evaluation of (20) takes more time for<br />
, even though a higher interpolation order has been used for<br />
. In contrast, it is slightly faster to generate the parametric<br />
ROM for , because the number of non-zero matrices in (23)<br />
amounts to 5 for and 15 for due to the different<br />
interpolation orders used to approximate the parameter<br />
dependency of the system matrices.<br />
Figure 2 shows the progress of the error indicator for<br />
the AMPXT run for generating the projection matrix for the<br />
nominal ROM. The frequency range of interest was set to<br />
Hz and AMPXT performed 10<br />
WCAWE iterations for the initial expansion and other<br />
4 iterations for a second expansion point with<br />
10 6 . We forced the projection matrices to be real, which<br />
doubled the number of columns added by the second<br />
expansion point. Thus, the resulting ROMs have a state space<br />
69
100%<br />
0.1%<br />
1e-04%<br />
1e-07%<br />
1e-10%<br />
1e-13%<br />
10 -2 10 0 10 2 10 4 10 6<br />
Frequency (Hz)<br />
Fig. 2. Progress of the AMPXT error indicator<br />
0.01%<br />
1e-04%<br />
1e-06%<br />
err for H(s)<br />
1e-08%<br />
err for H (s) E<br />
err for H (s) θ<br />
1e-10%<br />
10 2 10 3 10 4 10 5 10 6<br />
Frequency (Hz)<br />
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<br />
TABLE III<br />
RUNTIMES AND MAXIMUM RELATIVE ERRORS FOR PARAMETER SWEEP<br />
Fig. 3. Maximum relative errors of nominal ROM transfer function and<br />
Maximum relative error for parameter sweep w.r.t. 0.08%<br />
sensitivities according to (16)<br />
dimension of 72. Using the FOM reference solutions, we neighborhood of the nominal value.<br />
computed the exact relative errors of the ROM transfer For PMORinterp, a parameter sweep for the magnitude and<br />
function and its sensitivities w.r.t. and shown in phase of the ROM transfer function is plotted for selected<br />
Figure 3. The plots have a maximum of 0.06% and thus input-output-pairs in Figures 5 and 6. Compared to ,<br />
clearly demonstrate that the error indicator must not be clearly affects a broader portion of the frequency range.<br />
misinterpreted as an error bound, but is sufficient to monitor<br />
V. CONCLUSIONS AND OUTLOOK<br />
the convergence of ROM.<br />
In summary, the computation of 500 frequency samples of In this paper we presented an extension of existing MOR<br />
the transfer function and its sensitivities w.r.t. and takes methods for rapid and accurate computation of the frequency<br />
roughly 51 minutes for the FOM opposed to a total of 75 response and its sensitivities w.r.t. arbitrary parameters for<br />
seconds for generating and evaluating the parametric ROMs large scale finite element models. Optimization tasks with an<br />
with PMORnom. Hence, we obtained a speedup factor of objective function dependent on the transfer function will<br />
40.6.<br />
greatly benefit from the obtained speedup factor of more than<br />
For parameter sweeps, the speedup through model order 40.<br />
reduction is even more extreme. We considered an increment The lack of tools for the generation of parameter dependent<br />
of m for the suspension beam thickness , such that 21 full order models has been overcome with a system<br />
parameter values are obtained in the interval of m.<br />
interpolation approach that proved to be practical. For the<br />
sake of simplicity, we focused on interpolation of a single<br />
This results in a variation of rougly w.r.t the nominal<br />
parameter at a time, but the method can be easily extended to<br />
value of . For the Young’s modulus we considered 21<br />
multivariate polynomial interpolation involving cross terms<br />
parameter values in the interval of<br />
GPa, yielding<br />
for the interpolation polynomial. However, the more<br />
a variation of w.r.t the nominal value of . This sums parameters are involved, the more expensive the<br />
up to 41 distinct transfer function evaluations for 500<br />
10%<br />
frequency points each.<br />
Table III lists the corresponding runtimes, speedup factors, 1%<br />
and the maximum relative transfer function errors as defined<br />
in (16) taken over all 500 frequency points for all 41<br />
0.1%<br />
parameter values. Note that we did not use interpolated 0.01%<br />
parametric FOMs for the computation of the full order transfer<br />
0.001%<br />
functions which would have increased computation time even<br />
130 140 150 160 170 180 190<br />
further. Instead, we generated 41 single FOMs for fixed<br />
E (GPa)<br />
PMORnom<br />
parameter values and the time to generate these FOMs and<br />
PMORinterp<br />
export them to MATLAB ®<br />
0.001%<br />
was not accounted for the time<br />
10%<br />
measurements.<br />
1%<br />
Figure 4 shows a parameter dependent comparison of the<br />
0.1%<br />
maximum transfer function errors obtained with method<br />
PMORnom and PMORinterp over the frequency range of<br />
0.01%<br />
0.001%<br />
interest. Clearly, PMORinterp provides an accurate<br />
2.8 3 3.2<br />
approximation of the transfer function over the whole<br />
θ (µ m)<br />
3.4 3.6 3.8<br />
parameter range while PMORnom is only accurate in the Fig. 4. Maximum relative transfer function errors for parameter sweep<br />
FOM<br />
PMORnom<br />
PMORinterp<br />
Total costs for transfer function parameter sweep for 500<br />
frequency points and 41 parameter values<br />
25.1h<br />
Total costs for generation of param. ROMs w.r.t. and 74.9s<br />
Total costs for parameter sweep w.r.t. and 27.2s<br />
Speedup factor for parameter sweep<br />
885x<br />
Maximum relative error for parameter sweep w.r.t. 1.07%<br />
Maximum relative error for parameter sweep w.r.t. 78.93%<br />
Total costs for generation of parametric ROM w.r.t. 156.2s<br />
Total costs for generation of parametric ROM w.r.t. 371.8s<br />
Total costs for parameter sweep w.r.t. and 27.6s<br />
Speedup factor for parameter sweep<br />
163x<br />
Maximum relative error for parameter sweep w.r.t. 0.84%<br />
70
10 3<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
10 3<br />
Magnitude<br />
10 1<br />
10 -1<br />
10 -3<br />
10 -5<br />
10 2 10 3 10 4 10 5 10 6<br />
0<br />
-50<br />
in1/out1<br />
in2/out2<br />
in3/out3<br />
in4/out4<br />
Magnitude<br />
10 1<br />
10 -1<br />
10 -3<br />
10 -5<br />
10 2 10 3 10 4 10 5 10 6<br />
0<br />
-50<br />
in1/out1<br />
in2/out2<br />
in3/out3<br />
in4/out4<br />
Phase<br />
-100<br />
Phase<br />
-100<br />
-150<br />
-150<br />
10 2 10 3 10 4 10 5 10 6<br />
Frequency (Hz)<br />
Fig. 5. PMORinterp transfer function sweep for<br />
GPa,<br />
curves for nominal value<br />
GPa are emphasized<br />
computational costs will be (curse of dimensionality).<br />
PMORnom is a very efficient method allowing parameter<br />
sweeps for an arbitrary amount of parameters at the<br />
computational costs of only a single run of AMPXT for the<br />
nominal system. But for complex parameter dependencies, it<br />
does not deliver good accuracy for a broad parameter range as<br />
we demonstrated with the thickness of the suspension beam in<br />
section IV. In contrast, PMORinterp manages to capture the<br />
parameter dependence of the transfer function with a<br />
maximum relative error of less than 1% at the expense of a<br />
smaller speed up. Also, the more parameters are to be<br />
investigated, the more expensive PMORinterp becomes.<br />
For the application of (parametric) ROMs in system<br />
simulation, we have in-house tools to export these ROMs to<br />
behavior modeling languages like VHDL-AMS, Verilog-A,<br />
MAST ®, MATLAB ® /Simulink ® , and others. Hence, the<br />
ROMs can be coupled with other models for accelerated<br />
analysis of the system behavior.<br />
As part of current research, we seek to combine the<br />
promising work of [11] with an adaptive strategy to<br />
automatically select interpolation points, expansion points and<br />
the number of moments to be matched simultaneously w.r.t.<br />
multiple parameters .<br />
ACKNOWLEDGEMENT<br />
This paper is partially based on the project DIONYSYS<br />
which is supported by the German Federal Ministry of<br />
Education and Research under Grant No. 01M3084G. The<br />
authors of this paper are solely responsible for its content.<br />
REFERENCES<br />
[1] Antoulas, A. C.: Approximation of Large-Scale Dynamical Systems.<br />
Philadelphia, PA, USA: Society for Industrial and Applied Mathematics,<br />
2005.<br />
[2] Grimme, E. J.: Krylov projection methods for model reduction. Ph.D.<br />
dissertation, University of Illinois, 1997.<br />
[3] Grimme, E. J. and Gallivan, K.: A Rational Lanczos Algorithm for<br />
Model Reduction II: Interpolation Point Selection. Numerical<br />
Algorithms, vol. 12, pp. 33-63, 1998.<br />
[4] Reitz, S.; Bastian, J.; Haase, J.; Schneider, P.; Schwarz, P.: System level<br />
modeling of microsystems using order reduction methods. Symp.<br />
10 2 10 3 10 4 10 5 10 6<br />
Frequency (Hz)<br />
Fig. 6. PMORinterp transfer function sweep for<br />
m,<br />
curves for nominal value<br />
m are emphasized<br />
Design, Test, Integration and Packaging of MEMS/MOEMS", Cannes,<br />
France, 5-8 May 2002<br />
[5] Slone, R. D.; Lee, R.; Lee, J.-F.: Well-conditioned asymptotic waveform<br />
evaluation for finite elements. IEEE Trans. Antennas Propag., vol. 51,<br />
2003, pp. 2442-2447<br />
[6] Reitz, S.; Döring, C.; Bastian, J.; Schneider, P.; Schwarz, P.; Neul, R.:<br />
System level modeling of the relevant physical effects of inertial sensors<br />
using order reduction methods. Proceedings: Symposium on Design,<br />
Test, Integration and Packaging of MEMS/MOEMS, Montreux,<br />
Switzerland, 12 - 14 May 2004, pp.383-387<br />
[7] Köhler, A.; Reitzinger, S.: An adaptive multi-point multi-moment model<br />
order reduction algorithm for fast broadband simulation of large-scale<br />
3D electromagnetic models. In: Sommer, Ralf (ed.): ANALOG 2010.<br />
Entwicklung von Analogschaltungen mit CAE-Methoden: Vorträge der<br />
11. ITG/GMM-Fachtagung vom 22. bis 24. März 2010 in Erfurt, VDE-<br />
Verlag, 2010 (ITG-Fachbericht 221), S. 39-52<br />
[8] Webb, J.P.: Design sensitivity of frequency response in 3-D-finiteelement<br />
analysis of microwave devices. IEEE Trans. Magn., vol.38,<br />
2002, pp. 1109-1112<br />
[9] Webb, J.P.; , Finite-element analysis of the effect of geometric tolerances<br />
on performance over a frequency band. IEEE Transactions on<br />
Microwave Theory and Techniques, vol.52, no.1, pp. 306- 310, Jan.<br />
2004<br />
[10] Köhler, A.; Dyczij-Edlinger, R.; Farle, O.; Lohmann, B. (ed.) and Kugi,<br />
A. (ed.): “Schnelle Berechnung von Empfindlichkeiten in 3D EM<br />
Strukturen mittels Modellordnungsreduktion”. Tagungsband GMA<br />
Fachausschuss 1.30 Modellbildung, Identifikation und Simulation in der<br />
Automatisierungstechnik, Wien, 2010.<br />
[11] Farle, O. and Dyczij-Edlinger, R.: Numerically Stable Moment Matching<br />
for Linear Systems Parameterized by Polynomials in Multiple Variables<br />
With Applications to Finite Element Models of Microwave Structures.<br />
IEEE Transactions on Antennas and Propagation, 2010.<br />
[12] Panzer, H.; Mohring, J.; Eid, R. and Lohmann, B.: Parametric Model<br />
Order Reduction by Matrix Interpolation, at - Automatisierungstechnik,<br />
Oldenbourg Wissenschaftsverlag GmbH, 2010, vol.58, pp. 475-484<br />
[13] Köhler, A.; Clauß, C.; Reitz, S., Haase, J.; Schneider, P.; Troch, I. (ed.)<br />
and Breitenecker, F. (ed.): Snapshot-Based Parametric Model Order<br />
Reduction. Proceedings MATHMOD 09 Vienna - Full Papers CD<br />
Volume, 2009.<br />
[14] Köhler, A.; Reitz, S., Clauß, C.; Schneider, P. and Haase, J.:<br />
Parametrische Modellordnungsreduktion bei der automatisierten<br />
Modellgenerierung für den Elektronik- und Mikrosystemtechnikentwurf.<br />
9. Chemnitzer Fachtagung Mikrosystemtechnik, Chemnitz, 2009.<br />
[15] Schenk, O. and Gärtner, K.: Solving Unsymmetric Sparse Systems of<br />
Linear Equations with PARDISO, Journal of Future Generation<br />
Computer Systems, 20(3):475--487, 2004.<br />
[16] Schenk, O. and Gärtner, K.: On fast factorization pivoting methods for<br />
symmetric indefinite systems, Elec. Trans. Numer. Anal., vol. 23, pp.<br />
158-179, 2006<br />
71
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Integration of Ferroelectric BaTiO 3 on Metallic Ni<br />
Tapes for Power Generation<br />
Greg Collins, Emanuel Silva, Ming Liu, David Elam, Chunrui Ma,<br />
Andrey Chabanov, Arturo Ayon and Chonglin Chen<br />
The University of Texas at San Antonio<br />
One UTSA Circle<br />
San Antonio, TX 78249, USA<br />
Jie He, Jiechao Jiang and Efstathios Meletis<br />
The University of Texas at Arlington<br />
Arlington, TX 76019, USA<br />
Abstract- Ferroelectric BaTiO 3 thin films were integrated<br />
directly on metallic Ni tapes by using pulsed laser for energy<br />
harvesting applications. Microstructure studies from x-ray<br />
diffraction and electron microscopy indicate that the as-grown<br />
BaTiO 3 thin films have pure BaTiO 3 crystal phase which<br />
consists of the crystalline assemblage of nanopillars with<br />
average cross sections from 100 nm to 200 nm directly on the<br />
Ni tapes. The BaTiO 3 films have good interface structures and<br />
strong adhesion to the Ni metallic tapes. Dielectric<br />
measurements have shown the hysteresis loop at room<br />
temperature in the film with a large remnant polarization,<br />
indicating that the ferroelectric domains have been created in<br />
the as-deposited BTO films. The successful integration of<br />
ferroelectric thin films directly on metallic materials is<br />
considered to be very promising for the development of<br />
energy harvesting devices.<br />
I. INTRODUCTION<br />
Ferroelectric materials have been considered as the most<br />
important materials for energy harvesting and data<br />
storage due to their high dielectric constant and good<br />
insulating properties. Among them, Barium Titanate,<br />
BaTiO 3 (BTO), is one of the most important ferroelectric<br />
materials that has attracted great attention for its remarkable<br />
characteristics including high dielectric constant, good<br />
ferroelectric properties, and large electro-optic and nonlinear<br />
optic coefficients. Furthermore, this material has<br />
excellent piezoelectric properties resulting in broad<br />
applications in control systems, structural health monitoring<br />
and energy harvesting. Therefore, the major challenge is to<br />
successfully integrate BTO thin films directly on metallic<br />
substrates with optimum metal/film interface properties for<br />
various device applications such as supercapacitance and<br />
power generation, among others. In fact, various techniques<br />
have been developed to fabricate ferroelectric BTO thin<br />
film for device fabrications.<br />
Recently, BTO thin films have been deposited on various<br />
substrates including oxide single crystal and semiconductor<br />
substrates using a variety of techniques such as pulsed laser<br />
deposition (PLD), hydrothermal method, sol-gel processing,<br />
solid-state reactions, and metal-organic chemical vapor<br />
deposition [1-6]. However, many challenges remain,<br />
especially the interface-related issue observed when<br />
fabricating ferroelectric thin films on structural materials<br />
(steel, aluminum, titanium, etc.) for energy harvesting<br />
device development. Publications describing the fabrication<br />
of ferroelectric thin films on metallic materials were not<br />
available until the reports of our recent achievements of insitu<br />
fabrication of BTO thin films on the typical structural<br />
material Ni using PLD system [7-8]. In the report contained<br />
herein, we describe our recent achievements on the<br />
fabrication of ferroelectric BTO thin films directly on<br />
metallic Ni tapes with good crystallinity and excellent<br />
dielectric properties.<br />
II. EXPERIMENTAL<br />
BaTiO 3 thin films were deposited on amorphous nickel<br />
substrates in a PLD system using a KrF excimer laser with a<br />
wavelength of 248 nm with an energy density of about 2.5<br />
J/cm 2 and a laser repetition rate of 5Hz. The BTO thin films<br />
were fabricated with details that can be found from the<br />
literatures [7-8]. X-ray diffraction (XRD) was employed to<br />
understand the crystal phases and the transmission electron<br />
microscopy (TEM), plan-view and cross-section, were<br />
employed to study the microstructure of the as-grown films<br />
and interfacial layers. The dielectric properties were<br />
characterized by using a Radiant RT6000 for understanding<br />
the physical properties of the as-grown films and an Agilent<br />
AFM/PFM with lock-in amplifier was used to observe the<br />
piezoelectric response.<br />
III. CHARACTERIZATION<br />
Fig. 1 is the XRD θ-2θ pattern from the as-deposited<br />
BTO thin film on Ni showing that all the peaks are from the<br />
polycrystalline BTO phases and polycrystalline Ni<br />
substrate. These peak positions suggest that the Ni substrate<br />
is cubic phase and the BTO layer belongs the tetragonal<br />
phase.<br />
72
Intenstiy (counts/second)<br />
Fig. 1. X-ray data showing relative peak intensity.<br />
It is surprisingly found that the as-grown BTO film has a<br />
preferred c-axis oriented revealing from the stronger<br />
intensity from the (200) diffraction in the BTO film. The<br />
BTO films were found to have the tetragonal structure with<br />
lattice constant a = 4.00Å and c=4.03 Å. These results were<br />
verified by using both cross sectional and plan view TEM<br />
techniques. As seen in Fig. 2, the selected-area electron<br />
diffraction (SAED) pattern shows the as-grown films have a<br />
pure crystalline phase with a tetragonal structure with a<br />
space group of p4mm and lattice parameter of a=3.992Å<br />
and c=4.036 Å. The diffraction rings numbered in 1, 2, 3, 4,<br />
5 and 6 have a lattice spacing of 4.0 Å, 2.8 Å, 2.3 Å, 2.0<br />
Å, 1.8 Å, 1.64 Å and 1.4 Å, respectively, which can be<br />
identified as the (001), (101), (111), (002), (102) and (112)<br />
reflection of tetragonal BTO. The BTO films have a lateral<br />
width from 30 nm to 100 nm, as seen from the plan-view<br />
TEM. The BTO film is found to be very well bound the Ni<br />
tape with a sharp NiO interlayer in between with a thickness<br />
of about 100 nm. The BTO film has a thickness of about<br />
200 nm and consists of nanopillar structures with lateral<br />
dimensions of about 100 nm. This great achievement<br />
suggests that the BTO films directly on Ni tape has paved a<br />
way to develop supercapacitance devices and power<br />
generation for the energy harvest applications.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
To further understand the growth nature and physical<br />
property of the as-grown BTO films, Piezoelectric<br />
responsive Microscopy (PRM) was employed to study the<br />
multi- domain structures and domain distributions. As seen<br />
figure 3, ferroelectric domains are mainly perpendicular to<br />
the film surface with uniform polarization although there<br />
are about 15% ferroelectric domains with in plane<br />
polarized. The PRM results further confirm that the BTO<br />
film has a preferred c-axis oriented. The ferroelectric<br />
polarization hysteresis loop measurement was also<br />
performed at room temperature. The ferroelectricity was<br />
successfully achieved on the as-deposited BTO film. The<br />
room temperature ferroelectric response shows data for<br />
spontaneous polarization, remnant polarization, and<br />
coercive field from the as-deposited BTO layer can be seen<br />
in Fig. 3. It is known that the lattice dipole along the c-axis<br />
for a tetragonal perovskite structure provides the strongest<br />
ferroelectric properties associated with BTO. With BTO<br />
having d-components in the in- and out-of-plane directions<br />
according to the equation giving an optimal angle for<br />
displacement as ~52° [9], the ability to acquire specific<br />
orientation of the film is desirable for specific device<br />
application. The a-axis oriented BTO film cannot show<br />
ferroelectric hysteresis due to the randomly oriented<br />
polarization, the large spontaneous polarization obtained in<br />
the as-deposited film is consistent with the result of the<br />
microstructure measurement that the film has highly c-axis<br />
oriented texture structure. It is surprisingly found that the<br />
as-grown BTO films on Ni metal tapes with a NiO buffered<br />
layer exhibit very high resistivity value of 10 10 Ω•cm which<br />
well suits BTO’s ferroelectric reliance on a high dielectric<br />
constant. The ferroelectricity of the BTO films was<br />
evidenced from the hysteresis loop. The room temperature<br />
spontaneous polarization, remnant polarization, and<br />
coercive field from the as-deposited BTO layer can be<br />
obtained to be about 35 µC/cm 2 and 15 µC/cm 2 ,<br />
respectively, with a coercive field of 25 kV/cm. The<br />
piezoelectric response (Fig. 4) of the as-deposited BTO film<br />
was surprisingly found to be 130 (x 10 -12 C/N) which is<br />
about 30% larger than the values (90 – 100 x 10 -12 C/N) of<br />
BTO single crystalline and polycrystalline bulk<br />
materials.<br />
Fig. 2. TEM and SAED imagery showing grain characteristics.<br />
Fig. 3. Polarization response of BTO films at selected voltages.<br />
73
The large piezoelectric response might result from the<br />
uniform nanodomain structures as well as the NiO interlayer<br />
with a lattice constant of 4.18Å that closely matches BTO’s<br />
parameters. The BTO films offer a lead-free option to PZT<br />
and are found to be more chemically stable than some other<br />
materials when applied to metal substrate [10-11]. The<br />
nature of the mechanisms is under investigation and will be<br />
reported later on.<br />
Fig. 4. Piezoresponsive Force Microscopy measurement<br />
showing active areas.<br />
IV. SUMMARY<br />
In summary, we have demonstrated achievability to grow<br />
ferroelectric BaTiO 3 thin films directly on metallic Ni<br />
substrates by optimizing the growth parameters and<br />
conditions. The as-deposited BTO films have nanopillar,<br />
crystalline tetragonal structures with a good interface with<br />
respect to the substrate. The microstructural studies reveal<br />
that BaTiO 3 films are composed of crystalline assemblage<br />
of nanopillars with average cross sections from 100 nm to<br />
200 nm. The room temperature ferroelectric polarization<br />
measurements show that the ferroelectric domains have<br />
been created in the as-deposited BTO films. Successful<br />
fabrication of such ferroelectric films on the metallic<br />
substrates has significant importance for the development of<br />
new applications such as supercapacitance for energy<br />
storage and power generation for energy harvest. The work<br />
can be extended to integrate other ferroelectric oxide films<br />
with various promising properties to monitor the structural<br />
health materials and the energy harvest applications.<br />
ACKNOWLEDGMENT<br />
This work is partially supported by the National<br />
Science Foundation under Award Number NSF/CMS-<br />
0528873 and NSF/CMMI-0709293, the Army Research<br />
Office under award number 54484-RT-ISP, and the State of<br />
Texas through the TcSUH.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
REFERENCES<br />
[1] G. M. Davis, and M.C. Gower, “Epitaxial growth of thin films of<br />
BaTiO 3 using excimer laser,” Appl. Phys. Lett., vol. 55, pp. 112-<br />
114, July 1989.<br />
[2] K. Kajiyoshi, N. Ishizawa, and M. Yoshimura, “Heteroepitaxial<br />
growth of BaTiO 3 thin-films on SrTiO3 substrates under<br />
hydrothermal conditions,” Jpn. J. Appl. Phys., vol. 30, pp. L120-<br />
L123, April 1990.<br />
[3] S. Song, J. Zhai, and X. Yao, “Effects of buffer layer on the<br />
dielectric properties of BaTiO 3 thin films prepared by sol–gel<br />
processing,” Mater. Sci. and Engin. B, vol. 145, pp. 28-33,<br />
December 2007.<br />
[4] T. Garcia, P. Bartolo-Perez, E. de Posada, J.L. Pena, and M.<br />
Villagran-Muniz, “Studies of Pulsed Laser Deposition. Processes<br />
of BaTiO 3 Thin Films,” Surf. Coat. Techno., vol. 201, pp. 3621-<br />
3628, February 2006.<br />
[5] C.H. Lei, “The growth of BaTiO 3 films on (001) MgAl 2O 4<br />
substrates by pulsed laser deposition technique,” Thin Solid Films,<br />
vol. 515, pp. 1701-1707, December 2006.<br />
[6] A. Graff, S. Senz, D. Völtzke, H.P. Abicht, and D Hesse,<br />
“Microstructure evolution during BaTiO 3 formation by solid-state<br />
reactions on rutile single crystal surfaces,” J. Euro Ceramic Soc.,<br />
vol. 25, pp. 2201-2206, 2005.<br />
[7] Z. Yuan, J. Liu, J. Weaver, C. L. Chen, J. C. Jiang, B. Lin, V.<br />
Giurgiutiu, A. Bhalla, and R. Y. Guo, “Ferroelectric BaTiO 3 thin<br />
films on Ni metal tapes using NiO as buffer layer,” Appl. Phys.<br />
Lett., vol. 90, 202901, 2007.<br />
[8] J. C. Jiang, E. I. Meletis, Z. Yuan, J. Liu, J. Weaver, C. L. Chen, B.<br />
Lin, V. Giurgiutiu, R. Y. Guo, A. S. Bhalla, D. Liu, and K. W.<br />
White, “Orientation Preferred Structures in BaTiO3 Thin Films on<br />
Ni Substrates,” J. Nano Res., vol. 1, pp. 59-63, June 2008.<br />
[9] J. L. Ruglovsky, J. Y. Li, K. Bhattacharya and H. A. Atwater, “The<br />
effect of biaxial texture on the effective electromechanical<br />
constants of polycrystalline barium titanate and lead titanate thin<br />
films,” Acta Mater., vol. 54, pp. 3657-3663, March 2006.<br />
[10] J. G. Wu, Y. Wang, X. Yuanyu, D. Q. Xiao, J. G. Zhu, and Z. H.<br />
Pu, “Effects of Ag content on the phase structure and piezoelectric<br />
properties of (K 0.44-xNa 0.52Li 0.04Ag x)(Nb 0.91Ta 0.05Sb 0.04)O 3 lead-free<br />
ceramics,” Appl. Phys. Lett., vol. 91, pp. 132914-132914(3), June<br />
2009.<br />
[11] E. Ringgaard and T. Wurlitzer, “Lead Free CaTiO 3-Based<br />
Ceramics: Sintering, Phase Transitions and Dielectric Properties,”<br />
J. Eur. Ceram. Soc., vol. 25, pp. 2701-2706, 2005.<br />
74
11-13 May 2011, Aix-en-Provence, France<br />
, ,<br />
<br />
An Electromechanical Model for clamped-clamped Beam Type Piezoelectric<br />
Transformer<br />
Chi-Shao Chen 1 , Chia-Che Wu 2*<br />
1<br />
Graduate Student<br />
2*<br />
Assistant Professor 1<br />
1,2<br />
Department of Mechanical Engineering, National Chung Hsing University,<br />
250, Kuo Kuang Road, Taichung, Taiwan, 402<br />
Tel: +886-4-22840433 ext 419; Fax: +886-4-22877170;<br />
E-mail: josephwu@dragon.nchu.edu.tw<br />
Abstract- In this paper, an analytical solution of a fixed-fixed<br />
beam type piezoelectric transformer with Euler-Bernoulli beam<br />
assumption is proposed. The electromechanical equations are<br />
first derived for transient motions, and coupled expressions for<br />
the mechanical response and voltage output are obtained. The<br />
resulting equations are further reduced for the case of excitation<br />
around the first resonance frequency. Analyical solutions of<br />
mechanical response, voltage, current, and power outputs are<br />
presented. From analytical model, output voltage depends on<br />
the lengthes of two electrodes, the length of beam, and the<br />
Young’s modulus ratio and thickness ratio between PZT layer<br />
and substrate. The lengthes of input electrodes and output<br />
electrodes should be 0.22 time length of beam to achieve the<br />
largest output when the transformer is excited at first resonance<br />
frequency. The output voltages and the resonance frequencies of<br />
transformers are proportional and inversely proportional to the<br />
lengthes of beams, respectively. The combination of Young’s<br />
modului and thicknesses of PZT layer and substrate change the<br />
position of netural axis and the bending stiffness of beam,<br />
concurrently. However, output voltages of transformers depend<br />
not only on the postion of neutral axes but also on bending<br />
stiffnesses.<br />
I. Introduction<br />
Piezoelectric materials have the piezoelectric effect,<br />
which can convert vibration energy into electrical energy, so<br />
it can be used to make transformers for raising or lowering a<br />
voltage. Piezoelectric transformer (PT) offers many<br />
advantages over the small size, lighter with flat structure,<br />
electromagnetic field immunity, and high transforming ratio.<br />
The idea of a PT was first implemented by Rosen in 1956[1].<br />
It used the effect of couple between electrical and mechanical<br />
energy of piezoelectric materials. Exciting mechanical<br />
vibrations by the part of driver and output voltage can be<br />
induced by the part of generator. Most of the PTs are using<br />
the concept of Rosen-type PT such as uniformly-poled<br />
longitudinal PT [2], stacked disk-type PT [3, 4], and<br />
uniformly-poled disk type PT [5]. M. C. Do et al. used<br />
parallel connection of Rosen-type PTs to increase output<br />
power [6]. T. Inous et al.[7] developed a PT which is a<br />
combination of a longitudinal mode piezoelectric actuator<br />
and a longitudinal mode piezoelectric transducer transverse<br />
in parallel to achieve larger power. However, operating<br />
frequencies of transformers in the literature were usually<br />
from a few kHz to several hundred MHz. External oscillator<br />
and control circuit are required to satisfy the frequency<br />
requirement. However, they will substantially expend the size<br />
and the complexity of transformers.<br />
The PT is not only a mechanical system but also an<br />
electrical system. The electromechanical model approaches<br />
in the recently literature include single degree-of-freedom<br />
(SDOF) models [8], Rayleigh-Ritz method[9], equivalent<br />
circuit method[10], and expansion theory based on the Euler-<br />
Bernoulli beam assumptions [11]. The SDOF modeling<br />
approaches supposes a structure such as a cantilevered beam<br />
as a mass-spring-damper system which is convenient for<br />
coupling the mechanical part and electrical part of<br />
transformer. However, SDOF is just a simple approximation<br />
and it is limited to a single vibration mode. SDOF lacks of<br />
several important information of the system, such as the<br />
dynamic mode shape, the accurate strain or stress distribution<br />
along the beam. Rayleigh-Ritz method is a numerical<br />
approximation technique based on discretization of the<br />
continuous distributed parameter system and it allows<br />
predicting the electromechanical response in higher vibration<br />
modes. The Rayleigh-Ritz method can produce accurate<br />
results with only a small number of terms in the<br />
approximating series, which translates into a discrete model<br />
with a small number of degree of freedom. However, the<br />
Rayleigh-Ritz method can’t use in complex geometry and it’s<br />
not an exact solution. Equivalent circuit model is used to<br />
estimate the electrical characteristics of the PT, such as<br />
voltage ratio between input and output. But, it has no idea<br />
about the mechanical information since all parameters are<br />
transferred into electrical form and some coupled coefficient<br />
must be obtained from the experiments.<br />
Erurk and Inman [11] presented the exact<br />
electromechanical solution of a cantilevered piezoelectric<br />
energy harvester with Euler-Bernoulli beam assumptions.<br />
The electromechanical equations were derived for general<br />
transient motions from expansion series and coupled<br />
expression (not only single vibration mode) for mechanical<br />
response and voltage output were obtained. This method<br />
provides exact solutions of energy harvester. They also used<br />
internal strain rate damping and external air damping to<br />
achieve more accurate model. Backward coupling effect in<br />
the mechanical domain and the contribution from the other<br />
vibration mode were also considered in their model.<br />
In this paper, an analytical solution of a fixed-fixed beam<br />
type piezoelectric transformer with Euler-Bernoulli beam<br />
assumption is proposed. A clamped-clamped beam<br />
transformer consist of a fixed-fixed beam, a layer of<br />
75
piezoelectric film, a pair of electrodes for driver section and<br />
another pair of electrodes for generator section . The pair of<br />
electrodes for driver is located near one end of fixed-fixed<br />
beam, and another pair of electrodes for generator is located<br />
near the other end. Input voltage is applied to electrodes for<br />
driver to excite fixed-fixed beam, and output voltage is<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
generator from electrodes for generator. The input voltage is<br />
assumed to be harmonic in time. The resulting expressions<br />
for the coupled mechanical response and the electrical are<br />
then reduced for the particular case of harmonic behavior in<br />
time. Simple expressions for mechanical response, voltage,<br />
current, and power outputs are also presented.<br />
II. Derivation of the analytical model<br />
We consider the transformer shown in Fig. 1, which is a<br />
clamped-clamped uniform composite Euler-Bernoulli beam.<br />
A layer of PZT and two pairs of electrodes are perfectly<br />
bonded to the substrate. Input voltage is applied to one pair<br />
of electrodes on PZT layer to excite clamped-clamped beam,<br />
and output voltage is generated from the other pair of<br />
electrodes on PZT layer. The energy flows from the electrical<br />
energy of the input to the mechanical fields of vibration, then<br />
back to the output electric energy. Two pairs of electrodes are<br />
assumed to be perfectly conductive, and they cover surface of<br />
the PZT at the bottom and at the top. The lengths of<br />
electrodes is much larger than the thicknesses of PZT so that<br />
the electrical filed is uniform over the length of electrodes.<br />
The simple electrical circuit consists of a resistive load. The<br />
leakage resistance of PZT is much higher than the load<br />
resistance and it can be neglected in the electrical circuit. The<br />
capacitance of the PZT is considered as internal to the PZT,<br />
and it is not ignored although it is not shown in figure 1. The<br />
capacitance term will be shown in piezoelectric constitutive<br />
relations. The transformer is excited by one pair of PZT in<br />
figure 1. The Governing equation of motion can be written as<br />
[1]<br />
2<br />
5<br />
2<br />
∂ M x,<br />
t ∂ wx,<br />
t ∂wx,<br />
t ∂ wx,<br />
t<br />
c m 0<br />
2 s<br />
I c<br />
4<br />
a<br />
(1)<br />
2<br />
∂x<br />
∂x<br />
∂t<br />
∂t<br />
∂t<br />
Where w(x,t) is the transverse deflation of the beam<br />
relative to natural axis, c s I is the equivalent damping term of<br />
the composite cross section due to structural viscoelasticity<br />
( c s<br />
is the equivalent coefficient of strain rate damping and I<br />
is the equivalent area moment of inertia of PZT-substrate<br />
composite cross section), c a is the air damping coefficient, m<br />
is the mass per unit length of the beam, M is internal moment<br />
of cantilever can be written as [2]<br />
M<br />
<br />
x<br />
<br />
<br />
<br />
2<br />
∂ w x,<br />
t<br />
YI v(<br />
t)<br />
(2)<br />
∂x<br />
, t<br />
2<br />
Figure 1 clamped-clamped beam type transformer<br />
Where v(t) is the voltage across the PZT, is the<br />
piezoelectric coupling term and YI is the bending stiffness of<br />
the composite cross section given by<br />
<br />
bYs<br />
3<br />
3 3<br />
YI nhc<br />
(1 n)<br />
hb<br />
ha<br />
3<br />
(3)<br />
YP<br />
n =<br />
Y<br />
(4)<br />
Where Y p and Y s is Young’s modulus of PZT and<br />
substructure, h c is the position of the top of PZT layer from<br />
the neutral axis, h b is the position of the bottom of the PZT<br />
layer from the neutral axis, h a is the position of the bottom of<br />
the substructure layer from the neutral axis and the couple<br />
term can be written as<br />
s<br />
Ypd31b<br />
2 2<br />
( hc<br />
hb<br />
)<br />
2h<br />
(5)<br />
Where d 31 is piezoelectric coefficient, where h p is the<br />
thickness of PZT, b is the width of the beam. Because<br />
electrode of driver doesn’t cover entire cantilever but the<br />
region from 0 to x 1 , and electrode pair of sensor part covers<br />
from x 2 to L, then Eq. (2) should be multiplied by<br />
H( x)<br />
H(<br />
x x1)<br />
, where H(x) is the Heaviside function. So<br />
rewrite Eq. (2) as<br />
p<br />
2<br />
w(<br />
x,<br />
t)<br />
M ( x,<br />
t)<br />
YI v(<br />
t)<br />
H<br />
( x)<br />
H ( x x1<br />
) (6)<br />
2<br />
x<br />
Then employing Eq. (6) in Eq. (1), and considering moment<br />
generated from output electrode give<br />
4<br />
∂ w x<br />
YI<br />
4<br />
∂x<br />
d<br />
vin<br />
t<br />
dx<br />
5<br />
2<br />
, t ∂ wx,<br />
t ∂wx,<br />
t ∂ wx,<br />
t<br />
c I<br />
c<br />
4<br />
∂x<br />
∂t<br />
1 <br />
v<br />
dx <br />
x d<br />
x<br />
x <br />
<br />
s<br />
out<br />
t<br />
<br />
<br />
m<br />
∂t<br />
d<br />
dx<br />
Then we assume a solution of Eq. (8) in the form<br />
a<br />
∂t<br />
x<br />
x d<br />
x<br />
L<br />
2<br />
<br />
2<br />
<br />
<br />
0<br />
dx <br />
(7)<br />
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11-13 May 2011, Aix-en-Provence, France<br />
<br />
∑ ∞ 1 t<br />
r<br />
wr<br />
t<br />
<br />
w x,<br />
t <br />
r<br />
x<br />
r<br />
t<br />
<br />
(8) r<br />
t<br />
<br />
<br />
<br />
r1 vin<br />
t vout<br />
t e sin wdr<br />
t d<br />
0<br />
r1<br />
dr<br />
(20)<br />
wherer<br />
xare the normal modes of the system, and <br />
r<br />
t<br />
<br />
2<br />
where dr<br />
r<br />
1<br />
r<br />
is the damped angular frequency.<br />
are modal coordinate of clamped-clamped beam for the rth<br />
In order to obtain the electrical circuit equation, one<br />
mode. Mass-normalized modes can describe undamped free<br />
should consider the following piezoelectric constitutive<br />
vibration as [13]<br />
relation<br />
r<br />
x Ar<br />
cosh<br />
r<br />
x cosr<br />
x <br />
r<br />
sinh<br />
r<br />
x sin r<br />
x<br />
(9)<br />
T<br />
D3 x, t d31T1<br />
33E3<br />
(21)<br />
where r<br />
can be obtained from the characteristic equation<br />
given by<br />
<br />
cos r<br />
<br />
cosh r<br />
1<br />
(10)<br />
L L<br />
and is given by<br />
r<br />
cosh rL<br />
cos rL<br />
<br />
r<br />
(11)<br />
sinh L sin L<br />
The mode-shapes satisfy the orthogonality condition<br />
L<br />
0(<br />
r s)<br />
m s<br />
r<br />
( x)<br />
dx <br />
1(<br />
r s)<br />
x 0<br />
L<br />
4<br />
d <br />
r<br />
( x)<br />
0( r s)<br />
YI s<br />
dx <br />
4 2<br />
dx <br />
( r s)<br />
x 0<br />
r<br />
(12)<br />
2<br />
where r<br />
is the undamped natural frequency for rth mode<br />
given by<br />
2 YI<br />
r r<br />
2<br />
mL<br />
(13)<br />
and by eq. (12) we can obtain<br />
A r<br />
1 mL<br />
(14)<br />
Use eq. (7) in eq. (6) and with orthogonality condition<br />
given by eq. (11) yields the following equation:<br />
2<br />
<br />
r<br />
t<br />
<br />
2<br />
r <br />
rr<br />
t<br />
r<br />
r<br />
t<br />
<br />
r1vin<br />
t<br />
<br />
r<br />
2vout<br />
t<br />
0<br />
(15)<br />
where<br />
dr<br />
x<br />
dr<br />
x<br />
<br />
r1<br />
<br />
<br />
r 2<br />
<br />
(16)<br />
dx<br />
xx<br />
dx<br />
1<br />
xx2<br />
Since 1st normalized modes r<br />
xare symmetric in the x<br />
direction<br />
dr<br />
x<br />
dr<br />
x<br />
<br />
(17)<br />
dx<br />
xx<br />
dx<br />
1 xx2<br />
therefore <br />
r1<br />
equal <br />
r 2<br />
, then Eq. (15) rewrite as<br />
2<br />
<br />
r<br />
t<br />
<br />
2<br />
r <br />
r<br />
r<br />
t<br />
r<br />
r<br />
t<br />
<br />
r1v<br />
in<br />
t<br />
<br />
vout<br />
t<br />
0 (18)<br />
and<br />
csIr<br />
ca<br />
r<br />
<br />
(19)<br />
2YI<br />
2mr<br />
is the equivalent damping term that includes the effect of<br />
strain rate damping and air damping.<br />
The solution of Eq. (19) can be expressed by the Duhamel<br />
integral:<br />
r<br />
r<br />
where D 3 is electrical displacement, and T 1<br />
is axial stress in<br />
term of Young’s modulus of PZT Y p<br />
and axial strain S 1<br />
. E3<br />
is<br />
the electrical field of generator( E<br />
3<br />
vout<br />
( t)<br />
hp<br />
), and <br />
T 33<br />
is the<br />
permittivity at constant stress. It can be replaced by<br />
permittivity at constant strain, as [4]<br />
s T<br />
d31Y<br />
2 33<br />
<br />
33<br />
<br />
p<br />
(22)<br />
so Eq. (22) can be written as<br />
s vout<br />
D3 x, t d31Y<br />
pS1x,<br />
t<br />
<br />
33<br />
(23)<br />
hp<br />
Bending strain for PZT layer is not a constant, it increases<br />
or decreases linearly in polarization direction, y direction.<br />
The average bending strain can be expressed as a function of<br />
distance h pc of the center of the PZT layer (in thickness<br />
direction) to the neutral axis and curvature of the beam.<br />
2<br />
∂ w<br />
<br />
x,<br />
t s vout<br />
t<br />
<br />
D3<br />
x, t d31Y<br />
phpc<br />
<br />
2 33<br />
(24)<br />
∂x<br />
hp<br />
Electrical charge q(t) can be obtained by integrating the<br />
electrical displacement over the electrode area.<br />
2<br />
s<br />
L w<br />
<br />
x,<br />
t<br />
<br />
33<br />
q t D3dA<br />
( d31Yph<br />
pc<br />
vout<br />
t<br />
) bdx<br />
A <br />
<br />
xx<br />
2<br />
2<br />
x<br />
hp<br />
(25)<br />
Current can be obtained by first differential of<br />
electrical charge at time t.<br />
3<br />
s<br />
dq<br />
<br />
t<br />
L wx,<br />
t dvout<br />
t<br />
<br />
33b<br />
i t d31Y<br />
h b dx<br />
( L x2<br />
)<br />
2<br />
dt<br />
p pc<br />
<br />
<br />
x x2<br />
x<br />
t<br />
dt h<br />
(26)<br />
The output voltage across the resistive load is given by<br />
vout<br />
t<br />
Rli(<br />
t)<br />
<br />
<br />
3<br />
s<br />
L wx,<br />
t dv <br />
<br />
out<br />
t <br />
33b<br />
Rl<br />
d31Y<br />
phpcb<br />
dx <br />
( L x2)<br />
<br />
x<br />
x<br />
2<br />
2<br />
<br />
x<br />
t<br />
dt hp<br />
<br />
(27)<br />
The electrical circuit equation can be represented by<br />
s<br />
3<br />
v L<br />
out<br />
t dvout<br />
t b<br />
wx<br />
t<br />
<br />
33 ( L x2<br />
) d Yphpcb<br />
,<br />
31<br />
dx<br />
R<br />
x x<br />
2<br />
l<br />
dt h<br />
2<br />
p<br />
x<br />
t<br />
(28)<br />
Using Eq. (8) in Eq. (28) yields<br />
dv <br />
<br />
<br />
<br />
out<br />
t hp<br />
dr<br />
t<br />
<br />
v<br />
s<br />
out<br />
t (29)<br />
dt Rl<br />
33<br />
b( L x2)<br />
r1<br />
dt<br />
where<br />
p<br />
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11-13 May 2011, Aix-en-Provence, France<br />
<br />
L 2<br />
d31Y<br />
phpchp<br />
( ) d31Y<br />
phpch<br />
The steady-state output voltage eq. (31) can be rewritten as<br />
d x<br />
p dr<br />
x<br />
<br />
r<br />
<br />
dx <br />
s<br />
2<br />
s<br />
33( L x2)<br />
<br />
(30)<br />
<br />
<br />
xx<br />
dx ( )<br />
2<br />
33<br />
L x2<br />
dx<br />
1<br />
j <br />
<br />
<br />
xx<br />
2<br />
<br />
c j t dr<br />
t<br />
voute<br />
<br />
<br />
c <br />
r1<br />
dt<br />
(34)<br />
From Eq. (29), v out<br />
t<br />
yield<br />
then, using Eq. (33) in Eq. (34), the output voltage across the<br />
t<br />
<br />
L t<br />
<br />
dr<br />
<br />
t<br />
<br />
c<br />
<br />
resistive load due to the harmonic excitation can be expressed<br />
c<br />
vout<br />
t e e r<br />
dt c<br />
xx<br />
(31)<br />
2<br />
<br />
r1<br />
dt<br />
as<br />
<br />
<br />
Where τ c is time constant of the circuit given by<br />
jr<br />
r<br />
2 2<br />
s<br />
R<br />
l<br />
33b( L x2)<br />
r1<br />
r<br />
2 j<br />
rr<br />
<br />
c<br />
<br />
(32)<br />
vout<br />
<br />
v<br />
<br />
in<br />
h<br />
jr<br />
<br />
r<br />
1<br />
j<br />
c<br />
p<br />
<br />
<br />
2 2<br />
2 j<br />
<br />
III. Harmonic excitation<br />
In the application of transformer, input voltage is<br />
considered as a harmonic excitation, therefore (i.e.,<br />
t<br />
<br />
jwt<br />
vin<br />
vine<br />
, where v in<br />
, are the amplitudes of input<br />
voltage, ω is driving frequency, and j is the unit imaginary<br />
number). Steady state voltage outputs and beam responses<br />
are obtained. Since the system is linear, the mode shape of<br />
beam and voltage output should be also in the form of<br />
harmonic (i.e., where v out<br />
, and r<br />
are the amplitudes of output<br />
voltage, and modal coordinate of clamped-clamped beam).<br />
Then the modal equation of motion given by [15] can be<br />
reduced to<br />
jwt<br />
v v e<br />
t<br />
<br />
<br />
r in out<br />
r<br />
(33)<br />
2 2<br />
r<br />
2 jrr<br />
=========================================================<br />
The transforming ration around ω 1 is<br />
v<br />
v<br />
out<br />
in<br />
<br />
<br />
r1<br />
c 1 1<br />
2 2<br />
2<br />
2 2<br />
<br />
( 1<br />
2 ) 2 <br />
(<br />
) 2<br />
1<br />
1<br />
c<br />
1<br />
<br />
The phase angle between input and output voltage is simply<br />
2 2<br />
<br />
1<br />
21 1<br />
<br />
c(<br />
11<br />
1<br />
<br />
)<br />
sgn( 11)<br />
tan (<br />
)<br />
2 2<br />
<br />
(1 2 )<br />
2<br />
1<br />
1 c 1<br />
1<br />
r<br />
(35)<br />
If the transformer is excited around the natural<br />
frequency of the rth mode, the main contributions in the<br />
summation signs appearing in Eq. (34) and (35) are from the<br />
rth mode. In most cases, the mode of interest is the<br />
fundamental vibration mode of the transformer r=1.<br />
Therefore, it is a useful practice to consider the beam to be<br />
excited aroundω 1 . The reduced expression for the voltage<br />
across the load can be written as<br />
j<br />
c11<br />
vout<br />
<br />
v<br />
2 2<br />
in<br />
<br />
1<br />
2 jrr 1<br />
j<br />
c<br />
<br />
j<br />
c11<br />
(36)<br />
where sgn() is the signum function. Output power P can be expressed as v 2 out<br />
Rl<br />
given by<br />
2<br />
2<br />
vout<br />
( c11v<br />
in<br />
) Rl<br />
P <br />
(39)<br />
2 2<br />
2<br />
R<br />
2 2<br />
2<br />
l 1<br />
<br />
(1 21<br />
c1)<br />
21 1<br />
<br />
c(<br />
11<br />
1<br />
<br />
)<br />
IV. Parametric case study<br />
However, Erturk et al. [11] suggested that one can always use<br />
modal damping ratio ( <br />
4.1 Effect of the electrode length on output voltage<br />
r<br />
) obtained experimentally directly.<br />
1 0.01is used in this study.<br />
In this section, we analyze the transformer by proposed<br />
Table 1 Material, Geometric, and electromechanical parameters of the model<br />
analytical model. The geometric, material, and<br />
===========================================<br />
electromechancial parameters of the transformer are used in Length of the beam, L (mm) 100<br />
Table 1. The input of the transformer is due to the harmonic Width of the beam, b (mm) 10<br />
Thickness of the PZT, h<br />
p<br />
(mm) 0.25<br />
excitation (100V) at first resonance frequency. Steady state Thickness of the substrate, h<br />
s<br />
(mm) 0.5<br />
response of system is interested in. Before presenting the Mass density of the PZT, ρ<br />
p<br />
(kg/m 3 ) 7800<br />
Mass density of the substrate, ρ<br />
s<br />
(kg/m<br />
resulting voltage output and discussing the respective trend,<br />
3 ) 2300<br />
Young’s modulus of the PZT, Y<br />
p<br />
(GPa) 7<br />
mechanical damping coefficient have to be evaluated. With Young’s modulus of the substrate, Y<br />
s<br />
(GPa) 74<br />
the form of the differential equation given by Eq. (1), two Piezoelectric constant, d<br />
31<br />
(pm/V) -210<br />
s<br />
Permittivity, ε 33<br />
(nF/m) 15.3<br />
separate damping terms for the internal damping coefficient ===========================================<br />
(C s I) and the external viscous damping coefficient (Ca) are Figure 2 shows that voltage output for different lengths of<br />
assumed. C s I is assumed to be stiffness proportional, and C a output electrode pair and the attention is given to the first<br />
is assumed to be mass proportional. C s I/YI is equal to 1.243× resonance mode. The output electrode pair covers the region<br />
10 -5 s/rad and Ca/m is equal to 4.886 rad/s in Ref. [11].<br />
1<br />
c<br />
1<br />
1<br />
1<br />
r<br />
r<br />
c<br />
(37)<br />
(38)<br />
78
etween x 2 to L (remember that x 2 is measured from the left<br />
side of clamed end of the beam to electrode pair of sensor<br />
part, and L is the length of beam).From figure 2, the electrode<br />
pair can be used to covers from 0.77L to L and best output<br />
voltage can be obtained from the first vibration mode. To<br />
explain this, normalized mode shape for first vibration mode<br />
is shown in Figure 3. Strain for piezoelectric film is<br />
proportional to the curvature of beam. The strain distribution<br />
between 0.77L to L and 0.22L to 0.77L will cancel each other.<br />
Best output voltage will be obtained when the length of<br />
output electrode pair is 0.22L.<br />
Voltage (mv)<br />
5<br />
4.5<br />
4<br />
3.5<br />
3<br />
2.5<br />
2<br />
1.5<br />
1<br />
0.5<br />
0<br />
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35<br />
Nondimensional electrode length, x/L<br />
Fig. 2 Effect of nondimensional electrode length on output voltage<br />
Mass normalized mode<br />
shape<br />
2.00<br />
1.00<br />
-<br />
-1.00<br />
-2.00<br />
output voltage<br />
0 0.2 0.4 0.6 0.8 1<br />
Fig.3 The mode shape and curvature of cantilever<br />
4.2 Effect of cantilever length on output voltage and<br />
frequency<br />
deflection<br />
curvature<br />
Nondimensional beam coordinate, x/L<br />
Figure 4 shows that voltage output and resonance<br />
frequency for different lengths of cantilever lengths and the<br />
transformer is also excited at first vibration mode. In this<br />
study, material, geometric, and electromechanical parameters<br />
of transformers are shown in Table 1 but cantilever lengths<br />
vary from 5cm to 100cm. From figure 4, output is<br />
proportional to the square of cantilever length and 1 st<br />
resonance frequency is inversely proportional to the square of<br />
cantilever length. Longer beam exceeds larger output but<br />
lower excitation frequency.<br />
4.3 Effect of the thickness on output voltage<br />
Figure 5 shows that output for different thickness ratios<br />
between PZT and substrate when the transformer is excited at<br />
first resonance. In this study, material, geometric, and<br />
electromechanical parameters of transformers are also shown<br />
in Table 1 but thicknesses of substrate vary from 0.125mm to<br />
1mm. The best output can be obtained when the thickness<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
ratio between substrate and PZT is close to 2. Thickness ratio<br />
between substrate and PZT will change the position of<br />
neutral axis of beam. Figure 6(a) and 6(b) shows that neutral<br />
axis lies within substrate and within PZT, respectively. To see<br />
the stress distribution, tensile and compress stresses will<br />
cancel each other when neutral axis lies with PZT layer. We<br />
might think the best output will be obtained when the neutral<br />
axis is located in the contact plane between substrate and<br />
PZT. However, the maximum output voltage occurs when<br />
neutral axis lies within substrate (Fig. 6(a)).<br />
In this study, the thickness of PZT is 0.25 mm and<br />
thickness of substrate varies. When the thickness of substrate<br />
become larger, the distance from neutral axis to top of the<br />
beam will increase but bending stiffness will also increase.<br />
Larger bending stiffness will cause smaller deflection and of<br />
beam when actuator part is excited at the same input voltage.<br />
In the other word, maximum normal stress of beam will<br />
decreases when bending stiffness increases. The output<br />
voltage will be proportional to the summation of normal<br />
stress of beam. The maximum output voltage will occurs<br />
when neutral axis lies within substrate but not in the contact<br />
plane between substrate and PZT. Figure 7 shows that output<br />
voltages for different thickness ratio between PZT and<br />
substrate when choosing different substrate. Young’s<br />
modulus ratio between PZT and substrate varies from 1 to 2.8.<br />
Different Young’s modulus ratio leads to different thickness<br />
ratio to achieve best output of transformer. When the Young’s<br />
modulus ratios are 1 and 2.8, the best output are 0.078V and<br />
0.090V, respectively. The results are summarized in table 2.<br />
frequency (Hz)<br />
500<br />
450<br />
400<br />
350<br />
300<br />
250<br />
200<br />
150<br />
100<br />
50<br />
0<br />
5 20 35 50 65 80 95<br />
cantilever length (cm)<br />
Fig. 4 voltage output and nature frequency with different cantilever lengths<br />
Voltage (mV)<br />
80<br />
70<br />
60<br />
50<br />
40<br />
30<br />
20<br />
free vibration frequency<br />
out put voltage<br />
0.5 1 1.5 2 2.5 3 3.5 4<br />
thinkness ratio, substrate/PZT<br />
Fig. 5 output voltage vs. thickness ratio between substrate and PZT layers<br />
5<br />
4.5<br />
4<br />
3.5<br />
3<br />
2.5<br />
2<br />
1.5<br />
1<br />
0.5<br />
0<br />
voltage (V)<br />
output voltage<br />
79
Fig. 6<br />
Voltage (mv)<br />
90<br />
80<br />
70<br />
60<br />
50<br />
40<br />
30<br />
20<br />
10<br />
0<br />
PZT<br />
Subst<br />
rate<br />
Stress<br />
Neutral axis<br />
PZT<br />
subst<br />
rate<br />
(a)<br />
(b)<br />
Neutral axis lies (a) within substrate (b) within PZT<br />
0.5 1 1.5 2 2.5 3 3.5 4 4.5<br />
thinkess ratio, substrate/PZT layers<br />
Fig. 7 output voltage vs. thickness ratio when different Young’s module ratio<br />
Table 2 best output, thickness ratio and bending stiffness in different substrate<br />
Young’s modulus ratio 0.5 1.0 1.6 2.8<br />
Best thickness ratio 3.61 2.11 1.40 0.98<br />
Bending stiffness 0.063 0.026 0.017 0.012<br />
Best Output (V) 0.071 0.078 0.084 0.090<br />
V. Conclusion<br />
Young's modulus ratio 2.8<br />
Young's modulus ratio 1.6<br />
Young's modulus ratio 1<br />
In this paper, an analytical solution of a fixed-fixed beam<br />
type piezoelectric transformer with Euler-Bernoulli beam<br />
assumption is proposed. Mechanical response, voltage,<br />
current, and power outputs of piezoelectrical transformers are<br />
presented. The model is then used for parameter study. From<br />
analytical model, output voltage depends on the lengthes of<br />
two electrodes, the length of beam, and the Young’s modulus<br />
ratio and thickness ratio between PZT layer and substrate.<br />
The lengthes of input electrodes and output electrodes should<br />
be 0.22L to achieve the largest output when the transformer<br />
is excited at first resonance frequency. The output voltages<br />
and the resonance frequnecies of transformers are<br />
proportional and inversely proportional to the lengthes of<br />
beams, respectively. The combination of Young’s modului<br />
and thicknesses of PZT layer and substrate change the<br />
position of netual axis and the bending stiffness of beam,<br />
concurrently. However, output voltages of transfomers<br />
depend not only on the postion of neutral axes but also on<br />
bending stiffnesses.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
References<br />
[1] C. A. Rosen, “Ceramic transformers and Filters,”<br />
Proceeding of electronic Components Symposium,<br />
Washington, D. C., May 1-3 (1956) 205-211<br />
Stress<br />
[2] Y.-H. Hsu, C.-K. Lee, W.-H. Hsiao, “Optimizing<br />
piezoelectric transformer for maximum power transfer,”<br />
Smart Material and Structure, 12(2003) 373-83<br />
[3] R.-L. Lin, “Piezoelectric transformer characterization<br />
and application of electronic ballast,” PhD Dissertation,<br />
Virginia Polytechnic Institute and state University, USA,<br />
2001<br />
[4] E. M. Baker, W. Huang, D. Y. Chen, F. C. Lee, “Radial<br />
mode piezoelectric transformer design for fluorescent lamp<br />
ballast application,” 33 rd Annual IEEE Power Electronics<br />
Specialists, Cairns, Queensland, Australia, June 23-27 (2002)<br />
1289-94<br />
[5] J.-M. Seo, H.-W. Joo, H.-K. Jung, “Optimal design of<br />
piezoelectric transformer for high efficiency and high power<br />
density,” Sensors and Actuators A, 121 (2005) 520-526<br />
[6] M. C. Do, H. Guldner, “High output voltage DC/DC<br />
converter based on parallel connection of piezoelectric<br />
transformers,” International Symposium on Power<br />
electronics, Electrical Drives, Automation and Motion,<br />
Taormina, Italy, May 23-26, (2006) S18<br />
[7] T. Inoue, S. Hamamura, M. Yamamoto, A. Ochi, Y.<br />
Sasaki, “AC-DC converter based on parallel drive of two<br />
piezoelectric transformer,” Japanese Journal of Applied<br />
Physics, 47 (2008) 4011-4014<br />
[8] S. Roundy, P. K. Wright, J. M. Rabaey, “A Study of<br />
Low Level Vibrations as a Power Source for Wireless Sensor<br />
Nodes,” Computer Communications, 26 (2003) 1131-1144<br />
[9] N. E. duToit, B. L. Wardle, S.-G. Kim, “Design<br />
considerations for MEMS-scale piezoelectric mechanical<br />
vibration energy harvesters,” Integrated Ferroelectrics, 71<br />
(2005) 121-160<br />
[10] S. T. Ho, “Electromechanical Model of a Longitudinal<br />
Mode piezoelectric Transformer,” The Seventh International<br />
Conference on Power Electronics and Drive Systems,<br />
Bangkok, Thailand, November 27-30, (2007) 267-272<br />
[11] A. Erturk, D. J. Inman, “A distributed parameter<br />
electromechanical model for cantilevered piezoelectric<br />
energy harvesters,” Journal of Vibration and Acoustics, 130<br />
(2008) 041002<br />
[12] A. Erturk, and D.J. Inman, “On mechanical modeling<br />
of cantilevered piezoelectric vibration energy harvesters,”<br />
Journal of Intelligent Material Systems and Structures, 19<br />
(2008) 1311-1325<br />
[13] T. K. Caughey, and M. E. J. O’Kelly, “Classical normal<br />
modes in damped linear dynamic systems,” Journal of<br />
Applied Mechanics, 32 (1965) 583–588.<br />
80
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Study of Black Silicon Obtained by Deep Reactive<br />
Ion Etching – Approach to Achieving the Hot Spot of<br />
a Thermoelectric Energy Harvester<br />
K.N Nguyen 1 , D.Abi-Saab 1 , M. Malak 1 , P. Basset 1 , E. Richalot 2 , N. Pavy 1 , F. Flourens 1 , F. Marty 1 , D. Angelescu 1 ,<br />
Y. Leprince-Wang 3 , T.Bourouina 1<br />
1 Université Paris-Est, ESYCOM, ESIEE Paris, 2 Bd. Blaise Pascal, 93162, Noisy-le-Grand, France<br />
2 Université Paris-Est, ESYCOM, UPEMLV, 2 Bd. Blaise Pascal, 93162, Noisy-le-Grand, France<br />
3 Université Paris-Est, LPMDI, 2 Bd. Descartes, F-77454, Marne-la-Vallée Cedex 2, France<br />
Abstract- In this paper we study the enhanced absorption<br />
properties of micro/nano structured silicon surface under<br />
incident electromagnetic (EM) illumination and then its<br />
capacity to convert light to heat. We then simulate the optical<br />
reflectance of the 3D micro/nano silicon cones of different<br />
dimensions. Equipped with the favorable simulation results we<br />
fabricate black silicon with excellent anti-reflectivity by using<br />
deep reactive ion etching (DRIE) under cryogenic<br />
temperatures. Reflectance measurement with an integrating<br />
sphere is approximately 1% in the optical wavelength range.<br />
Following this, black silicon with integrated resistance<br />
temperature detector (RTD) is developed to investigate its<br />
efficiency of the photo-thermal conversion.<br />
I. INTRODUCTION<br />
Research in the area of electromagnetic energy harvesting<br />
has been done over the past decade. In the optical<br />
wavelength range, one can use photovoltaic conversion as<br />
well as photo-thermal conversion, both of whose<br />
efficiencies depend on material properties. Either<br />
mechanism is possible on silicon, with various efficiencies<br />
depending on level of doping and wavelength of incident<br />
light. Also, microstructuring the surface of silicon can lead<br />
to noticeable enhancement of conversion efficiency.<br />
In this paper we study the photo-thermal conversion<br />
behavior of black silicon obtained by cryogenic DRIE, with<br />
the prospect of producing a hot spot intended to integrate a<br />
thermoelectric energy harvester consisting of a vertical<br />
superlattice [1]. In order to fuel such thermoelectric<br />
elements by solar radiation, the optimization of the hot spot<br />
is crucial. To this end, we propose in this paper to develop a<br />
light-absorbing layer with extremely low reflectivity so as<br />
to maximize heating of the hot spot under the effect of<br />
electromagnetic (EM) radiation in the visible and nearinfrared<br />
ranges. The hot spot is made of black silicon, a<br />
material consisting of dense (sub)-micrometer cones which<br />
lead to multiple reflections of incident photons and hence to<br />
light trapping and absorption. An air cavity etched on the<br />
back side thermally insulates the hot spot, which is heated<br />
by incident light focused by a microlens (Fig. 1).<br />
Fig. 1. Sketch of the target material.<br />
The natural reflectance of a flat silicon/air interface is<br />
around 30% because of the high refractive index of silicon.<br />
Surface texturing is an effective technique to reduce the<br />
reflectance of the Si surface. Several techniques have been<br />
previously studied for forming different profiles such as wet<br />
etching [2], femtosecond laser pulse [3], reactive ion<br />
etching [4] and deep reactive ion etching (DRIE) [5,6].<br />
Maskless texturing of a polished silicon wafer can result in<br />
the formation of “grass”-like structures that appears black to<br />
the human eye, hence the name “black silicon” [4,5]. DRIE<br />
texturing is a well-known technique to obtain black silicon<br />
surfaces of low reflectance, in a controlled and repeatable<br />
manner. The DRIE is based on inductively-coupled plasma<br />
(ICP) of sulphur hexafluoride (SF 6 ) and allows anisotropic<br />
etching of silicon by taking advantage of a passivation<br />
mechanism in the side walls. While the Bosch technique is<br />
the most widely used process in DRIE, we have in this work<br />
used the cryogenic process to fabricate black silicon.<br />
II. FEM SIMULATIONS AND RESULTS<br />
The influence of process parameters on the black silicon<br />
structure has been actively studied [5] but no in depth study<br />
concerning the impact of its 3D geometry on the reflectivity<br />
of the incident EM radiation has been done. The simulations<br />
of the reflectance of a model black silicon surface hence are<br />
performed with Ansoft’s HFSS (High Frequency Structural<br />
Simulator) based on the Finite Element Method. Although<br />
81
Fig. 2. a) Diagram of the structure simulated by HFSS TM b) 3D sketch of<br />
the simulated surface.<br />
black silicon can be built in various shapes such as spikes,<br />
“penguin-like” structures, columns and pyramids, the<br />
simulations are performed with cones since it is one of the<br />
shapes that provides better absorption [7]. In this paper, we<br />
focus on textured surfaces formed by cones of dimensions<br />
(height and width) varying between 150 nm and 5 µm under<br />
different directions of the incident field.<br />
A. Description of the 3D surface<br />
The simulated structure consists of a silicon substrate on<br />
which identical cones are periodically repeated along the x-<br />
and y- axis. The structure is defined by its out-of-plane<br />
height (h), base diameter (d), and in-plane periodicity (p), as<br />
represented in Fig. 2. According to Floquet’s theorem [8],<br />
the structure periodicity induces the field pseudo-periodicity<br />
and allows us to reduce the computation time by restricting<br />
it to a single lattice unit with biperiodic boundary conditions<br />
[9]. The surface is excited by an incident monochromatic<br />
plane wave with a wavelength tuned from 430 nm to 1000<br />
nm. The reflectance is obtained by calculating the ratio<br />
between the reflected and incident energies passing through<br />
a surface S. The surface S has the same dimensions as the<br />
elementary cell and is placed above the simulated cone and<br />
parallel to the periodicity plane.<br />
B. Results<br />
1) Influence of the height of cones<br />
To evaluate the influence of cone height on the<br />
reflectivity of black silicon, we have simulated a structure<br />
with identical cones with varying height. The periodicity<br />
and the width of cones are of 1.5 µm while the height is<br />
varied from 3.5 µm to 5 µm. This structure is illuminated in<br />
a direction normal to the substrate. The results shown in<br />
Fig. 3 clearly indicate that the reflectivity in the visible light<br />
spectrum decreases uniformly while increasing cone height<br />
because of multireflection of the incident light on the silicon<br />
3D surface. We can notice that while the dependence of<br />
reflectivity on height is modest in the infrared limit (1000<br />
nm), it is significant when observed at lower wavelengths<br />
(430-600 nm).<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
shown in Fig. 4. It appears that at constant periodicity the<br />
reflectivity decreases with increasing cone diameter while<br />
the diameter is lower than the periodicity. The curve slope<br />
decreases from the point where the bases of the structure are<br />
in contact. The very large reflectivity reduction before this<br />
point can be understood by the large reduction of the planar<br />
surface between the cones. Then the cone bases start to<br />
overlap. Before the planar surface disappears completely,<br />
the reflectivity starts to increase slightly, which can then be<br />
explained by the reduction of the angle between the incident<br />
field and the normal of the cone lateral surface, and by the<br />
induced reduction of the cone aspect ratio. We observe that<br />
the lowest reflectivity is obtained for a cone diameter<br />
approximately 30% larger than the structure periodicity.<br />
Decreasing the cone periodicity is also recommended for<br />
lower reflectivity.<br />
Fig. 3. Cone height influence on simulated reflectivity.<br />
Fig. 4. Cone width influence on simulated reflectivity.<br />
2) Influence of the diameter of cones<br />
The impact of the cone diameter on the reflectivity is<br />
studied by simulating micrometer size cones of constant<br />
periodicity (p = 1.5 µm) and height (h = 3.5 µm), whose<br />
diameter is varied from 1 to 2.08 µm. A normal incidence is<br />
considered. The simulated reflectivity of these structures is<br />
Fig. 5. Example of simulated reflectivity with respect to the electric field<br />
incident angle.<br />
82
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
3) Influence of the incident electric field angle<br />
Since promising results on the physical parameters of<br />
silicon conical structures for the lowest reflectivity are<br />
found as above, textured surfaces consisting of micrometer<br />
and sub-micrometer cones with high steepness and high<br />
density are simulated to study the variation of the<br />
reflectivity with respect to the incident field direction. θ i is<br />
the incident field angle from the normal incidence.<br />
Simulations are performed with the variation of θ i from 0°<br />
to 85° on a textured silicon surface (Fig. 5). The periodicity,<br />
diameter and height of sub-micrometer cones are 150 nm,<br />
190 nm and 910 nm respectively. As shown in Fig. 5, high<br />
density cones exhibit a low reflectivity in the visible range<br />
for incidence angles up to 50° from the normal of the<br />
surface. A similar effect is observed for the micrometer size<br />
structures whose dimensions are presented in Fig. 4.<br />
Fig. 6. SEM image of Black Si obtained by a cryogenic DRIE process.<br />
III.<br />
EXPERIMENTS AND RESULTS<br />
A. Fabrication<br />
Based on the simulation results presented above, we<br />
performed an experimental study of black silicon<br />
fabrication and characterization. The black silicon was<br />
obtained using O 2 -SF 6 cryogenic deep reactive ion etching<br />
on 4-inch polished (100) silicon wafers. By varying the<br />
process parameters such as bias voltage, temperature, gas<br />
pressure, and RF power, we can obtain various structure<br />
geometries [10]. The wafers were subjected to DRIE at<br />
cryogenic temperatures without any mask. SF 6 gas<br />
generates F * radicals for chemical etching of silicon leading<br />
to volatile SiF 4 whereas O 2 gases produce O * radicals for<br />
silicon sidewall passivation with Si x O y F z . Such wafers were<br />
treated under different plasma conditions in order to obtain<br />
different textures of black silicon. Guided by the simulation<br />
results, we attempt to obtain the best compromise between<br />
density, periodicity and aspect ratio of the silicon cones.<br />
Black silicon with high-density of (sub)-micrometer cones<br />
(Fig. 6) was achieved. In order to investigate the photothermal<br />
conversion of the black silicon, we have designed<br />
and fabricated a thin film platinum resistance temperature<br />
detector (Pt RTD) at micrometer scale to measure its<br />
temperature change under exposure to various intensities of<br />
light (Fig. 7). Platinum was used because of its stability,<br />
precision and its linear relationship between resistance and<br />
temperature. A four probe resistance measurement was<br />
performed to eliminate contact resistances and increase the<br />
accuracy of the measurement.<br />
B. Characterization<br />
The surface morphologies of black silicon were<br />
investigated by scanning electron microscopy (SEM). The<br />
cones have a width of about 350 nm and a height of about<br />
1.4 µm, which were directly extracted from side-view SEM<br />
image (Fig. 6). The average periodicity of the structure is<br />
570 nm, calculated from the azimuthally averaged intensity<br />
of the Fourier Transform (FT) of a top-view image (see<br />
Fig.8).<br />
Fig. 7. SEM image of Pt RTD surrounded by Black Si.<br />
Fig. 8. Top-view SEM image of Black Si and azimuthal average of the<br />
Fourier Transform of the top view image.<br />
The reflectance of black silicon has been measured for<br />
wavelengths between 400 nm and 1000 nm by a<br />
spectrometer (Maya 2000 Pro from Ocean Optic) with an<br />
integrating sphere coupled to a halogen light source. NIST<br />
reflectivity standards were used for calibration. Fig. 9 plots<br />
the measured reflectance from planar and DRIE-textured<br />
surface under normal incidence. Our black silicon is found<br />
to exhibit a reflectance of ~1% in the visible range without<br />
anti-reflection films.<br />
83
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
Guided by the favorable simulation results, conical black<br />
silicon wafer was fabricated by DRIE under cryogenic<br />
temperatures with diameter of 350 nm, height of 1.4 µm and<br />
periodicity of 570 nm. The cones are fabricated in a<br />
collective manner over the whole wafer area. This structure<br />
presents excellent antireflective behavior over the 400 – 950<br />
nm spectral range with a reflectance ~ 1% in the visible<br />
range. This reflectance level is among the best published in<br />
the literature for plasma-etched black silicon. We have<br />
successfully measured the variation in the photo-thermal<br />
effects of black silicon with varying incident light<br />
intensities albeit notably without thermal insulation<br />
substrates. Current work in progress consists of improving<br />
the prototype for the light-thermal conversion and finally<br />
Fig. 9. Measured reflectance spectra of black silicon under normal<br />
incidence.<br />
thermoelectric conversion.<br />
Relative resistance variation (ppm)<br />
8000<br />
6000<br />
4000<br />
2000<br />
0<br />
0 0,5 1 1,5 2<br />
Incident light intensity (mW/mm²)<br />
Fig. 10. Resistance variation of Pt RTD under different incident light<br />
intensities.<br />
The device was tested by irradiating it with different<br />
intensities of visible light coming from a halogen light<br />
source. The resistance variation (Fig. 10) ΔR/R is nearly<br />
7000 ppm for an incident light intensity of 1.6 mW/mm 2<br />
equivalent to a temperature increase of nearly 2°C and to a<br />
photo-thermal conversion of 1250°C/(W/mm 2 ) for this<br />
device. It is worth mentioning that this first trial was<br />
performed on a device suspended on a membrane but with<br />
no additional thermal insulation from the thermally<br />
conductive substrate.<br />
IV. CONCLUSION<br />
We have simulated the optical reflectance of the 3D black<br />
silicon structures consisting of cones of sub-micro and<br />
micrometer dimensions with different heights and diameters<br />
for the optical wavelengths. We observed that a black<br />
silicon structure with the sharpest and high density cones is<br />
expected to obtain the lowest reflectivity. It is obtained<br />
when the cones diameter are about 30% larger than the<br />
periodicity. Besides, the influence of the direction of the<br />
incident field on the reflection of black silicon cannot be<br />
neglected. It is shown that angle of the incidence from the<br />
normal surface has almost no influence up to 40° on a low<br />
reflective surface.<br />
ACKNOWLEDGMENT<br />
The authors would like to thank to EADS Foundation by<br />
whom this work is funded through the project TESEER.<br />
REFERENCES<br />
[1] J. Parasuraman, M. Bardoux, D. Angelescu, P. Basset, T.<br />
Bourouina and P. Chantrenne, "Development of vertical<br />
superlattices in silicon for on-chip thermal management",<br />
Proceeding of the 16th International workshop on Thermal<br />
investigations of ICs and Systems, Barcelona (THERMINIC'10),<br />
Barcelona, Spain, 2010.<br />
[2] Howard M. Branz, Vernon E. Yost, Scott Ward, Kim M. Jones,<br />
Bobby To and Paul Stradinset, “Nanostructured black silicon and<br />
the optical reflectance of graded-density surfaces”, Applied<br />
Physics Letters, vol. 94, N° 23, 2009.<br />
[3] M. Y. Shen, C. H. Crouch, J. E. Carey and E. Mazur,<br />
“ Femtosecond laser-induced formation of submicrometer spikes<br />
on silicon in water”, Applied Physics Letters, vol. 85, N°. 23, 6<br />
December 2004.<br />
[4] B.M. Damiani, R. Lüdemann, D.S. Ruby, S.H. Zaidi, A. Rohatgi,<br />
“Development of RIE-textured silicon solar cells”, Photovoltaic<br />
Specialists Conference, 2000. Conference Record of the Twenty-<br />
Eighth IEEE, 15-22 Sept. 2000, pp.371–374.<br />
[5] Henri Jansen, Meint de Boer, Henk Wensink, Ben Kloeck, Miko<br />
Elwenspoek, “The black silicon method. VIII. A study of the<br />
performance of etching silicon using SF6/O2-based chemistry with<br />
cryogenical wafer cooling and a high density ICP source,”<br />
Microelectronics Journal, vol. 32, pp. 769-777, 2001.<br />
[6] F. Marty , L. Rousseau, B. Saadany, B. Mercier, O. Francais, Y.<br />
Mita, T. Bourouina “Advanced etching of silicon based on deep<br />
reactive ion etching for silicon high aspect ratio microstructures<br />
and three-dimensional micro- and nanostructures”,<br />
Microelectronics Journal, Vol. 36,issue 7, Jul. 2005 pp. 673-677.<br />
[7] M.Halbwax et al, “Micro and nano-structuration of silicon by<br />
femtosecond laser: Application to silicon photovoltaic cells<br />
fabrication”, Thin solid films, Vol. 516, issue 20, pp. 6791-6795<br />
(2008).<br />
[8] R. Petit,Ed., “Electromagnetic Theory of Gratings”, Berlin,<br />
Germany : Springer-Verlag,1990.<br />
[9] E. Richalot, M. Bonilla, M.-F. Wong, V. Fouad-Hanna, H.<br />
Baudrand, J. Wiart, “Electromagnetic Propagation into Reinforced<br />
-Concrete Walls”, IEEE Trans. Microwave Theory Tech., Vol. 48,<br />
No. 3, March 2000, pp.357-366.<br />
[10] R. Dussart et al, “Silicon columnar microstructures induced by an<br />
SF6/O2 plasma,” Journal of Physics D: Applied Physics, 38, 3395,<br />
2005.<br />
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<br />
Performance Evaluation of MEMS Piezoelectric<br />
Inertial Energy Generator<br />
Aliza Aini Md Ralib 1 , Anis Nurashikin Nordin 1 , Raihan Othman 2 , Hanim Salleh 3<br />
1 Department of Electrical and Computer Engineering, International Islamic University Malaysia<br />
2<br />
Department of Science in Engineering, International Islamic University Malaysia<br />
3 Department of Mechanical Engineering, Universiti Tenaga Nasional Malaysia<br />
Abstract- Vibration based inertial energy generators have<br />
become significantly popular due to the growing demand of<br />
wireless sensor networks which need miniature, portable,<br />
long lasting and easily recharged sources of power. Usage of<br />
hazardous batteries is an unacceptable solution to power up<br />
the densely populated nodes due to their bulky sizes and high<br />
battery replacement cost. As such, the viability of ‘green’<br />
microelectromechanical (MEMS) vibration based inertial<br />
energy generator has become even more dominant. This<br />
paper reports the design and simulation of a cantilever<br />
piezoelectric inertial energy generator based on bulk silicon<br />
micromachining for wireless condition monitoring in power<br />
plants. Power plants generate ambient vibrations in the low<br />
kHz range which can be harvested to power the wireless<br />
condition monitoring circuits. Output power of the system<br />
will be enhanced when it is operated at the ambient resonance<br />
frequency. This paper discusses the effect of various lengths,<br />
shapes and volume of the cantilever beam, to its natural<br />
resonant frequency. The effect of different piezoelectric<br />
material with the maximum output power produced is also<br />
highlighted. The design and finite element modeling was<br />
conducted using MEM PZE module in Coventorware TM .<br />
I. INTRODUCTION<br />
Conventional battery power sources have large<br />
maintenance and short lifetimes, making it unsuitable for<br />
applications in low power wireless sensor nodes. Energy<br />
harvesting devices capable of converting wasted ambient<br />
energy to useful electrical power could be one of the<br />
favorable solutions. Ambient mechanical vibration is the<br />
most versatile renewable energy source because it has<br />
infinite lifetimes and eliminates the disposal issues of<br />
waste batteries.<br />
Inertial energy generators harvest the energy from<br />
motion by damping the internal motion of a proof mass<br />
suspended within the device when the device is vibrating.<br />
The electro mechanical conversion is done by a transducer<br />
typically electromagnetic, electrostatic or piezoelectric.<br />
Among these methods, piezoelectric transducers have the<br />
simplest configuration and highest efficiencies [1]. This<br />
paper emphasizes on the design and simulation of a<br />
MEMS piezoelectric inertial generator based on bulk<br />
silicon micromachining to power up wireless condition<br />
monitoring circuits at power plants. Low ambient resonant<br />
frequencies are needed to acquire optimize harvested<br />
output power. This paper is organized as follows. Section II<br />
explains the proposed device structure. Section III presents<br />
the simulation analysis, results and discussion. The<br />
simulation results are divided into three main sections that<br />
discuss the effect of various lengths, shapes and volume of<br />
the cantilever beam to its natural resonant frequency, the<br />
effect of resonant frequency to the output power and the<br />
effect of the different piezoelectric thin film material to the<br />
maximum output power produced respectively. Finally, the<br />
conclusion is given in Section IV.<br />
II. DEVICE STRUCTURE<br />
Fig. 1 shows the MEMS piezoelectric cantilever inertial<br />
generator device operating in transversal mode (d 31 mode) for<br />
low frequency vibration. The miniature device consists of a<br />
single cantilever structure that comprised of silicon substrate,<br />
a piezoelectric layer and two layers of top and bottom<br />
electrodes. A silicon proof mass is fabricated at the free end<br />
to adjust the resonant frequency. The piezoelectric energy<br />
conversion can be described using the equivalent linear<br />
spring mass system as shown in Fig. 2 where z=x-y is the net<br />
displacement, K eq is the equivalent spring constant , C is the<br />
damping coefficient and M eq is the equivalent lumped mass<br />
[2].<br />
Fig. 1. 3D view of MEMS piezoelectric inertial generator operated in<br />
d31 mode<br />
Fig. 2. Schematic representation of the cantilever beam fixed at one end<br />
and the equivalent first order model of a resonant system<br />
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ω =<br />
K<br />
n<br />
eq<br />
K<br />
M<br />
3EI<br />
=<br />
L<br />
eq<br />
eq<br />
3<br />
beam<br />
(1)<br />
(2)<br />
Fig. 3. Side and Top view of MEMS Piezoelectric Inertial Generator<br />
TABLE I<br />
DEVICE DIMENSIONS<br />
Material Length<br />
(m)<br />
Width<br />
left (m)<br />
Width<br />
right (m)<br />
Thickness<br />
(m)<br />
Silicon 5039 672 2076 500<br />
Substrate<br />
Aluminium 5039 672 2076 5<br />
Zinc Oxide 5039 672 2076 10<br />
Aluminium 5039 672 2076 5<br />
Silicon Proof<br />
Mass<br />
1039 2076 2076 300<br />
The device dimensions, side and top view of the MEMS<br />
piezoelectric inertial generator are shown in Table I and Fig.<br />
3 respectively. When the cantilever is excited due to ambient<br />
vibration, it will induce the mechanical stress in the<br />
piezoelectric layer. The stress induced strain in the zinc oxide<br />
piezoelectric crystal and consequently generates surface<br />
charges due to unbalanced ions. The charges generated are<br />
proportional to output voltage produced and is extracted<br />
through metallization of top and bottom layer of aluminium<br />
electrodes that located in between of the piezoelectric layer.<br />
Equations (1) to (4) show the relation between the volume<br />
increments of the cantilever beam to its resonant frequency.<br />
If we assume that the micro generator operates at resonance<br />
frequency, the resonant frequency of spring mass system, n<br />
is shown in (1) where M eq is the equivalent mass and K eq is<br />
the equivalent stiffness. The equivalent stiffness for the<br />
cantilever beam can be calculated as shown in (2) where E is<br />
Young Modulus (GPa), I is the moment of inertia, and L beam<br />
is the length of the beam, W beam , and H beam are width and<br />
height of the beam respectively. The equivalent mass is<br />
directly related to the length, width and height of the<br />
cantilever beam. The higher the volume of the cantilever<br />
beam, the lower the resonant frequency applied at the<br />
proposed design as shown in (1). At the resonance<br />
frequency, the cantilever vibrates at maximum deflection and<br />
consequently maximum potential difference will be produced<br />
which directly impact the output power produced [2].<br />
33<br />
Meq = Mm + Mbeam<br />
140<br />
M = ρ * W * H * L<br />
beam beam beam beam<br />
III.<br />
SIMULATION ANALYSIS<br />
The prototype device was modeled using finite<br />
element simulator (FEM), MEM PZE Coventorware that<br />
provides automated design solutions for MEMS devices.<br />
Designer module was used to define the virtual fabrication<br />
steps, draw the layout and generate three dimensional<br />
model of the prototype device [3]. The structure consists<br />
of Silicon / Aluminium / Zinc Oxide / Aluminium / Silicon<br />
proof mass. The detailed of the fabrications steps defined<br />
in the designer module as shown in Table II. Silicon bulk<br />
micromachining is applied in the fabrication steps to<br />
design the cantilever beam. Undoped silicon wafer<br />
with thickness of 500 m was used as the substrate.<br />
Aluminium was chosen as top and bottom electrode with<br />
thickness of 5 m each to measure the output voltage<br />
produced. Zinc oxide was chosen as the piezoelectric layer<br />
because it requires relatively low deposition temperature,<br />
has high piezoelectric coupling coefficient and excellent<br />
bonding to substrate materials such as silicon [4]. The<br />
proposed device utilized back-etched silicon proof mass in<br />
order to simplify the fabrication process of the devices. A<br />
series of piezoelectric cantilever beams with various<br />
volumes functioning in the range of 230 Hz until 1.5 kHz<br />
were simulated. The cantilever beam operates in<br />
transversal mode (d 31 mode) where top and bottom<br />
electrodes were used to measure the output voltage<br />
produced. A comparison of power output produced<br />
between two different piezoelectric thin films materials<br />
(zinc oxide and aluminium nitride) is also highlighted.<br />
There are two simulation conditions applied: closed circuit<br />
conditions and open circuit conditions [5]. To simulate<br />
closed circuit conditions (when load resistance, R L = 0 as<br />
shown in equivalent circuit in Fig. 4), the potentials of the<br />
electrodes are set to 0 V. For open circuit conditions, one<br />
electrode is set to TiePotential, and the other to a potential,<br />
e.g. 0 V [5]. The load resistance is fixed to 50 for a<br />
series of simulation.<br />
(3)<br />
(4)<br />
TABLE II<br />
FABRICATION STEPS OF MEMS PIEZOELECTRIC INERTIAL GENERATOR<br />
Nu. Step Name Layer name Material Name Thickness Mask Name Comment<br />
0 Substrate Substrate Silicon_100 500 m Substrate Mask<br />
1 Stack Material Bottom_electrode Aluminium 5 m<br />
2 Stack Material Piezoelectric Zinc Oxide 10 m<br />
3 Stack Material Top_electrode<br />
K<br />
Aluminium 5 m<br />
4 Straight Cut eq<br />
ω<br />
n<br />
=<br />
Cantilever Back etching Depth:495 m<br />
5 Straight Cut M<br />
(1)<br />
Proof Mass Back etching Depth:200m<br />
eq<br />
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<br />
Fig. 4. Equivalent Circuit of MEMS Piezoelectric Energy Harvester<br />
III.<br />
SIMULATION RESULT AND DISCUSSION<br />
The simulation results indicate that the displacement at z<br />
direction of fundamental mode 1 provides highest<br />
displacement as shown in Fig. 5. A total of 30 harmonic<br />
frequency steps within range of 220 to 290 Hz were<br />
generated. Fig. 6 to 8 show the graph of displacement, output<br />
voltage and output power versus frequency applied<br />
respectively. The graphs show the expected sharp change and<br />
the peak as the frequency approach the fundamental<br />
resonance Mode 1 value that is 233 Hz.<br />
It is essential for the MEMS piezoelectric inertial generator<br />
to operate at the resonance frequency to harvest optimum<br />
power. Resonant frequency of 233 Hz provides the maximum<br />
displacement of vibration, output voltage and power. Fig. 6<br />
and 8 illustrate maximum displacement of 420µm and<br />
maximum output power of 3.02µW at mode 1 resonance<br />
frequency of 232 Hz. It shows that the device needs to<br />
operate at resonance mode to harvest optimum output power.<br />
Fig. 7. Output voltage produced versus frequency<br />
Fig. 8. Output power produced versus frequency<br />
A. The effect of various volume of the cantilever beam<br />
to its natural frequency<br />
Low resonant frequency is essential in MEMS<br />
piezoelectric inertial generator because most of the ambient<br />
vibrations are at very low frequencies [1]. The power output<br />
will only be optimized if the miniature devices are operated<br />
at the resonance mode. Table III and Fig. 9 illustrate the<br />
volume increment of the cantilever beam decreases the<br />
resonant frequency and increases the peak output power<br />
produced. The larger the volume of the cantilever beam, the<br />
lower the resonant frequency.<br />
Fig. 5. Mem Mech Piezoelectric Analysis<br />
B. The effect of resonant frequency to the peak output<br />
power<br />
Output power of the system will be optimized when it is<br />
operated at the ambient resonance frequency. MEM PZE<br />
piezoelectric analysis was done to compute the output power<br />
produced at the resonance frequency. To obtain resonance<br />
TABLE III<br />
Fig.6. Displacement of vibration versus frequency<br />
SIZES, RESONANT FREQUENCY AND OUTPUT POWER OF PIEZOELECTRIC INERTIAL<br />
GENERATOR<br />
Design<br />
Number<br />
Wbeam Lbeam Resonant<br />
Frequency<br />
Peak Output<br />
Power ( R L=50<br />
ohm)<br />
1 722 2044 1542.00 0.7260<br />
2 442 3018 976.40 0.5765<br />
3 166 3288 620.87 0.8350<br />
4 1010 4299 463.50 1.8216<br />
5 672 5039 232.0 3.0560<br />
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<br />
Design Number<br />
Fig. 9. Design number versus applied resonant frequency<br />
mode of the structure, modal analysis was performed. Thirty<br />
modes were simulated for six different volume of<br />
piezoelectric energy harvester. The mode with the highest<br />
displacement was identified as mode of interest [3].<br />
A range of 230 Hz to 1.5 kHz resonant frequency for five<br />
different designs is simulated to get maximum power output<br />
at the ambient resonant frequency. Fig. 10 shows the effect of<br />
resonance frequency to the output power produced. The<br />
lower the resonant frequency, the higher the output power<br />
produced. The highest output power is 3.056 W at 232 Hz<br />
resonant frequency. The power output produced is enough to<br />
power the wireless condition monitoring circuits since power<br />
plants generate ambient vibrations at low frequencies.<br />
C. The effect of chosen piezoelectric material on output<br />
power.<br />
MEMS technology has introduced the concept of<br />
many functional materials with new fabrication process<br />
which allow a creation of miniature devices consuming less<br />
power, reliable and integrate multiple functions. One of the<br />
main functional materials is piezoelectric thin film.<br />
Piezoelectric materials develop charge on the sample surfaces<br />
when exposed to applied stresses or vibration. [6].<br />
The choice of the piezoelectric thin films depends<br />
on the process complexity, piezoelectric coupling coefficient<br />
and CMOS compatibility. The most common materials used<br />
are aluminium nitride (AlN), zinc oxide (ZnO) and lead<br />
zicronate titanate (PZT). PZT provides highest coupling<br />
coefficient. However, PZT thin film deposition is very<br />
complex and hazardous due to lead contamination. AlN and<br />
ZnO are both wurtzite structure materials with the polar<br />
direction [6]. Low deposition temperature is kept during<br />
sputtering process to obtain high quality of piezoelectric AlN<br />
and ZnO films and to allow complete integration capabilities<br />
with CMOS technology [6]. The low deposition temperature<br />
also offers the use of standard Al for metallization layers<br />
(electrodes). The sputter deposition techniques for both AlN<br />
and ZnO is also well-known standard deposition and is less<br />
complex compared to the deposition of PZT. Therefore, the<br />
simulations are shown to discuss the effect of AlN and ZnO<br />
piezoelectric materials only.<br />
Fig. 10. Output power produced versus applied resonant frequency<br />
The main material properties for the analysis are<br />
piezoelectric strain coefficient; dielectric constant and<br />
stiffness matrix. The material properties for ZnO and AlN<br />
piezoelectric material are summarized in Table IV and Table<br />
V [5] [7]. The Dielectric entries shown are relative values to<br />
vacuum permittivity 0 = 8.85 x 10 -12 c/(vm) [5].<br />
The aim of the simulation analysis is to compare the<br />
output power produced at the resonance frequency for two<br />
different piezoelectric material used: AlN and ZnO. Table<br />
VII and Fig. 11 show that zinc oxide provides more energy<br />
output compared to aluminium Nitride (AlN). At 242 Hz<br />
resonant frequency, the peak power output produced for ZnO<br />
is 3.0560 µW and 0.8000 µW for AlN for the same volume<br />
TABLE IV<br />
ZINC OXIDE AND ALUMINUM NITRIDE PIEZOELECTRIC PROPERTIES<br />
Parameter ZnO AlN<br />
Density ( kg/m 3 ) 5.8 e -13 3.2 e -15<br />
Coupling Coefficient, k 0.33 0.24<br />
Relative dielectric<br />
constant, <br />
10.9 10.5<br />
TABLE V<br />
ELASTIC CONSTANTS FOR ZINC OXIDE , [C ZnO] 6X6 AND ALUMINUM NITRIDE , [C AIN] 6X6<br />
C COLUMN X ROW ZnO (N/m 2 ) AlN (N/m 2 )<br />
C 11 2.907 x 10 5 3.450 x 10 5<br />
C 21 1.21 x 10 5 1.250 x 10 5<br />
C 22 2.907 x 10 5 3.450 x 10 5<br />
C 31 1.051 x 10 5 1.200 x 10 5<br />
C 32 1.051 x 10 5 1.200 x 10 5<br />
C 33 2.109 x 10 5 3.950 x 10 5<br />
C 44 4.430 x 10 4 1.100 x 10 5<br />
C 55 4.240 x 10 4 1.180 x 10 5<br />
C 66 4.240 x 10 4 1.180 x 10 5<br />
TABLE VI<br />
PIEZOELECTRIC STRAIN COUPLING MATRIX COEFFICIENT FOR ZINC OXIDE , [dZnO] 3X6<br />
AND ALUMINUM NITRIDE , [d AIN] 3X6<br />
Parameters ZnO (C/m 2 ) AlN (C/m 2 )<br />
d 31 -5.430 x 10 -6 -5.800 x 10 -1<br />
d 33 1.167 x 10 -5 1.55<br />
d 15 1.134 x 10 -5 4.800 x 10 -1<br />
- d 15 -1.134 x 10 -5 -4.800 x 10 -1<br />
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<br />
BRIEF BIOGRAPHY OF THE AUTHOR<br />
Aliza Aini Md Ralib received the B. Eng (Hons) Computer and<br />
Information Engineering from the International Islamic University Malaysia<br />
(IIUM), Kuala Lumpur, Malaysia in 2006 and currently working toward<br />
Master Degree in Electronic Engineering. Her main research interests are<br />
VLSI and MEMS. She currently involves in the design and simulation of<br />
MEMS piezoelectric energy harvesting devices for wireless sensor in power<br />
plant application.<br />
Fig. 11. Output power produced versus applied resonant frequency for<br />
two different piezoelectric materials<br />
TABLE VII<br />
OUTPUT POWER FOR TWO DIFFERENT PIEZOELECTRIC MATERIALS<br />
Design<br />
( Wbeam x beam)<br />
Design 4 :<br />
1010 x 4299<br />
Design 5 :<br />
672 x 5039<br />
Resonant<br />
Frequency<br />
(Hz)<br />
Output Power (µW)<br />
AlN ZnO<br />
463.5 0.2000 1.8216<br />
242 0.8000 3.0560<br />
of cantilever beam. At 463.5 Hz resonant frequency, the peak<br />
output power produced for ZnO is 1.8216 µW compared to<br />
smaller output power for AIN material 0.200 µW. This is due<br />
to the fact that the zinc oxide has higher piezoelectric<br />
coupling coefficient compared to aluminium nitride as shown<br />
in Table IV.<br />
IV. CONCLUSION<br />
We presented in this paper the mechanical finite element<br />
solution of MEMS piezoelectric energy harvesting generating<br />
at low frequencies. Cantilever structure has been identified as<br />
the most suitable structure for maximum energy conversion.<br />
The performance of the miniature devices were evaluated by<br />
varying the volume of the cantilever beam (length and width<br />
of the cantilever beam). Volume increment of the cantilever<br />
beam decreases the resonant frequency and increases the<br />
peak power output produced. The effect of the resonance<br />
frequency to the output power was also discussed in this<br />
paper. The lower the resonance frequency, the higher the<br />
output power produced. Selection of piezoelectric material<br />
was also essential to get the optimize output power. Zinc<br />
oxide piezoelectric material produced larger output power<br />
compared to aluminium nitride at the same resonance<br />
frequency due to the zinc oxide has higher piezoelectric<br />
coupling coefficient compared to aluminium nitride.<br />
ACKNOWLEDGMENT<br />
The research was supported by the R&D grant from Tenaga<br />
Nasional Berhad Malaysia and collaboration between<br />
Universiti Tenaga Nasional Malaysia and International<br />
Islamic University Malaysia.<br />
Anis Nurashikin Nordin received the B. Eng. Degree in Computer and<br />
Information Engineering from the International Islamic University Malaysia<br />
(IIUM), Kuala Lumpur, Malaysia in 1999, and the M.S degree in Computer<br />
Engineering from the George Washington University (GWU), Washington<br />
DC, in 2002, and the D. Sc. Degree in Electrical and Computer Engineering<br />
at GWU. Currently, she is a lecturer at International Islamic University<br />
Malaysia (IIUM).Her main research interests are VLSI and RF MEMS,<br />
SAW Resonators, and particularly oscillators.<br />
Raihan Othman received the B Sc..in Physics, and the M.Sc degree in<br />
ThinFilm Technology, and Ph.D in Electrochemical Power Sources, all<br />
from the University of Malaya (UM) , Kuala Lumpur. Currently, he is a<br />
lecturer at International Islamic University Malaysia (IIUM). His main<br />
research interests are metal-air electrochemical system and biological fuel<br />
cells.<br />
Hanim Salleh received the B. Sc in Agricultural Engineering from the Cum<br />
Laude, Georgia, USA, and the M.S degree in Agricultural Process<br />
Engineering from the Universiti Putra Malaysia and PhD in sound and<br />
vibration studies at University of Southampton, United Kingdom. Currently,<br />
she is a senior lecturer at Universiti Tenaga Nasional Malaysia (UNITEN)<br />
since 2007.Her main research interests are dynamic systems, vibrations<br />
control, instrumentation and energy harvesting.<br />
REFERENCES<br />
[1] Jong Cheol Park; Jae Yeong Park; Yoon-Pyo Lee; , "Modeling<br />
and Characterization of Piezoelectric d 33-Mode MEMS Energy<br />
Harvester," Microelectromechanical Systems, Journal of , vol.19,<br />
no.5, pp.1215-1222, Oct. 2010<br />
[2] Priya, Shashank, Inman, Daniel J. (2009). Energy Harvesting<br />
Technologies.<br />
[3] Aini Md Ralib, A.; Nurashikin Nordin, A.; Salleh, H.; ,<br />
"Simulation of a MEMS piezoelectric energy harvester," Design<br />
Test Integration and Packaging of MEMS/MOEMS (DTIP), 2010<br />
Symposium on , vol., no., pp.177-181, 5-7 May 2010<br />
[4] Anis Nurashikin Nordin (2008) Design, Implementation and<br />
Characterization of Temperature Compensated SAW Resonators<br />
in CMOS Technology for RF Oscillators.<br />
[5] Coventorware 2010 Application Notes<br />
[6] Bassiri-Gharb, Nazanin (2008) Piezoelectric and Acoustic<br />
Materials for Transducer Applications. Springer USpp. 413-430.<br />
[7] Coventerware Version 2006. MEMS Design and Analysis<br />
Tutorials, Volume 1.<br />
[8] Shad Roundy, Jan M.Rabaey and Paul Kenneth Wright (2003).<br />
Energy Scavenging for Wireless Sensor Networks. Kluwer<br />
Academic Publishers.<br />
[9] Paradiso, J.A.; Starner, T.; , "Energy scavenging for mobile and<br />
wireless electronics," Pervasive Computing, IEEE , vol.4, no.1,<br />
pp. 18- 27, Jan.-March 2005<br />
[10] W.J. Choi • Y. Jeon• J.-H. Jeong• R. Sood• S.G. Kim (2006)<br />
“Energy Harvesting MEMS device based on thin film<br />
piezoelectric cantilevers”, Electroceramics Journal 2006, Volume<br />
17, Numbers 2-4, December 2006.<br />
[11] Ralib, A.A.M.; Nordin, A.N.; Salleh, H.; , "Theoretical modeling<br />
and simulation of MEMS piezoelectric energy harvester,"<br />
Computer and Communication Engineering (ICCCE), 2010<br />
International Conference on , vol., no., pp.1-5, 11-12 May 2010<br />
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<br />
Parameter Design of Triaxial Microaccelerometers<br />
with Piezoelectric Thin-Film<br />
Jyh-Cheng YU<br />
National Kaohsiung First University of Science and Technology<br />
2 Jhuoyue Rd.,Nanzih , Kaohsiung City 811,Taiwan, R.O.C.<br />
Chungda Lee<br />
Department of Mechanical and Automation Engineering<br />
I-Shou University, 1, Sec. 1, Syuecheng Rd., Dashu, Kaohsiung 840, Taiwan, R.O.C.<br />
Abstract - This study proposes an analytical model for a<br />
high sensitivity piezoelectric thin film triaxial<br />
microaccelerometer, and investigates the influence of the<br />
fabrication processes on the parameter design. The structure<br />
design is consisted of four parallel suspension beams, a<br />
central seismic mass, and eight piezoelectric thin film<br />
transducers. The sensitivity consistence between out-of-plane<br />
and in-plane accelerations is a key issue for the following<br />
signal processing. A simplified system modeling scheme<br />
based on anisotropic material properties using area moment<br />
method and laminated beam theory is presented and applied<br />
to the parameter design. An optimized thickness ratio<br />
between piezoelectric thin film and the silicon substrate of<br />
the laminated supporting beam is derived to maximize the<br />
sensitivity. The study shows that the aspect ratio of the<br />
seismic mass is the deterministic factor to the differences<br />
among triaxial sensitivities. The triaxial sensitivity<br />
performances for two structure designs with the seismic<br />
masses fabricated using chemical wet etching and deep<br />
reactive ion etching (DRIE) are compared. The design using<br />
DRIE provides more even triaxial sensitivity, while the<br />
design using wet etching shows cost advantage with<br />
additional parameter constraints due to the required<br />
compensation pattern for convex corner etching.<br />
I. INTRODUCTION<br />
For the last two decades, the design and research of<br />
piezoelectric sensing accelerometers have drawn great<br />
attention due to the advantages of low cost, energy saving,<br />
simple structure, high sensitivity, and excellent dynamic<br />
performance. Bulk micromachining based piezoelectric<br />
accelerometers have a lower detection level that is suitable<br />
for precision measurement. Piezoelectric accelerometers<br />
consisted of a single seismic mass and cross supported beams<br />
have been investigated for uniaxial measurement [1][2][3]<br />
and triaxial measurement [4] applications. Most analytical<br />
models assume piezoelectric thin film for simplification.<br />
Hindrichsen [5] proposed an analytical model for a PZT thick<br />
film triaxial accelerometer based on anisotropic material<br />
tensors and Euler’s beam.<br />
In general, accelerometers with PZT thick film<br />
transducers have higher charge and voltage sensitivities in<br />
comparison with those thin film devices in the same given<br />
dimensions. Another configuration of piezoelectric thin-film<br />
triaxial microaccelerometers consisted of parallel beam<br />
suspensions and a central seismic mass has been studied<br />
[6][9]. Parallel beam suspensions provide higher sensitivity<br />
for in-plane accelerations comparing with those using a cross<br />
beam structure.<br />
Different fabrication processes of seismic mass, such as<br />
chemical wet etching and deep reactive ion etching (DRIE),<br />
affect the parameter design for device dimensions. Most of<br />
the triaxial accelerometers do not have equal sensitivity for<br />
all the three directions. Accelerometers fabricated using<br />
DRIE of a thick SOI wafer could provide better consistence<br />
of sensitivity in all the three orthogonal axes with a proper<br />
parameter design. On the other hand, wet etching process has<br />
a great cost advantage with dimensional limitations due to<br />
sloping walls and a required compensation pattern for convex<br />
corners.<br />
This work presents the system modeling and parameter<br />
design based on fabrication constraints of a triaxial<br />
piezoelectric accelerometer. The proposed configuration<br />
suspension of the device adopts parallel beams at both ends<br />
of a seismic mass using etching of (100) SOI wafer. This<br />
study adopts the area moment method to determine the<br />
stiffness matrix of the supporting beams describing the<br />
relationship between the end forces and moments and the<br />
boundary conditions. The elastic property of the suspension<br />
beams considers both the silicon beams and the piezoelectric<br />
films using the laminated beam theory based on anisotropic<br />
material properties. Analytical models of the resonant<br />
frequency and the sensitivity are verified with the results<br />
using finite element method to justify the model accuracy.<br />
Three axial sensitivities for applicable conditions of design<br />
geometry due to different fabrication processes are<br />
compared.<br />
II. DESIGN OF MICROACCELEROMETER<br />
The proposed microaccelerometer consists of a<br />
quadri-beam suspension, a seismic mass, and eight<br />
90
piezoelectric displacement transducers. Each of the two<br />
opposite ends of the seismic mass is supported by two<br />
parallel suspension beams. The thickness of suspension<br />
beams is defined by the silicon-on-insulator (SOI). Two<br />
transducers are patterned on each suspension. From a proper<br />
interconnection among the transducers, triaxial accelerations<br />
can be measured without cross-axis interference [9].<br />
The design of the accelerometer depends on the processes<br />
of bulk micromachining for the seismic mass. The seismic<br />
mass can be fabricated from either Deep Reactive Ion<br />
Etching (DRIE) as shown Figure 1, or chemical wet etching<br />
as shown in Figure 2, following by dry etching to release the<br />
suspension beams. Flexibility for the crystalline orientation<br />
of suspension beams and high aspect ratio of seismic mass<br />
are possible for DRIE. The suspension beams are often [100]<br />
oriented in the previous studies [3][4][5][6]. Chemical wet<br />
etching of (100) silicon wafer using such as KOH and<br />
TMAH has the cost advantage over dry etching. However,<br />
the sloping walls of the seismic mass cause additional<br />
constraints in the dimensional design on account of the<br />
required convex corner compensation in the masking layer of<br />
wet etching. Also, the crystalline orientation of the<br />
suspension beams will be restricted to [110].<br />
Figure 1 Design of the accelerometer using DRIE dry etching for the seismic<br />
mass<br />
Figure 2 Design of the accelerometer using wet etching for the seismic mass<br />
The out-of-plane (z axial) acceleration will result in a<br />
symmetric vibration, and in-plane (x and y axial)<br />
accelerations will produce asymmetric and torsional<br />
vibrations. The inertia force will introduce bending and<br />
torsional stress on the beams that will produce electrical<br />
charge by the piezoelectric transducers. The seismic mass<br />
from dry etching often provides higher sensitivity because<br />
the distance between the surfaces of suspension beams to the<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
center of mass is larger in contrast with that of the seismic<br />
mass from wet etching.<br />
III. ANALYTICAL MODEL<br />
The system model of the accelerometer is consisted of a<br />
mechanical subsystem and an electric subsystem. The<br />
mechanical subsystem is assumed a spring-mass-damper<br />
system. The equivalent stiffness of the suspension and the<br />
equivalent seismic mass are two deterministic factors for the<br />
structure modeling. Laminated beam theory is applied to<br />
obtain the equivalent bending rigidity of the supported beams.<br />
The application of area moment method provides the<br />
formulation of suspension stiffness for three vibration modes.<br />
The piezoelectric transducers convert beam stresses into<br />
output charge.<br />
A. Equivalent bending rigidity of the laminated beam<br />
The derivation of the structure model is based on the<br />
following assumptions:<br />
(1) the influence of electrodes on the beam stiffness is<br />
negligible;<br />
(2) the seismic mass and rim of the structure are rigid;<br />
(3) the deflections of substrate material and piezoelectric<br />
films of supporting beams observe linear elasticity and<br />
Hooke’s Law;<br />
(4) the piezoelectric material is anisotropic;<br />
(5) the supported beams are wide and flat, and thin beam<br />
theory (Euler-Bernoulli beam model) is applied; the<br />
stresses in the z-direction and the strains in the<br />
y-direction are negligible compared with others [7].<br />
Therefore,<br />
0543<br />
(1)<br />
0542<br />
The constitutive equation can be simplified as follows<br />
for an orthotropic piezoelectric thin film such as PZT:<br />
1<br />
p p<br />
CCC<br />
p<br />
1 3, 2, 1,<br />
0 1<br />
<br />
<br />
<br />
<br />
<br />
2 p,21<br />
p,22<br />
CCC<br />
p,23<br />
0 <br />
2 <br />
<br />
<br />
<br />
<br />
(2)<br />
<br />
3 p p<br />
CCC<br />
p<br />
3 3, 2, 1,<br />
0 3<br />
<br />
<br />
<br />
<br />
6 <br />
000<br />
C<br />
p 6 6,<br />
6<br />
<br />
From the stress/strain assumptions in (1) and the expression<br />
for σ 3 in (2), we obtain<br />
C<br />
3<br />
<br />
(3)<br />
C<br />
p 3 1,<br />
1<br />
p 3 3,<br />
Similarly, with the aid of (3), stress σ 1 can be represented as<br />
follows:<br />
<br />
<br />
<br />
CC <br />
<br />
(4)<br />
<br />
<br />
,13CC<br />
pp<br />
,31<br />
CE (5)<br />
,13 pp<br />
,31<br />
1<br />
C<br />
p 1 1,<br />
1<br />
EP<br />
<br />
1<br />
C <br />
p 3 3,<br />
pP 1 1,<br />
C<br />
p 3 3,<br />
where E p is the effective modulus of elasticity for the<br />
piezoelectric film.<br />
91
11-13 <br />
May 2011, Aix-en-Provence, France<br />
Also, the cross section of the supported beams is assumed<br />
<br />
to be wide and flat, and under the plane stress condition.<br />
Therefore, the effective modulus of elasticity, E b , for the<br />
silicon beams can be derived similarly as (6)<br />
CS,13CS,31<br />
EB<br />
CS,11<br />
<br />
(6)<br />
CS,33<br />
(a) Symmetric bending<br />
where C S,11 , C S,13 , C S,13 and C S,33 are the stiffness coefficients<br />
of silicon beams.<br />
Here we apply laminated beam theory to derive the<br />
equivalent bending rigidity [8]. The laminated beam can be<br />
treated as an equivalent beam of the same material as<br />
substrate layer with the width of the piezoelectric layer<br />
adjusted according the elasticity ratio to the substrate<br />
material’s as shown in Figure 3. The distance from the<br />
neutral axis Y to the interface can then be obtained as (7).<br />
a <br />
1<br />
2<br />
2<br />
Btb<br />
B b<br />
2<br />
Pt<br />
p<br />
E E<br />
E t E t<br />
P P<br />
By parallel axis theorem, the equivalent moment of<br />
inertia about Y-axis can be derived, and the equivalent<br />
bending rigidity of the composite beam is as follows [9]:<br />
3<br />
3<br />
t<br />
<br />
<br />
<br />
P 2 2 t<br />
<br />
b 2 2<br />
( EI <br />
<br />
Y<br />
)<br />
eq<br />
EPwb<br />
tPa<br />
tPa<br />
EBwb<br />
tba<br />
tba<br />
(8)<br />
3<br />
3<br />
<br />
EP<br />
wb<br />
<br />
E<br />
B<br />
(7)<br />
(b) Asymmetric bending<br />
(c) Torsional bending<br />
Figure 4 Three principal vibration modes of the proposed device<br />
Figure 5 Beam with arbitrary boundary conditions<br />
Assume the supporting beams are fixed at the rim of the<br />
vibration device, from the boundary conditions shown in<br />
Figure 6, the stiffness matrix can be simplified as (10).<br />
Figure 3 Cross section of the equivalent piezoelectric film/Si beam<br />
B. Derivation of suspension stiffness using area moment<br />
method<br />
The elastic property of the suspension beams considers<br />
both the silicon beams and the piezoelectric films using the<br />
laminated beam theory. The supported beams are assumed<br />
wide and flat, and Euler-Bernoulli or thin beam theory<br />
applies. The stresses in the z-direction and the strains in the<br />
y-direction are negligible in comparison with others to<br />
simplify the model. The relationship between the end forces,<br />
moments, and the boundary conditions as shown in Figure 4<br />
and Figure 5 can be described using the stiffness matrix as<br />
(9).<br />
F<br />
<br />
M<br />
<br />
F<br />
<br />
M<br />
1<br />
1<br />
2<br />
2<br />
k<br />
<br />
k<br />
<br />
<br />
k<br />
<br />
<br />
k<br />
11<br />
21<br />
31<br />
41<br />
k<br />
k<br />
k<br />
k<br />
12<br />
22<br />
32<br />
42<br />
k<br />
k<br />
k<br />
k<br />
13<br />
23<br />
33<br />
43<br />
k<br />
k<br />
k<br />
k<br />
14<br />
24<br />
34<br />
44<br />
u1<br />
<br />
<br />
<br />
<br />
<br />
1<br />
<br />
<br />
u<br />
2 <br />
<br />
<br />
<br />
<br />
<br />
2 <br />
(9)<br />
Figure 6 Free body diagram of the beam with one fixed end<br />
F<br />
m k11<br />
k12<br />
u<br />
<br />
<br />
M<br />
m k21<br />
k22<br />
(10)<br />
<br />
This study adopts the area moment method [10] to<br />
determine the stiffness matrix [k] of the supporting beam.<br />
12EI<br />
6EI<br />
<br />
F<br />
m 3 2 <br />
<br />
l l<br />
u<br />
<br />
<br />
<br />
M<br />
<br />
6EI<br />
4EI<br />
m <br />
<br />
<br />
(11)<br />
<br />
2<br />
l l <br />
The displacement conditions of the supporting beams for<br />
three vibration modes are shown in Table 1.<br />
Table 1 Displacement conditions of supported beams<br />
u<br />
θ<br />
Symmetric δ m 0<br />
Asymmetric l m α / 2 α<br />
Torsional (w m – w b ) β / 2 0<br />
92
Figure 7 Displacement conditions of supported beams for symmetric<br />
bending<br />
m<br />
Figure 8 Displacement conditions of supported beams for asymmetric<br />
bending<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
where (EI Y ) eq is the equivalent bending rigidity about Y<br />
axis ,and l b is the length of the suspension beam.<br />
Follow the similar procedure, the equivalent asymmetric<br />
stiffness and the equivalent torsional stiffness can be derived<br />
as listed in Table 6.<br />
The equivalent mass of the structure considers the seismic<br />
mass and the effective mass of the supported beams. A simple<br />
integration of the structure will give the seismic mass:<br />
<br />
m <br />
<br />
lmwmh<br />
<br />
l<br />
w t<br />
m<br />
m b<br />
m<br />
h<br />
2<br />
m<br />
( l<br />
m<br />
4<br />
wm<br />
)cot<br />
h<br />
3<br />
3<br />
m<br />
2 <br />
cot <br />
<br />
(17)<br />
where is the density of Silicon, = 54.7° for anisotropic<br />
wet etching of silicon, and = 90° for dry etching.<br />
Figure 9 Displacement conditions of supported beams for torsional bending<br />
C. Analytical formula of resonant frequency<br />
The resonant frequency of the structure is determined by<br />
the equivalent stiffness of the suspension and the equivalent<br />
system mass. The out-of-plan resonant frequency f n and<br />
in-plan resonant frequencies f n,a f n,t can then be represented<br />
as follows:<br />
f<br />
1<br />
K<br />
s<br />
n,<br />
s<br />
(12)<br />
2<br />
mt,<br />
s<br />
K<br />
a<br />
n,<br />
a<br />
<br />
2<br />
J<br />
(13)<br />
t,<br />
a<br />
f<br />
f<br />
1<br />
1<br />
K<br />
t<br />
n,<br />
t<br />
<br />
2<br />
J<br />
(14)<br />
t,<br />
t<br />
where m t,s is the equivalent total mass and K s is the stiffness<br />
for the symmetric vibration mode, J t,a is the equivalent<br />
moment of inertia and K a is the torsional stiffness for the<br />
asymmetric vibration mode, and J t,t is the equivalent moment<br />
of inertia and K t is the stiffness for the torsional vibration<br />
mode.<br />
For symmetric bending, the end force relates to the<br />
displacement from (11) as follows,<br />
F m 11<br />
k u<br />
(15)<br />
The stiffness of the system for symmetric vibration can be<br />
given as<br />
K<br />
s<br />
48( EI )<br />
(16)<br />
Y eq<br />
3<br />
lb<br />
Figure 10 Geometric parameters for the proposed accelerometer design<br />
The effective mass of the supporting beams, m eq , can be<br />
derived by kinetic energy method [10].<br />
m ( )<br />
eq<br />
cm mb (18)<br />
where m b is the actual mass of the laminated composite beam<br />
and c m is the effective constant which is 13/35 for symmetric<br />
bending. Similarly, the equivalent rotational moment of<br />
inertia is consisted of the equivalent rotational moment of<br />
inertia of the seismic mass and the effective rotational<br />
moment of inertia of the supported beams.<br />
D. Analytical formula of sensitivity response<br />
The general system modeling for three primary vibration<br />
modes shown in Figure 4 can be represented as (19):<br />
2<br />
e s ni<br />
( s)<br />
Si<br />
<br />
2<br />
2<br />
ai<br />
s<br />
1<br />
s 2nis<br />
<br />
(19)<br />
ni<br />
where S i is the corresponding open circuit voltage<br />
sensitivity and ω ni = 2π×(f n,i ) is the corresponding resonant<br />
frequency for each primary vibration mode i.<br />
The electric subsystem is determined by the piezoelectric<br />
transducers. When the inertial force of the seismic mass is<br />
acting on the beam, an infinitesimal charge will be generated<br />
due to the stresses on the piezoelectric transducer. Integration<br />
of the charge along the suspension beams yields the charge<br />
output of piezoelectric transducers.<br />
For instance of symmetric vibration, the application of<br />
Hook’s law provides the displacement of the seismic mass<br />
under vertical acceleration a Z<br />
.<br />
m a<br />
t z<br />
( lb)<br />
<br />
m<br />
<br />
(20)<br />
Ks<br />
u<br />
93
The deflection equation of the Euler’s beam can be<br />
derived from the boundary conditions.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
2<br />
)23 x xlx () ( u<br />
16.57<br />
6.39 6.39 0 0 0 <br />
<br />
6.39 16.57 6.39 0 0 0<br />
<br />
<br />
<br />
bm)2(6)()(<br />
zxux zxl<br />
6.39 6.39 16.57 0 0 0 <br />
10 2<br />
C S<br />
)/(1<br />
100 <br />
mN (28)<br />
<br />
009 <br />
6<br />
<br />
09 <br />
6 .<br />
<br />
<br />
<br />
9 <br />
6 .<br />
19.44<br />
3.52 6.39 0 0 0 <br />
t<br />
xlx<br />
) <br />
bm<br />
<br />
a<br />
(2(6)(<br />
3.52 19.44 6.39 0 0 0<br />
<br />
<br />
<br />
2<br />
6.39 6.39 16.57 0 0 0 <br />
10 2<br />
C S<br />
mN )/(1<br />
110 <br />
<br />
(29)<br />
<br />
009 <br />
6<br />
<br />
09 <br />
6 .<br />
<br />
<br />
<br />
0 <br />
9 .<br />
bm<br />
(21)<br />
The strain distribution for the beam is as follows,<br />
<br />
(22)<br />
Therefore, the average strain of the piezoelectric film along<br />
the beam due to the bending moment is as (23)<br />
m<br />
(23)<br />
From the free-body diagram of Figure 7, the bending<br />
moment along the supported beam for a given displacement<br />
δ Z of the seismic mass can be represented as follows:<br />
M )( t <br />
P EP<br />
<br />
1<br />
a <br />
<br />
<br />
(25)<br />
I<br />
eq 2 EB<br />
<br />
If the stresses other than σ 1 caused by bending of the<br />
piezoelectric film are negligible and no external electrical<br />
field is applied, the contribution from an infinitesimal portion<br />
of the piezoelectric material to the total charge is as follows:<br />
3 mdD<br />
13<br />
(26)<br />
where d 31 is the transverse piezoelectric charge to stress ratio.<br />
Eight transducers in total are interconnected such that<br />
triaxial accelerations can be measured selectively.<br />
A proper interconnection of eight transducers in Figure 4<br />
can eliminates cross axis sensitivity, and provides accurate<br />
measurement of triaxial accelerations using the proposed<br />
configuration [9]. From simple integration of the generated<br />
charge along the beams, the sensor’s open-circuit voltage<br />
sensitivity S Z can be derived as follows:<br />
V 4Q<br />
Z , bs 31<br />
tlEdm<br />
P <br />
b <br />
Pt<br />
SZ<br />
<br />
a<br />
(27)<br />
aZ<br />
ZCa<br />
EI<br />
e<br />
<br />
33 qY<br />
2)( <br />
2<br />
where C is the capacitance of the piezoelectric transducer,<br />
and <br />
33<br />
is the dielectric permittivity of piezoelectric film.<br />
Similar procedures can be applied to asymmetric and<br />
torsional vibrations. Summary of analytical modeling<br />
parameters for the tri-axial acceleration sensing are listed in<br />
Table 6.<br />
IV. VERIFICATION USING FEM MODEL<br />
Typically, the structure of the proposed<br />
microaccelerometer is fabricated using a (100) SOI wafer.<br />
There is no specific dimensional constraint if DRIE is<br />
applied to fabricate the se(11)ismic mass. The suspension<br />
beams can be aligned to either [100] or [110] of the silicon<br />
crystalline directions. However, if chemical wet etching is<br />
applied for the seismic mass, the suspension beams are<br />
aligned to [110]. The corresponding stiffness matrices of the<br />
silicon supported beams for [100] and [110] aligned are<br />
shown in (28) and (29) respectively.<br />
EI <br />
ze<br />
EI )( <br />
1 The accuracy 2)( of the analytical model is verified<br />
qY ze qY<br />
6<br />
using a<br />
M )( <br />
<br />
2<br />
3<br />
(24) 3-D ANSYS FEM model. Anisotropic material properties of<br />
lb<br />
lb<br />
PZT and silicon are introduced to both the analytical models<br />
The average bending stress σ 1 of the piezoelectric film due to and the FEM analysis. The results of resonant frequency and<br />
the bending moment is as follows:<br />
the open circuit voltage sensitivity are compared in Table 4<br />
and Table 5. The small error percentages demonstrate<br />
satisfactory accuracy.<br />
Table 2 Dimensions of the tri-axial accelerometer with the seismic mass<br />
from wet etching<br />
l b w b t b t p l m w m h m<br />
800 200 20 1 2200 2200 490<br />
(unit: μm)<br />
Table 3 Dimensions of the tri-axial accelerometer with the seismic mass<br />
from dry etching<br />
l b w b t b t p l m w m h m<br />
200 60 10 1 1000 1000 490<br />
(unit: μm)<br />
Table 4 Verification of the resonant frequency for chemical wet etching<br />
design<br />
Analytical<br />
Model<br />
FEM<br />
Error<br />
Symmetric 1.17 1.17 0.3%<br />
Asymmetric 2.56 2.49 2.8%<br />
Torsional 1.72 1.70 1.0%<br />
(Unit: KHz)<br />
Table 5 Verification of the sensitivity for chemical wet etching design<br />
Analytical<br />
Model<br />
FEM<br />
Error<br />
Symmetric 40.90 39.90 2.5%<br />
Asymmetric 6.48 6.40 1.2%<br />
Torsional 9.05 8.77 3.2%<br />
(Unit: mV/g)<br />
Table 5 Resonant frequency and output sensitivity for dry etching design<br />
Resonant frequency<br />
(KHz)<br />
Voltage sensitivity<br />
(mV/g)<br />
Symmetric 11.3 7.04<br />
Asymmetric 16.8 2.83<br />
Torsional 13.2 3.43<br />
IV. PARAMETER STUDY<br />
The objective of the design is to obtain high and even<br />
sensitivity in three principal directions with the constraints of<br />
the minimum bandwidth and the device size. For the seismic<br />
mass fabricated from wet etching, the walls are sloping as<br />
Figure 10. Also, a pertinent convex corner compensation<br />
pattern is required for the masking design in etching to<br />
94
protect the corners from undercut. Therefore, parameter<br />
design of the structure of the accelerometer will be<br />
constrained by the fabrication processes. If a typical <br />
band compensation [11] is used as shown in Figure 11, the<br />
constraints among the size of the seismic mass and the length<br />
of the supported beams are listed in (30) to (32).<br />
m<br />
23 hw<br />
m<br />
(30)<br />
b<br />
21.1<br />
hl<br />
m<br />
(31)<br />
m<br />
23 hl<br />
m<br />
(32)<br />
Table 1 and Table 2 show the parameters of two sample<br />
devices fabricated using chemical wet etching and DRIE<br />
respectively. The thickness of the seismic mass is assumed to<br />
be 490 (μm) for a typical 4-inch (100) wafer. The device size<br />
of the accelerometer using wet etching is larger owing to the<br />
compensation pattern required for the convex corners of<br />
seismic mass.<br />
Also, longer suspension beams and a larger seismic mass<br />
cause a lower resonant frequency and larger sensitivity. The<br />
triaxial sensitivities for the accelerometer fabricated using<br />
DRIE are more consistent than those using wet etching. The<br />
sensitivity due to symmetric vibration is the largest while the<br />
sensitivity on account of asymmetric vibration is the lowest.<br />
The difference can be reduced if a thicker wafer is available,<br />
which results in a thicker seismic mass.<br />
26.1<br />
Figure 11 Convex corner compensation using band<br />
For instance of the design using wet etching, select three<br />
different thickness of the silicon substrate beam and vary the<br />
thickness of the piezoelectric thin film. The sensitivity of<br />
out-of-plane acceleration is shown in Figure 12. As expected,<br />
the thinner the supporting beams, the higher the response<br />
sensitivity. However, the maximum sensitivity occurs at t p /t b<br />
= 0.7~0.8.<br />
2<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Response sensitivity for out-of-plane<br />
acceleration<br />
0.1<br />
0.09<br />
0.08<br />
0.07<br />
0.06<br />
0.05<br />
0.04<br />
0.03<br />
0.02<br />
0.01<br />
0<br />
0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3<br />
tp / tb<br />
Figure 12 Parameter study for the thickness ratio between piezoelectric film<br />
and silicon substrate of supporting beam (t p /t b )<br />
Here, we define the deviation coefficient as the<br />
performance index for triaxial accelerometer.<br />
<br />
<br />
2 max(<br />
i<br />
) min( SS<br />
i<br />
)<br />
DC (33)<br />
max( ) min( SS<br />
)<br />
i<br />
i<br />
where max(S i ) is the maximum sensitivity and min(S i ) is the<br />
minimum sensitivity among triaxial sensitivities. A smaller<br />
DC represents a smaller variation among triaxial sensitivities.<br />
Our investigation shows that the thickness ratio t p /t b , the<br />
aspect ratio of seismic mass w m /l m , and the width and depth<br />
ratio of supporting beam w b /l b have a negligible influence on<br />
DC. However, the thickness to width ratio of the seismic<br />
mass, h m /l m has a great impact on DC. The thicker the mass,<br />
the smaller the deviation coefficient.<br />
IV. CONCLUSION<br />
This study presents an analytical model for a high<br />
sensitivity tri-axial piezoelectric micro-accelerometer using<br />
piezoelectric thin film sensing. The modeling results of<br />
resonant frequency and sensor sensitivity are compared with<br />
FEM analysis. The proposed analytical model shows a good<br />
consistency under pertinent assumption of the design of<br />
transducers and suspension beams. The influence of the<br />
fabrication processes on the performance of the design is<br />
investigated. The model provides a good insight to the sensor<br />
design and can be applied to future design optimization.<br />
ACKNOWLEDGMENT<br />
The authors would like to thank the National Science<br />
Council (NSC 99-2221-E-327-029) and I-Shou University<br />
(ISU 94-02-15) of the Republic of China, Taiwan for<br />
financially supporting this research.<br />
Keywords: Area moment method, Piezoelectric thin film,<br />
triaxial microaccelerometer, Wet etching, System modeling<br />
REFERENCE<br />
t b (μm)<br />
[1] Yu J., Lan C.: System modeling of microaccelerometer using<br />
piezoelectric thin films. Sensors and Actuators A, 2001; 88, 2: 178-186.<br />
[2] Wang Q, Yang Z, Li F, Smolinski P: Analysis of thin film piezoelectric<br />
microaccelerometer using analytical and finite element modeling.<br />
Sensors and Actuators A, 2004; 113, 1: 1–11.<br />
[3] Wang L.-P., Wolf R.A., Jr.; Wang Y., Deng K.K., Zou L., Davis R.J.,<br />
Trolier-McKinstry S.: Design, fabrication, and measurement of<br />
20<br />
15<br />
10<br />
95
high-sensitivity piezoelectric microelectromechanical systems<br />
accelerometers, Journal of Microelectromechanical Systems, 2003, 12,<br />
4: 433-439.<br />
[4] Kunz K, Enoksson P, Stemme G: Highly sensitive triaxial silicon<br />
accelerometer with integrated PZT thin film detectors. Sensors and<br />
actuators A, 2001; 92: 156-160.<br />
[5] Hindrichsen C.C., Almind N.S., Brodersen S.H., Hansen O., and<br />
Thomsen E.V.: Analytial model of a PZT thick-film triaxial<br />
acceleromter for optimum design, IEEE Sensor Journal, 2009, 9, 4:<br />
419-429.<br />
[6] Zhu M, Kirby P, Lim MY: Lagrange’s formalism for modeling of a<br />
triaxial microaccelerometer with piezoelectric thin-film sensing,<br />
Sensors Journal, IEEE, 2004; 4, 4: 455 – 463.<br />
[7] Van Kampen RP, Woffenbuttel RF: Modeling the mechanical behavior<br />
of bulk-micromachined silicon accelerometers. Sensors and actuators A,<br />
1998; 64: 137-150.<br />
[8] Gere J. M., Timoshenko S. P. (1997) Mechanics of Materials, 4 th ed.,<br />
PWS <strong>Publishing</strong> Company, Boston, MA..<br />
[9] Yu J., Lee C., Chang C., Kuo W., Chang C.: Modeling Analysis of a<br />
Tri-Axial Microaccelerometer with Piezoelectric Thin-Film Sensing<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Using Energy Method, paper accepted, to appear in Journal of<br />
Microsystem Technologies.<br />
[10] Thomson W. T. and Dahleh M. D., Theory of vibration with application,<br />
5th ed., Prentice-Hall, New Jersey, 1998.: 23<br />
[11] Lang W.: Silicon Microstructuring Technology, Materials Science and<br />
Engineering: R: Reports, 1996 , 17, 1: 1-55.<br />
Brief biography of the corresponding author:<br />
Jyh-Cheng Yu graduated from the National Taiwan University,<br />
ROC, in 1985, and received the M.S. and Ph.D. degrees from<br />
the Department of Mechanical Engineering, the Ohio State<br />
University, USA. He now serves as a professor in the<br />
Department of Mechanical and Automation Engineering at<br />
the National Kaohsiung First University of Science and<br />
Technology. His research interests include design and<br />
manufacturing of piezoelectric microsensors, LCD back light<br />
module, and engineering optimization.<br />
Structure Stiffness<br />
Mass /<br />
Moment of Inertia<br />
Open Circuit<br />
<br />
s<br />
Sensitivity EI<br />
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S<br />
<br />
Table 6. Summary of modeling parameters for the tri-axial acceleration sensing<br />
Symmetric<br />
Asymmetric<br />
Torsional<br />
Mode<br />
Mode<br />
Mode<br />
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tlEdm<br />
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3 1,,<br />
bmp p<br />
pa 3 c p 1,,<br />
pb t<br />
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96
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Modeling and Experimental Validation of Levitating<br />
Systems for Energy Harvesting Applications<br />
Giorgio De Pasquale, Sonia Iamoni, Aurelio Somà<br />
Department of Mechanics, Politecnico di Torino, Corso Duca degli Abruzzi 24, 10129, Torino, Italy<br />
giorgio.depasquale@polito.it, sonia.iamoni@polito.it, aurelio.soma@polito.it.<br />
Abstract- The diamagnetic levitation principle is used to<br />
design an innovative typology of suspension for energy<br />
harvesting devices applied to very low frequency vibrating<br />
environments. The static configuration of the magnetostructural<br />
coupling is investigated starting from the theory of<br />
magnetism and a discrete numerical model is finally<br />
developed. The experimental validation is provided with<br />
measurements conducted by dedicated samples with a<br />
diamagnetic proof mass levitating in a magnetic field<br />
generated by permanent magnets. The results presented in this<br />
work provide important indications to the designer of<br />
microsystems for energy harvesting and the modeling<br />
approach proposed represent a valid design tool for coupled<br />
systems.<br />
I. INTRODUCTION<br />
Energy harvesting is a very promising strategy for the<br />
supplying of small systems and sensors that need energetic<br />
autonomy. Many applications may benefit from selfpowered<br />
systems, especially those related to sensing<br />
purposes in high energy vibrating environments: diagnostic<br />
systems for vehicles, structural monitoring, wireless sensors<br />
networks, measurement systems in laboratory facilities, etc.<br />
Very common problems related to the harvesting of energy<br />
from vibrations are the selection of transduction principle,<br />
the amplification of harvester bandwidth, the introduction of<br />
tuning systems, the duty cycle dimensioning and the global<br />
efficiency improvement. Many applications (e.g. sensing<br />
systems for vehicles, buildings, human body, etc.) imply<br />
very low vibration frequencies from the environment; this<br />
introduces additional problems to the tuning of the harvester<br />
and generally leads to higher proof masses and to<br />
limitations on miniaturization and integration. For these<br />
cases, the suspensions based on magnetic levitation<br />
represent a very promising opportunity to reduce the<br />
response of the harvester by preserving its small<br />
dimensions: compared to traditional mechanical<br />
suspensions, the stiffness of the magnetic interface is<br />
several orders of magnitude lower. Similar benefits interest<br />
MEMS energy harvesters, where very small masses are<br />
used [1-3]. Furthermore, the powerless functioning of these<br />
suspensions is very appreciable for the energetic efficiency<br />
of harvesters. The application of magnetic suspensions<br />
increases sensitively the lifetime of the harvesting device<br />
because the mechanical fatigue effects usually produced in<br />
the structural suspensions under alternate loads are<br />
completely avoided. Other advantages are given by the<br />
removal of mechanical bended elements, which are<br />
responsible to several energy dissipations sources:<br />
thermoelastic damping in the material, air damping under<br />
the suspensions, etc. [4]. The theoretical study of magnetic<br />
suspension was presented in some previous works, where<br />
analytic models and simulations were used [5,.6]; the<br />
magneto-structural coupling and the damping effect<br />
introduced by eddy currents were also described by Elbuken<br />
et al. [7]. Conversely, experimental measurements on<br />
levitating systems are not so diffused in literature [8, 9].<br />
This work describes the behavior of a magnetic<br />
suspension constituted by a layer of permanent magnets and<br />
a levitating diamagnetic proof mass. The static<br />
configuration of the suspension was studied by a finite<br />
element (FE) model; the results provided by the<br />
experimental validation are in good agreement with the<br />
levitation distance theoretically predicted. The models and<br />
characterizations presented are referred to a macrodimensional<br />
prototype of magnetic suspension. This is due<br />
to the easiness of fabrication and assembling and to the fact<br />
that micro fabrication techniques of magnets are still not<br />
completely mature, even if some promising samples were<br />
presented before [10]. However, the results obtained are<br />
suitable for the dimensioning of micro-scaled suspensions<br />
with similar topologies by a scaling procedure. The<br />
parametric approach was adopted in defining the geometry<br />
and topology of the specimen; a similar strategy was<br />
preferred also by Alqadi [11] for its analytic formulation.<br />
II. SAMPLES AND EXPERIMENTS<br />
The levitating system considered is represented by some<br />
layers of square permanent magnets and a diamagnetic<br />
square proof mass. This configuration is suitable to the<br />
fabrication of capacitive devices with magnetic suspension;<br />
for instance, Fig. 1 [8] represents a MEMS accelerometer<br />
with levitating proof mass and interdigitated comb drives<br />
detection. A similar architecture can be considered for the<br />
fabrication of diamagnetically levitating capacitive energy<br />
harvesters.<br />
The rare-earth permanent magnets are made of NdFeB<br />
and coated with Ni-Cu-Ni on the surface; they are oriented<br />
in the ‘opposite’ configuration [6] and arranged in N planar<br />
layers with four magnets each (Fig. 2). The diamagnetic<br />
material used for the levitating mass is pyrolytic graphite.<br />
The schematic configuration of the levitating system is<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
97
BACKGROUND<br />
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May 2011, Aix-en-Provence, France<br />
represented in Fig. 3 and its nominal dimensions and<br />
<br />
III. THEORETICAL<br />
material properties are listed in Table I.<br />
pyrolytic graphite density ρ 2200 kg/m 3 1⁄ 2 <br />
. (10)<br />
The measurements were performed by the optical The magnetic properties of materials are identified by the<br />
strategy; the laser sensor Keyence LK-G82 (50kHz parameters described below. The induced magnetization <br />
sampling frequency, 0.2µm±0.05% accuracy) was used to persists in the permanent magnets even if the external<br />
measure the static configuration of the system. The real magnetic field is removed; generally, it is given by<br />
thickness of every levitating mass was estimated as the<br />
<br />
average of 9 detections (1.707·10 -3 (1)<br />
average variance among<br />
where <br />
all the masses); the levitation height was also measured in<br />
is the magnetic susceptibility. The magnetic flux<br />
the different configurations of the system.<br />
density is related to by the equation<br />
(2)<br />
where is the magnetic permeability of free space<br />
(4 · 10 ⁄ ). Under the hypothesis of 1, valid<br />
for diamagnetic materials, the combination of Eqs. (1) and<br />
(2) gives the relations<br />
(3)<br />
1 (4)<br />
where 1 is the relative magnetic permeability<br />
and is the magnetic permeability.<br />
Diamagnetic materials are characterized by very small<br />
negative (that means slightly smaller than 1),<br />
paramagnetic materials have very small positive ( <br />
slightly higher than 1) and are weakly attracted by magnetic<br />
Fig. 1. Application of the magnetic suspension to real devices [8]. fields, ferromagnetic materials have large positive ( <br />
much larger than 1) and are strongly attracted by magnetic<br />
fields. Thanks to their properties, diamagnetic materials are<br />
able to generate a weak opposite field when inserted into an<br />
external magnetic field; consequently, in particular<br />
conditions, the magnetic force acting on the diamagnetic<br />
mass may balance the gravity force and produce levitation.<br />
To estimate the magnetic force, the single dipole of the<br />
diamagnetic material (e.g. atoms, molecules, ions, etc.) has<br />
Fig. 2. Image of the levitating system.<br />
to be considered. Each dipole has an individual<br />
y<br />
characteristic magnetization . The unit of volume ∆ has<br />
a magnetization<br />
z<br />
∑ <br />
∆ . (5)<br />
w<br />
The single dipole immerged in the magnetic field with<br />
o x flux density has the potential energy<br />
(6)<br />
o x l<br />
then, the elementary diamagnetic force acting on the dipole<br />
can be calculated as<br />
. (7)<br />
The diamagnetic force per unit volume is<br />
Fig. 3. Configuration of the levitating system.<br />
∑ <br />
∆ (8)<br />
and, from Eq. (3), it results<br />
TABLE I<br />
NOMINAL DIMENSIONS AND MATERIAL PROPERTIES<br />
1⁄ 2 (9)<br />
Description Symbol Value Unit<br />
IV. MODELING<br />
NdFeB magnets side w 20 mm<br />
NdFeB magnets thickness t’ 3 mm<br />
The static levitation distance of the proof mass can be<br />
NdFeB magnets layers N 1-2-3 -<br />
predicted by considering the diamagnetic force for unit<br />
NdFeB coercive force H c 860÷995 kA/m volume in the vertical direction as expressed by Eq. (9);<br />
χ<br />
NdFeB mag. susceptibility x,y -85·10 -6 -<br />
instead, the horizontal contributions are opposite in<br />
χ z -450·10 -6 -<br />
pyrolytic graphite side l 10 mm direction and self balanced. For orthotropic materials, it<br />
pyrolytic graphite thickness t 0.3-0.5-0.7-0.9-1.0 mm results<br />
98
A. Continuous domain<br />
Starting from a given configuration of permanent magnets<br />
in terms of shape, dimensions and polar orientation, the<br />
magnetic flux density can be calculated in the region of<br />
space surrounding the magnets as a 3-dimensional vector<br />
field. The diamagnetic proof mass is parallel to the plane of<br />
magnets and is situated in the same space region; its midplane<br />
is initially positioned at the distance from the<br />
magnets surface. The diamagnetic force in vertical direction<br />
can be estimated for all the points of the proof<br />
mass, assuming that each contribution of the force acts on<br />
the infinitesimal volume d d d d. The total vertical<br />
diamagnetic force at the height of the mid-plane is<br />
d<br />
(11)<br />
<br />
where is the proof mass volume. The static equilibrium of<br />
the diamagnetic proof mass is given by the following<br />
relation:<br />
(12)<br />
where is the gravity force acting on the proof<br />
mass, is the proof mass density and is the acceleration<br />
of gravity. Depending to the vertical position of the mass,<br />
the force balance may not to be verified and a recursive<br />
calculation is needed. In the further step of the calculation,<br />
the magnetic force should be evaluated at the height . The<br />
next value of the height must be assumed according to the<br />
following cases:<br />
if , then d<br />
if , then d<br />
The recursive calculation ends when <br />
at the levitation distance L .<br />
B. Discrete modeling<br />
In this case, starting from the configuration of permanent<br />
magnets, the magnetic flux density is calculated in the<br />
region of space surrounding the magnets as a 3-dimensional<br />
discrete vector field. This means that the value of is<br />
calculated only in specific points that are comparable to the<br />
nodes of a finite element model. The distribution of can<br />
be calculated in the discrete domain, starting from a given<br />
configuration of the permanent magnets, for example by<br />
using a commercial FEM simulator.<br />
The diamagnetic proof mass is parallel to the plane of<br />
magnets and is situated in the same space region; its midplane<br />
is initially positioned at the discrete distance from<br />
the magnets surface. The diamagnetic force in vertical<br />
direction can be estimated for all the discrete points<br />
of the proof mass, assuming that each contribution of the<br />
force acts on the discrete volume ∆ ∆ ∆ ∆. The<br />
central finite difference method can be used for this<br />
calculation. The total vertical diamagnetic force at the<br />
height of the mid-plane is<br />
∑∆ <br />
(13)<br />
As described before, the vertical equilibrium between<br />
diamagnetic and gravity forces has to be considered with a<br />
recursive calculation; the height of the mass at the further<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
step must be assumed according to the following cases:<br />
if , then ∆<br />
if , then ∆<br />
The recursive calculation ends when <br />
at the levitation distance L .<br />
C. Approximations<br />
The discrete modeling approach can be simplified by the<br />
following assumption: in case of thin proof mass, the<br />
variation of the diamagnetic force along the thickness can<br />
be neglected. This means that const. over the<br />
thickness. Thanks to this approximation, only the mid-plane<br />
of the proof mass can be considered in the modeling.<br />
The diamagnetic force in vertical direction can<br />
be estimated only for the discrete points of the mid-plane,<br />
assuming that each contribution of the force acts on the<br />
discrete volume ∆, which is placed across the mid-plane.<br />
The total vertical diamagnetic force at the height of the<br />
mid-plane is<br />
<br />
<br />
∆ <br />
(14)<br />
∆ ∑<br />
V. FEM SIMULATION OF MAGNETIC FIELD<br />
The distribution of the magnetic field around the<br />
permanent magnets in the described configuration ( 1)<br />
was calculated with a 3D FEM simulation by the<br />
commercial tool ANSYS 13.0. The elements solid96 were<br />
used to model the magnets and the surrounding air; the<br />
mesh size was 0.5mm and the coercive force <br />
750 kA⁄ m was assumed for the magnets. The first<br />
magnetization curve represented in Fig. 4 was used for<br />
NdFeB. The FEM model is shown in Fig. 5.<br />
The magnetic field distribution in the surrounding air was<br />
calculated; the value of at the vertical height 1mm is<br />
reported in the diagrams of Fig. 6. Due to the magnets<br />
orientation, the simulation results show the following<br />
symmetries of the magnetic field: , <br />
, | | | |.<br />
Fig. 4. First magnetization curve of NdFeB magnets.<br />
99
Fig. 5. Lower view of the FEM model with permanent magnets in the<br />
opposite configuration and air volume.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
VI. NUMERICAL CALCULATION OF STATIC LEVITATION<br />
The experimental levitation distance of the proof mass<br />
with 10mm and 1mm was assumed for the<br />
calculation of the magnetic force that was successively<br />
compared to the gravity force in order to verify the<br />
equilibrium. As described in the next section, the<br />
experimental levitation distance of the mid-plane of the<br />
proof mass in the configuration given is L, <br />
1.0687mm; the distribution of the magnetic field from the<br />
FEM model was calculated in correspondence to the vertical<br />
positions 1mm ∆ in the discretized domain, where<br />
∆ 0.5mm is the mesh size. The magnetic forces in<br />
horizontal directions are self-balanced, that reduces the<br />
equilibrium calculation at the vertical direction.<br />
Only one half of the proof mass was considered due to the<br />
symmetry of the levitating system; this part of the graphite<br />
volume was divided in several portions (1mm wide) as<br />
described by Fig. 7. The magnetic force was computed on<br />
the nodes situated along the longitudinal axis of each<br />
portion and then extended to the volume of the entire slice.<br />
Finally, the total magnetic force acting on the proof mass<br />
was calculated by adding all the contributions.<br />
The magnetic forces are calculated in correspondence to<br />
the nodes , situated along the longitudinal axis of each<br />
graphite slice (), as indicated in Fig. 7.<br />
Fig. 7. Division of the proof mass in portions and nodal magnetic force<br />
distribution.<br />
Fig. 6. Distribution of the magnetic field components at 1mm <br />
L, calculated from FEM simulation.<br />
The central finite difference method was used to compute<br />
the derivative of the magnetic flux density reported in Eq.<br />
(10). By introducing<br />
<br />
<br />
(15)<br />
it results, in the discretized domain,<br />
,<br />
<br />
,,<br />
∆<br />
(16)<br />
where the vertical index L, 1mm corresponds to<br />
coordinate of the experimental levitation height of the proof<br />
mass mid-plane and 1 L, ∆, as represented in<br />
Fig. 8.<br />
100
Fig. 8. Central finite difference method applied to the magnetic flux<br />
density.<br />
The nodal magnetic force is<br />
, <br />
,,<br />
∆<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
(17)<br />
where L, in the case considered.<br />
The magnetic force acting on each slice of the proof mass<br />
is related to the real volume of that portion of graphite.<br />
Depending to the number of nodes present on the<br />
longitudinal axis of one slice, each nodal force is referred to<br />
a small volume centered on the same node. For each<br />
graphite portion, it is related to the unit volume by the<br />
ratio<br />
<br />
<br />
<br />
(18)<br />
where is the number of nodes on the longitudinal axis.<br />
The magnetic force acting on the -th portion of the proof<br />
mass is<br />
<br />
∑ , <br />
(19)<br />
where , is the magnetic force per unit volume<br />
calculated in the node , . The total magnetic force acting<br />
on the entire proof mass is<br />
<br />
<br />
2·∑ ∑ , . (20)<br />
VII.<br />
RESULTS<br />
The levitation height was measured by the laser sensor [9]<br />
on the configuration with 1, 2, 3, nominal graphite side<br />
10mm and nominal graphite thickness <br />
0.3, 0.5, 0.7, 0.9, 1.0mm; the results, referred to the midplane<br />
of the proof mass, are reported in Fig. 9.<br />
the experimental levitation distance L, 1.0687mm in<br />
the configuration with 1 and 1mm; the measured<br />
thickness of the proof mass is 0.9117mm. The half<br />
graphite mass was divided in 7 portions ( 1, … ,7) and<br />
the coefficient for the volume correction was calculated for<br />
every portion. The magnetic force acting on each slice of<br />
graphite was calculated and compared to the corresponding<br />
value of the gravity force acting on the same portion. The<br />
mesh size used was 0.5mm that gives the number of nodes<br />
along the longitudinal axis of each portion. The results are<br />
listed in Table II.<br />
Portion<br />
<br />
Number<br />
of nodes<br />
<br />
TABLE II<br />
NUMERICAL CALCULATION RESULTS<br />
Volumetric<br />
coeff.<br />
(α)<br />
Magnetic<br />
force<br />
<br />
[mN]<br />
Gravity<br />
force<br />
<br />
[mN]<br />
<br />
<br />
1 27 0.52 0.211 0.261 0.81<br />
2 23 0.52 0.269 0.221 1.22<br />
3 19 0.53 0.090 0.181 0.50<br />
4 15 0.54 0.106 0.141 0.75<br />
5 11 0.56 0.102 0.100 1.01<br />
6 7 0.60 0.082 0.060 1.37<br />
7 3 1.00 0.016 0.020 0.80<br />
The gravity force acting on the entire graphite mass is<br />
1.970mN. The numerical calculation of the magnetic<br />
force was conducted at the levitation height of the graphite<br />
mid-plane 1mm L, . Due to the high uncertainty<br />
about the magnets coercive force, two values of <br />
representative of its range of variability were considered.<br />
For 750 kA⁄ m, the total magnetic force results<br />
1.751mN, corresponding to the error with the gravity<br />
force of about 11.1%. For 900 kA⁄ m, the total<br />
magnetic force results 2.409mN, corresponding to<br />
the 22.3% error with the gravity force. Following the first<br />
approximation of linear interpolation, the actual value of<br />
coercive force for the current magnets is about 800 kA⁄<br />
m<br />
In conclusion, the force equilibrium is demonstrated by the<br />
discrete model, with relative small errors that are<br />
addressable to (a) the experimental error in the evaluation of<br />
levitation height, (b) the approximations introduced by the<br />
discretization, (c) the uncertainty about the material<br />
properties and the magnetization curve of the magnets.<br />
Further investigations will provide more detailed results in<br />
terms of resolution of the levitation height and material<br />
parameters evaluation; different configurations of the<br />
levitating system will be considered as well.<br />
Fig. 9. Experimental levitation height of the proof mass mid-plane in<br />
different configurations: N=1 (◦), N=2 (), N=3 ().<br />
The calculation of the magnetic force was conducted at<br />
VIII. CONCLUSIONS<br />
The numerical model developed and presented in this<br />
paper was used to calculate the magnetic force acting on a<br />
proof mass of diamagnetic material levitating on permanent<br />
magnets in the ‘opposite’ configuration. Starting from the<br />
magnetic field distribution obtained with a commercial<br />
FEM simulator, the nodal magnetic force on the proof mass<br />
was calculated at the vertical position corresponding to the<br />
experimental levitation height. The resulting magnetic force<br />
was compared to the gravity force acting on the proof mass<br />
101
11-13 <br />
May 2011, Aix-en-Provence, France<br />
and the vertical equilibrium was verified quite precisely.<br />
<br />
The relatively small error was attributed to the uncertainties<br />
about materials and measurements and to the<br />
approximations of the discretized model.<br />
REFERENCES<br />
[1] W.C. Tang, M.G. Lim, and R.T. Howe, “Electrostatic comb drive<br />
levitation and control method,” J. Microelectromech. S., vol. 1, pp.<br />
170-178, 1992.<br />
[2] S.W. Chyuan and Y.S. Liao, “Computational study of the effect of<br />
finger width and aspect ratios for the electrostatic levitating force<br />
of MEMS combdrive,” J. Microelectromech. S., vol. 14, pp. 305-<br />
312, 2005.<br />
[3] E.M. Yeatman, “Applications of MEMS in power sources and<br />
circuits,” J. Micromech. Microeng., vol. 17, pp. S184-S188, 2007.<br />
[4] G. De Pasquale, C. Siyambalapitiya, A. Somà, and J. Wang,<br />
“Performances improvement of MEMS sensors and energy<br />
scavengers by diamagnetic levitation,” proc. of ICEAA, Torino,<br />
Italy, pp. 465-468, 2009.<br />
[5] H. Chetouani, B. Delinchant, and G. Reyne, “Efficient modelling<br />
approach for optimization of a system based on passive<br />
diamagnetic levitation as a platform for bio-medical applications,”<br />
in Compel, vol. 26, Emeral Group <strong>Publishing</strong>, 2007, pp. 345-355.<br />
[6] F. Barrot, “Acceleration and inclination sensors based on magnetic<br />
levitation. Application in the particular case of structural health<br />
monitoring in civil engineering,” PhD thesis, EPFL, Lausanne,<br />
Switzerland, 2008.<br />
[7] C. Elbuken, M.B. Khamesee, and M. Yavuz, “Eddy current<br />
damping for magnetic levitation: downscaling from macro- to<br />
micro-levitation,” J. Phys. D: Appl. Phys., vol. 39, pp. 3932-3938,<br />
2006.<br />
[8] D. Garmire, H. Choo, R. Kant, S. Govindjee, C.H. Séquin, R.S.<br />
Muller, and J. Demmel, “Diamagnetically levitated MEMS<br />
accelerometers,” proc. of Transducers and Eurosensors, Lyon,<br />
France, pp. 1203-1206, 2007.<br />
[9] G. De Pasquale, C. Siyambalapitiya, S. Iamoni, and A. Somà,<br />
“Characterization of low-stiffness suspensions based on<br />
diamagnetic levitation for MEMS energy harvesters,” proc. Power<br />
MEMS, Leuven, Belgium, pp. 77-80, 2010.<br />
[10] H. Chetouani, V. Haguet, C. Jeandey, C. Pigot, A. Walther, N.M.<br />
Dempsey, F. Chatelain, B. Delinchant, and G. Reyne,<br />
“Diamagnetic levitation of beads and cells above permanent<br />
magnets,” proc. of Transducers and Eurosensors, Lyon, France,<br />
pp. 715-718, 2007.<br />
[11] M.K. Alqadi, H.M. Al-Khateeb, F.Y. Alzoubi, and N.Y. Ayoub,<br />
“Effects of magnet size and geometry on magnetic levitation<br />
force,” Chin. Phys. Lett., vol. 24 p. 2664, 2007.<br />
102
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Molecular Dynamic Simulation of Nanoparticle Size<br />
Effect on Melting Point of Gold<br />
P. Nayebi 1,2 , M.Shamshirsaz 2 , K. Mirabbaszadeh 1 , E. Zaminpeyma 1 , M.B. Asgari 3<br />
1 Physics Department, 2 New Technologies Research Center, Amirkabir University of Technology (Tehran Polytechnic),<br />
4 Niroo Research Institute<br />
4244 Hafez Ave., P.B. 15875-4413. Tehran, Iran<br />
E-mail: shamshir@aut.ac.ir<br />
Abstract- Molecular dynamics simulation of<br />
the crystallization<br />
behavior of liquid gold (Au) nanoparticles, with 1, 2, 3, 4, 5 and 6 nm<br />
in diameter, on cooling has been carried out based on the embeddedthe<br />
structure of the<br />
atom-method potential. It is demonstrated that<br />
fully crystallized particle is polycrystalline face-centered cubic<br />
(FCC). The FCC structure of the gold nanoparticle is proved<br />
energetically the most stable form. The final structure of<br />
nanoparticles are affected by cooling time and size of nanoparticles.<br />
Increasing the size of nanoparticles, the melting point of<br />
nanoparticles will be increased.<br />
Keywords: Molecular Dynamic Simulation, Melting point,<br />
Gold nanoparticle, Crystallization<br />
I. INTRODUCTION<br />
Atomistic simulation techniques such as molecular dynamics<br />
(MD) have become a powerful tool in the field of<br />
nanotechnology as they provide a physical insight in<br />
understanding various phenomena on atomic<br />
scale and enable<br />
one to predict some properties of nanomaterials. It is expected<br />
that atomistic simulation could gradually play a key role in<br />
complementing experiments in the field of nanomaterials,<br />
because the experimental exploration of nanomaterials will<br />
inevitably encounter many technical difficulties not to mention<br />
the cost problems.<br />
In the present paper we report the results of molecular<br />
dynamic simulations of the crystallization<br />
of molten gold<br />
nanoparticles with an embedded atom method for potential as<br />
used previously by the authors [1,2].<br />
II. SIMULATION MODEL<br />
Molecular dynamic simulations are effective tools for<br />
understanding the melting process of finite systems at the<br />
atomistic level. MD simulations weree performed on<br />
unsupported spherical face-centered cubic (fcc) gold particles,<br />
with 1,2,3,4,5 and 6nm in diameter, containing 6699 atoms for<br />
6nm-particle, 3925 atoms for 5nm-particle, 1985 atoms for<br />
4nm-particle, 887 atoms for 3nm-particle, 249 atoms for 2nm-<br />
verlet algorithm<br />
particle and 43 atoms for 1nm-particle. The<br />
for calculating the simulation parameters is applied. The<br />
particle was subjected to a periodic boundary condition in all<br />
directions at given temperatures<br />
under zero external pressure.<br />
The EAM potential is an inter-atomic potential developed by<br />
Daw and Baskes [3, 4] for metals. A simulation box with<br />
dimensions of 300* 300* 300A˚˚ is created; hence the periodic<br />
boundary condition will not affect particle size changing.<br />
Furthermore, this simulation box<br />
is very larger than the cut off<br />
radius of EAM potential. The particles are placed in the center<br />
of simulation box. Au nanoparticles are generated by cutting a<br />
sphere from bulk crystals in its<br />
equilibrium state[11]. As a<br />
sample, the 4nm-particle is employed in simulation with VMD<br />
software as shown in Fig. 1.<br />
Fig.1. A sphere of 4nm- Au nanoparticle, generated by cutting a sphere from<br />
bulk crystal<br />
As a first step, the MD run 50ps was performed with a time<br />
step of 1fs at 2000K in a canonical ensemble using the<br />
Langevin Temperature thermostat, which is slightly above the<br />
melting temperature of bulk Au, to melt the whole particle.<br />
Then, temperature was decreased to 50K with Langevin<br />
temperature thermostat. In the present study, the MD<br />
simulations were carried out with four different cooling times:<br />
50,170, 500 and 1000ps.<br />
III. RESULTS AND DISCUSSION<br />
In order to investigate what kind of Au structure can be<br />
observed just after solidification, the internal energy of<br />
particles is calculated. The variation of the internal energy of<br />
the 4nm-Au nanoparticle with temperature for different<br />
cooling time is presented in Fig. 2.<br />
103
Internal Energy(eV)<br />
-7100<br />
50 150 250 350 450 550 650 750 850<br />
-7150<br />
-7200<br />
-7250<br />
-7300<br />
500ps<br />
170ps<br />
-7350<br />
50ps<br />
-7400<br />
-7450<br />
1000ps<br />
Temprature(K)<br />
Fig. 2. Internal energy of 4nm nanoparticle versus temperature<br />
with different cooling time.<br />
All of the internal energies decrease similarly as the<br />
temperature decreases down to about 550- 600 K, regardless<br />
of the cooling time. However, the internal energies show a<br />
quite different feature below almost 500 K. For the cooling<br />
time of 50ps, the internal energy decreases almost linearly<br />
with decreasing temperature. While for the cooling time 1000<br />
ps and same temperature range, internal energy drops slower<br />
compare to 50 ps and 170ps cooling time. The final internal<br />
energies at 50K tends to decrease as cooling time decreases,<br />
indicating the different phase transformation routes<br />
experienced with temperature. Nanoparticle at cooling time<br />
1000ps are more stable, because it has smaller internal energy.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Fig. 4 shows the atomic arrangements of the 4nm gold<br />
nanoparticle at 50K cooled with a different cooling time. It<br />
can be seen that the particle cooled with 50ps has completely a<br />
random arrangement. Also, the particles cooled with 170 and<br />
500ps show a random arrangement, though the partial<br />
crystalline feature in the outer shell can be seen in Fig. 4(b, c).<br />
This is while the particles cooled at a rate of 1000ps have a<br />
more crystallized structure. Fig. 5 shows the variation of the<br />
internal energy of the Au nanoparticles with temperature<br />
during cooling time of 1000ps.<br />
However the whole of internal energy decreases, but it drops<br />
at different temperature. This shows that melting point of gold<br />
nanoparticles is changed with their sizes. The physical<br />
properties of nano-particles are expected to deviate from those<br />
of the bulk metal because of both on account of their vastly<br />
increased ratio of surface atoms to internal ones and of course<br />
as the result of their different electronic structure. The melting<br />
point of Au nanoparticles at cooling time of 1000ps for 1-6nm<br />
is demonstrated in Fig 6.<br />
(a)<br />
(b)<br />
Fig 3 shows the surface morphologies of 4nm gold<br />
nanoparticles at 50K and cooled with a different cooling time.<br />
The surface atoms of the particle cooled with 50ps exhibit a<br />
random arrangement, whereas those of the particles with<br />
1000ps show a crystalline feature.<br />
(c)<br />
Fig. 4. Atomic arrangement of 4nm Au nanoparticle with<br />
different cooling time: a) 50ps b) 170ps c) 500ps d) 1000ps<br />
(d)<br />
-3100<br />
50 150 250 350 450 550 650 750 850 950<br />
-3120<br />
-3140<br />
-3160<br />
(a)<br />
(b)<br />
-3180<br />
-3200<br />
-3220<br />
-3240<br />
-3260<br />
-3280<br />
-3300<br />
(c)<br />
(d)<br />
Fig. 3: Surface morphology of 4nm-Au with different cooling<br />
time: a) 50ps b) 170ps c) 500ps d) 1000ps<br />
Fig.5. Variation of the internal energy of the Au nanoparticles<br />
with temperature during cooling with the time of 1000ps.<br />
104
700<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Melting Point(K)<br />
650<br />
600<br />
550<br />
500<br />
450<br />
400<br />
350<br />
300<br />
1 2 3 4 5 6<br />
Nanoparticle Size(nm)<br />
Fig 6. Melting point of Au nanoparticles with a cooling time<br />
of 1000ps for 1-6nm<br />
In Fig 7, the melting points of Au nanoparticles encapsulated<br />
in silica obtained experimentally [12, 13] are compared with<br />
the simulation results obtained here for different particle size.<br />
Both of these results show a reduction of melting point with a<br />
decrease in size of gold nano-particles. As it can be seen,<br />
however the trends of curves are similar but there exist a<br />
deviation of simulation results from experimental data for the<br />
melting point of each range of nanoparticles sizes. As the<br />
example, the melting point of Au particles with the 5 nm size<br />
range (~Au3600) is found by simulation about 830°K in<br />
literature while for the almost same size nanoparticles<br />
(~Au3925), a melting temperature of about 600K is calculated<br />
here. This difference in melting temperature can be due to the<br />
different number of atoms used in MD. Also, the nanoparticles<br />
of 2 nm size (about Au200) and by extrapolation the<br />
nanoparticles of 1 nm size (~Au30) ,become liquid at 550°K<br />
and 430 °K respectively in literature while in this study, it is<br />
found a melting temperature of about 500K for 2nm size<br />
(~Au249) and 400K for 1nm size (~Au43) particles.<br />
Fig 7. Melting point of Au nanoparticles versus particle<br />
diameter obtained by a) MD simulation b) experimental data<br />
in literature [12,13]<br />
IV. CONCLUSIONS<br />
Molecular dynamic simulation of the crystallization<br />
behavior of a liquid gold (Au) nanoparticle, with 1, 2, 3 and 4<br />
nm in diameter, on cooling has been carried out.<br />
It has been found that the final structure of a gold nanoparticle<br />
strongly depends on cooling time during crystallization from<br />
liquid. The increase of cooling time from 50ps up to 1000ps<br />
makes the structure of nanoparticle closer to FCC structure.<br />
As size of particle increases from 1nm to 4nm, the final<br />
structure of the particle tends more to a crystalline structure.<br />
The internal energies of the particles for cooling time of<br />
1000ps drop slower comparing to those related to 50ps and<br />
170 cooling time. This means that the atoms in the outer shell<br />
tend faster to become a crystalline structure.<br />
The crystallization of the surface atoms has a greater<br />
contribution to the decrease in internal energy comparing to<br />
the core crystallization. It is confirmed that the FCC structure<br />
is energetically the most stable form for a gold nanoparticle.<br />
The internal energy of nanoparticle at cooling time of 1000ps<br />
is smaller than the other cooling time rate simulated in this<br />
work. Increasing the size of nanoparticles, the melting point of<br />
nanoparticles will be increased.<br />
REFERENCES<br />
[1] K.Mirabbaszadeh,E. Zaminpeyma, P. Nayebi,J Cluster Sci<br />
(2008)19,623-629<br />
[2] K. Mirabbaszadeh, P.Nayebi, E. Zamipeyma, J Cluster Sci<br />
(2008)19,661-670<br />
[3] M.S. Daw, M.I. Baskes, Semiempirical, quantum<br />
mechanical calculation of hydrogen embrittlement in metals,<br />
Phys. Rev. Lett. 50 (1983) 1285–1288.<br />
[4] M.S. Daw, M.I. Baskes, Embedded-atom method:<br />
derivation and application to impurities, surfaces, and other<br />
defects in metals, Phys. Rev. B 29(1984) 6443–6453.<br />
[5] M.S. Daw, M.I. Baskes, Phys. Rev. B 29 (1984) 6443.<br />
[6] S.M. Foiles, M.I. Baskes, M.S. Daw, Phys. Rev. B 33<br />
(1986) 7983.<br />
[7] A.E. Carlsson, in: H. Ehrenreich, D. Turnbull (Eds.), Solid<br />
State Phys., vol. 43, Academic Press, New York, 1990.<br />
[8] M.S. Daw, S.M. Foiles, M.I. Baskes, Mater. Sci. Rep. 9<br />
(1993) 251.<br />
[9] C.L. Cleveland, U. Landman, T.G. Schaaff, M.N.<br />
Shafigullin, P.W. Stephens, R.L. Whetten, Phys. Rev. Lett. 79<br />
(1997) 1873;M.D. Wolf, U. Landman, J. Phys. Chem. A 102<br />
(1998) 6129.<br />
[10] J.W.M. Frenken, P. Stolze, Phys. Rev. Lett. 82 (1999)<br />
3500.<br />
[11] M.S. Daw, M.I. Baskes, Phys. Rev. Lett. 50 (1983)1285.<br />
[12] H.B. Liu, J.A. Ascencio, M. Perez-Alvarez, and M.J.<br />
Yacaman: Surface Science, 2001, vol.491,pp.88-98.<br />
[13] K. Dick, T. Dhanasekaran, Z. Xhang, and D. Meisel: J.<br />
Am. Chem. Soc, 2002, vol.124(10),pp.2312-2317<br />
105
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Stress Identification of Thin Membrane Structures by<br />
Dynamic Measurements<br />
Steffen Michael 1 , Christoph Schäffel 1 , Sebastian Voigt 2 , Roy Knechtel 3<br />
1 IMMS GmbH, Ehrenbergstr. 27, 98693 Ilmenau, Germany<br />
2 TU Chemnitz, Chair in Microsystems and Precision Engineering, Reichenhainer Str. 70, 09107 Chemnitz, Germany<br />
3 X-FAB Semiconductor Foundries AG, Haarbergstr. 67, 99097 Erfurt, Germany<br />
A fast identification method of membrane stresses is<br />
investigated for an early stage of the manufacturing process.<br />
The approach consists of performing optical measurement of<br />
the out-of-plane modal responses of the membrane. This<br />
information is used in an inverse identification algorithm<br />
based on a FE model by an optimization.<br />
I. INTRODUCTION<br />
The development of the two criteria costs and reliability<br />
is essential for the further growth of the MEMS market like<br />
microphones. Efficient test procedures on wafer level can<br />
reduce costs significantly by the detection of faulty sensors<br />
before the subsequent packaging and assembly steps. The<br />
presented method deals with an approach for a fast and<br />
accurate stress identification of thin membranes by using<br />
the sensitivity of their modal frequencies versus the stress.<br />
MEMS devices usually do not permit direct parameter<br />
measurement of mechanical parameters. The indirect<br />
parameter identification by modal frequencies was first<br />
presented in [1], [2]. Up to now the approach is used mostly<br />
for the identification of geometrical parameters like<br />
membrane thicknesses [3], [4]. In this case the approach<br />
competes against other methods like optical ones. In<br />
contrast the method has a unique feature with regard to the<br />
identification of tensile stressed membranes like<br />
microphones – another non-destructive method on wafer<br />
level is not known.<br />
Perforated circular SiN membranes with a thickness of<br />
300 nm and a diameter of 1000 µm are investigated. The<br />
perforation is required by the technology – the membrane<br />
structure is deposited on a sacrificial layer which is<br />
removed at the end of the processing through the<br />
perforation holes. The formed cavity with a height of 1µm<br />
causes a squeeze film damping in conjunction with an<br />
absent resonance rice under ambient atmosphere.<br />
Correspondingly the measurements are done in a vacuum<br />
prober.<br />
II. HARDWARE SETUP<br />
The measurement setup consists on a vacuum probe<br />
station from Cascade and a laser Doppler vibrometer<br />
integrated in the Micro System Analyzer MSA500 from<br />
Polytec. The laser beam of the vibrometer scans<br />
automatically over a user defined grid at the surface of the<br />
membrane.<br />
Fig. 1: Measurement setup<br />
The vibration of passive devices like the membrane<br />
structures is realized by electrostatic forces. A probe needle<br />
is connected to a high voltage (up to 400V) excitation signal<br />
controlled by a chirp signal of the measurement system. The<br />
needle is positioned above the device surface. With respect<br />
to a high excitation force the gap between the needle and<br />
the membrane is smaller than 100µm. The setup permits the<br />
excitation of modal frequencies up to 4MHz.<br />
III. IDENTIFICATION ALGORITHM<br />
The approach can be subdivided into three different<br />
phases. First of all a sensitivity analysis has to be done to<br />
check whether the modal frequencies are sensitive versus<br />
the interesting parameters. In case of a multidimensional<br />
problem the orthogonality of the parameter space has to be<br />
tested furthermore.<br />
Following to the sensitivity analysis a characterization<br />
phase is done. Frequency response functions (FRF) are<br />
measured with a fine grid of measurement points to check<br />
the mode shapes and adapt the finite element (FE) model if<br />
needed.<br />
The results shown here refer to measurement data of the<br />
characterization phase. In case of testing complete wafers<br />
the measurement time should be minimized. The<br />
measurement time depends proportional on the number of<br />
measurement points. The identification approach is based<br />
on frequency values which permits the reduction of<br />
measurement points to one.<br />
106
o<br />
o<br />
o<br />
o<br />
o<br />
o<br />
o<br />
o<br />
Check applicability<br />
Analytic / FE- modelling<br />
Check sensitivity<br />
Check orthogonality<br />
Development of test structures<br />
Characterization<br />
Fine grid of measurement points<br />
Selection of frequency modes for<br />
identification<br />
Parametet identification & validation<br />
Adaption of FE model<br />
Wafer-Test<br />
Fig. 2: Phases of the parameter identification<br />
The measurement time of a one point measurement is 2<br />
seconds. The measurement respectively software system is<br />
not yet optimized, the lower measurement time limit given<br />
by physics is about 200 milliseconds.<br />
Precondition for the identification is on one hand the<br />
measurement unit which delivers a FRF, and the simulation<br />
unit with a parameter matrix as result on the other hand. The<br />
automatic identification is done by a tool implemented in<br />
C++ with respect to a fast data processing. The<br />
identification tool can be structured into three submodules.<br />
The frequency values has to be extracted from the measured<br />
FRF which is done in one submodule, and the parameter<br />
matrix is approximated by usually polynomials in another<br />
submodule due to a fast and efficient data handling. Based<br />
on an user defined accuracy (default value 0.1%) the degree<br />
of the polynomial is selected by the program.<br />
Finally the optimization respectively identification is<br />
realized by the nonlinear least square method.<br />
Measurement<br />
system<br />
Frequency<br />
response<br />
Peak detection<br />
Identification tool<br />
Optimization<br />
FE-Simulation<br />
Polynomial<br />
approximation<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
first step a conventional algorithm searches for local maxima<br />
considering the estimated signal-to-noise ratio (SNR). At the<br />
peaks found, starting values for a nonlinear least square fit to<br />
the Lorentzian function<br />
Parameter<br />
matrix<br />
i<br />
2<br />
, ih<br />
LfL<br />
, ip 2 2<br />
,<br />
)(<br />
+−<br />
, i hi<br />
)(<br />
f<br />
= (1)<br />
fff<br />
with the peak amplitude L p,i , the peak frequency f p,i and the<br />
half-width f h,i. of the ith peak are estimated. The iterative<br />
fitting procedure based on Levenberg-Marquardt algorithm<br />
eliminates wrongly preselected peaks and delivers the peak<br />
parameter including the quality factor.<br />
A. FE Modeling and Simulation<br />
The FE model which delivers the parameter matrices is<br />
implemented in Ansys. The ratio thickness to lateral<br />
dimension of the membrane leads to a modeling by twodimensional<br />
shell elements. The default mesh of the<br />
membrane perforated by several thousand holes will be<br />
irregular. To prevent such an inefficient irregular mesh<br />
substructures are generated. Square areas with a centered<br />
hole permit a regular meshing.<br />
Fig. 4: FE modell with prestructured membrane elements<br />
A prestressed modal analysis as well as a prestressed<br />
harmonic analysis is performed. The multitude of small<br />
structures causes a large number of finite elements<br />
respectively nodes. With regard to the measurement time the<br />
membrane symmetry is used by the calculation of a quarter<br />
model. Symmetric boundary conditions are applied to the<br />
static analysis. The modal analysis is executed with three<br />
load steps with different symmetry conditions at the x and y<br />
axes (symmetric/symmetric, asymmetric/symm. and<br />
asym./asym.) to deliver all modal frequencies .<br />
For the modeling of the squeeze film damping the<br />
corresponding element types of Ansys are used. The macro<br />
RMFLVEC.MAC which extracts the damping parameters<br />
from the modal frequencies is adapted to the quarter model<br />
with the multiple loadsteps.<br />
Sensor parameter<br />
Fig. 3: Structure of the parameter identification<br />
From the measured FRF, the peak frequency values are<br />
extracted automatically by a two level algorithm. Within a<br />
a) f 11 b) f 12<br />
Fig. 5: Simulated modal frequencies<br />
107
IV.<br />
A. Simulation and Sensitivity Analisys<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
STRESS IDENTIFICATION<br />
With respect to the identification phases a sensitivity<br />
analysis is done for the membrane structure. Parameters<br />
which have to be considered beside the interested ones are<br />
parameters with relevant tolerance ranges. The membrane<br />
thickness is such a parameter – due to technological reasons<br />
the thickness varies within a range of ±5%.<br />
(∂ f 1<br />
/∂ h) ∆ h/f 1<br />
[%]<br />
6<br />
5<br />
4<br />
3<br />
2<br />
1<br />
f 1<br />
[kHz]<br />
60<br />
50<br />
40<br />
30<br />
20<br />
10<br />
0<br />
0.42<br />
Fig. 6: First modal frequency versus membrane thickness and stress<br />
As is apparent from Fig. 6 which show the results of the<br />
two dimensional parameter simulation for the first modal<br />
frequency the most sensitive parameter is the stress. An<br />
approximation of the functional dependency is done with<br />
regard to a quantitative analysis. The default expansion is a<br />
polynomial one. In this case rational functions are used for<br />
the stress motivated by the plate theory [5] on the one hand<br />
and the curve characteristic of Fig. 6 on the other hand The<br />
frequency mode f i,j is given by<br />
with the membrane thickness h and the stress s.<br />
Based on partial derivatives of the approximated course<br />
of the function the sensitivity of the modal frequencies<br />
versus the parameters is determined. Fig. 7 shows the<br />
sensitivity normed on the maximum thickness variation of<br />
5%. In case of a tensile stressed membrane the varying<br />
thickness can be neglected – a relevant sensitivity of the<br />
modal frequencies versus the thickness is given only in case<br />
of a stress-free or compressive stressed membrane.<br />
0<br />
0 2 4 6 8 10 12 14 16 18 20<br />
s [MPa]<br />
Fig. 7: Normed sensitivity of the first modal frequency versus membrane<br />
thickness<br />
B. Measurement Results<br />
Measurements are done at three different wafers at a<br />
pressure range between 0.005 mbar and 0.1 mbar. The<br />
pressure range is determined by the resonance rice on one<br />
hand and a minimal peak width to be detectable by the FFT<br />
on the other hand. The measured quality factors show a<br />
good accordance with the simulated ones given by the<br />
harmonic analysis of the FE model.<br />
0.41<br />
20<br />
0.4<br />
15<br />
10<br />
0.39<br />
5<br />
z [µm] 0.38 0<br />
s [MPa]<br />
10 5 measurement data<br />
simulated data<br />
10 4<br />
10 3<br />
2/1<br />
ji 1,<br />
2<br />
++=<br />
3<br />
),(),(),(<br />
sjipsjipjip f<br />
3/1<br />
+<br />
4<br />
5<br />
6<br />
⋅++<br />
),( shjiphjip sjip 10 2<br />
(2)<br />
10 -5 10 -4 10 -3 10 -2 10 -1 10 0<br />
2/1<br />
3/1<br />
p [mbar]<br />
7<br />
8<br />
),(<br />
⋅+⋅+<br />
shjipshjip<br />
Q-factor<br />
Fig. 8: Q-factor versus ambiance pressure<br />
The first three modal frequencies are used for the<br />
identification of the membrane stress. Mode shapes are<br />
investigated at some samples to guarantee the right<br />
classification of the frequency peaks to the corresponding<br />
modes.<br />
Fig. 9: Measured mode shape f 1,1<br />
108
C. Identification Results<br />
The identified tensile stresses at 36 measured dies vary<br />
between 24MPa and 81MPa due to their different position at<br />
the test wafers.<br />
TABLE 1<br />
IDENTIFIED STRESS OF MEMBRANE SAMPLES<br />
f 1,1<br />
[kHz]<br />
f 1,2<br />
[kHz]<br />
f 2,2<br />
[kHz]<br />
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<br />
s r [MPa]<br />
55.2 88.1 117,3 28.75 ± 0.33<br />
60.9 97.2 130.5 34.99 ± 0.16<br />
69.9 109.7 147.1 45.57 ± 0.98<br />
69.1 110.2 148.0 45.50 ± 0.10<br />
[4] Michael, S. at al, “MEMS parameter identification on wafer level<br />
using laser Doppler vibrometry”, Smart Systems Integration 2007,<br />
Editor T.Gessner, VDE Verlag, 2007, pp. 321-328<br />
[5] Dickinson, S.M., “The Buckling and Frequency of Flexural<br />
Vibration of Rectangular Isotropic and Orthotropic Plates Using<br />
Raleigh’s Method”, Journal of Sound and Vibration, 1978, 61(1),<br />
pp. 1-8<br />
The identification is based on the first three modal<br />
frequencies which results in an over-determined problem<br />
which permits a quantitative evaluation of the identification<br />
results. Theoretically the stress values should be identically;<br />
practically measurement and modeling errors will cause<br />
different values. The particular stress values differ within a<br />
range of 2% which shows a good model quality.<br />
90<br />
80<br />
Wafer 1<br />
Wafer 2<br />
Wafer 3<br />
70<br />
s r<br />
[MPa]<br />
60<br />
50<br />
40<br />
30<br />
20<br />
0 20 40 60 80 100<br />
Die index<br />
Fig. 10: Identified stress at test wafers across the x axes<br />
V. CONCLUSION<br />
We have presented an approach for the fast and accurate<br />
stress identification of thin membranes which the uses the<br />
sensitivity of their modal frequencies versus stress. The<br />
approach is well suited for an efficient process control of<br />
stress sensitive membranes like microphones on wafer level.<br />
REFERENCES<br />
[1] Smith, N.F. et al, “Non-Destructive Resonant Frequency<br />
Measurement on MEMS Actuators”, 39 th Annual International<br />
Reliability Physics Symposium, Orlando, FL, USA, 2001,<br />
Proceedings, pp. 99-105<br />
[2] Tanner, D.M. et al: “Resonant frequency method for monitoring<br />
MEMS fabrication”, Reliability, Testing and Characterization of<br />
MEMS/MOEMS II, San Jose, CA, USA, 2003, Proceedings, pp.<br />
220-228<br />
[3] Gerbach, R. et al: „Numerical Identification of Geometric<br />
Parameters from Dynamic Measurement of Grinded Membranes<br />
on Wafer Level”, 7 th Conf. on Thermal, Mec.l and Multiphysics<br />
Simulation and Experiments in Micro-Electronics and Micro-<br />
Systems EuroSimE, Como, Italy, 2006, Proceedings, pp. 223-228<br />
109
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<br />
Meso-Scale Actuator Design For The Integrated Dynamic<br />
Alignment Of A Lenslet Array Within a Package<br />
Stefan Wilhelm, Robert W. Kay, Marc P.Y. Desmulliez<br />
Microsystems Engineering Centre (MISEC),<br />
Institute for Integrated Systems (IIS),<br />
School of Engineering and Physical Sciences, Heriot-Watt University,<br />
Edinburgh EH14 4AS, Scotland, United Kingdom<br />
Tel: +44 (0)131-451-8316<br />
Keywords- LTCC, actuator, packaging, optical lenses<br />
Abstract- This paper describes the design of an LTCCprocess<br />
compatible meso-scale actuator for the six degrees of<br />
freedom dynamic adjustment of micro-optical components, in<br />
particular the alignment of a microlens array on top of a UV-<br />
LED array. The lens array is specified to have an active area of<br />
3mm x 3mm, the GaN array is 5mm x 5mm x 450!m. The focal<br />
length is 65!m. The actuator must enable the collimation or<br />
the focusing of the optical beams emanating from the LED array.<br />
INTRODUCTION<br />
A great variety of micro-devices encompasses multiple<br />
interacting electronic, electro-mechanical, electrochemical<br />
or optoelectronic components that require to be aligned statically<br />
or dynamically (real-time). Static alignment can be<br />
achieved with the help of high precision pick-and-place machines<br />
with control feedback combined with a bonding process,<br />
such as U.V. curable glue or flip-chip bonding using<br />
reflowed solder balls. There are however instances where an<br />
alignment has to be performed after the sealing of the package.<br />
In such cases, structures with temporary actuation functionalities<br />
designed within the micro-devices can be exposed<br />
to external fields, providing thereby precise positioning<br />
with the help of external or temporary internal feedback.<br />
Dynamic alignment requires the manufacturing of permanent<br />
actuators within the device and must fit the requirements<br />
for power consumption, response time, force, deflection<br />
range and long term reliability. Conventionally, the<br />
function of the package is to provide electrical interconnection,<br />
heat transfer and protection against mechanical, electromagnetic<br />
and chemical influences. The additional ability<br />
of the package to provide actuation and feedback elements<br />
for aligning statically or dynamically opens up interesting<br />
opportunities for new applications such as the microscope<br />
on a chip, and greater ease of packaging by relaxing positioning<br />
tolerances at the assembly stage. This paper aims to<br />
offer an example of such a meso-sale actuation for optoelectronic<br />
application using Low Temperature Cofired Ceramics<br />
(LTCC). In that respect, a MEMS post process based solution<br />
for the alignment of the microlens array has already<br />
been reported in [1].<br />
LTCC is an established multi-layer-process, which enables<br />
to integrate electrical, fluidic or optical interconnections<br />
and passive circuit components together with mechanical<br />
structures in one solid ceramic body. Applications include<br />
electrical packaging, RF-systems, micro-fluidics [2],<br />
sensors [3] and actuators [4]. The 3D laminated device is<br />
composed of paper-thin flexible sheets consisting of alumina,<br />
glass and organic binders [5]. These so-called green<br />
sheets can encompass layer-interconnection vias, cavities<br />
and flexures, whose patterning can be performed using laser-machining,<br />
powder blasting [6], punching and embossing<br />
[7]. Metal tracks, resistors, solder masks, sacrificial inlays,<br />
high-! materials and magnetic components like ferrite<br />
[8] can be applied using thick-film screen-printing. After<br />
being separately processed, the layers are laminated and cofired<br />
into a single body at temperatures of approximately<br />
900ºC.<br />
The variety of applicable materials and the standardized<br />
process make LTCC a preferred candidate for a meso-scale<br />
actuator. The challenge is to devise a low-cost LTCCprocess<br />
compatible design, which compensates for the accuracy<br />
limits of the process whilst satisfying the requirements<br />
of maximum stroke of 10!m for the application envisaged.<br />
DESIGN OF THE PACKAGE ACTUATOR<br />
Six degrees of freedom actuation of the optical system requires<br />
the generation of translational forces and momentums<br />
for three linear independent axes. In macro manipulators,<br />
this is often realized by cascading independent polar<br />
and linear axes. As complex three-dimensional structures<br />
increase significantly the complexity of the LTCC manufacturing<br />
process, actuation elements and restoring force elements<br />
were selected, which can be placed “in plane” by<br />
structuring single layers using screen and stencil printing.<br />
Hence, the device requires planar actuation elements that<br />
generate lateral and vertical forces. The optical system can<br />
be rotated and tilted by generating these forces at a specified<br />
distance of the rotation/tilt centre point.<br />
110
The proposed solution for this problem consists of three<br />
functional LTCC layers and two sacrificial layers. The vertical<br />
and lateral forces are generated using electrostatic and<br />
magnetic actuation, respectively. This concept allows the<br />
integration of the actuator in a conventional LTCC circuit.<br />
Additional process steps include via filling, metallization,<br />
screen printing of sacrificial material and mechanical structuring.<br />
For the manufacturing of the prototype, the self constrained<br />
(“zero shrinkage”) system HeraLock HL2000<br />
(91!m cofired thickness) was selected. The appropriate<br />
metallization pastes are TC0302 for tracks and TC0303 for<br />
vias. The sacrificial material PST-CARB-SP is based on<br />
nano carbon particles and produced by the company Thick<br />
Film Technologies TM .<br />
LTCC layers, structured elements, metallization and sacrificial<br />
layers were implemented as parametrical components<br />
(“sheet parts”) with the 3D CAD software Inventor.<br />
This software provides useful features for parametric designing,<br />
assembly collision checks, as well as file formats<br />
that can be imported by CAM systems. Laser manufacturing<br />
files, powder-blasting masks, via stencil and screen design<br />
are automatically updated when model parameters change.<br />
A. Functional design overview<br />
The actuated part (“rotor”) of the meso-scale device is a<br />
hexagonal area layer in the centre of the device, which is<br />
mechanically structured and double side printed with metallization<br />
as shown in figure 1. The actuator contains a cutout<br />
for the insertion of the optical system, two 80º buckled<br />
beam flexures to provide the restoring forces and electrical<br />
connections to the bulk of the device (“stator”), and three<br />
120º shifted actuator units, each consisting of two planar<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
coil pairs.<br />
For the generation of the lateral forces, the six coil pairs<br />
create a magnetic field in response to a current passing<br />
through the windings. The underside and the topside coils<br />
are connected in the centre of each coil using vias, and with<br />
the outer bulk using the topside and underside metallization<br />
of the flexures. The bottom side coils are mirrored, so that<br />
driving currents generate collinear Lorentz-forces when an<br />
external magnetic field is applied as shown in figure 2. The<br />
external stator fields are provided by nickel-plated neodymium<br />
magnet pairs (each magnet 5mm x 1.5mm x 1mm, 1.3<br />
T), which are placed after the cofiring step above and below<br />
the coils. As each actuation unit includes two 90º shifted<br />
coils, the geometric sum of the force vectors of all independent<br />
driven coils enable to generate the lateral forces in<br />
x and y, as well as a rotation around the z-axis.<br />
The vertical electrostatic actuation is generated with the<br />
surfaces of the twelve coils. The top and bottom counter<br />
electrodes are placed at a distance of 25!m form the coils.<br />
Carbon sacrificial material is situated between the electrode<br />
layer and the centre layer to provide the gap. To increase the<br />
magnitude of the forces, the coils surfaces have been extended<br />
by filling the inner areas and attaching rectangular<br />
areas at the outer winding tracks. The combination of all independent<br />
pull forces enables deflections in the z-direction<br />
and tilt around the x- and y-axes.<br />
B. Combined actuator principle<br />
According to the dimensions of the microlens array, the<br />
device has to move a load of 248!N (volume density 2.25<br />
g/cm 3 ) in addition to the actuator mass. The applied actuation<br />
methods were chosen out through the careful study of<br />
Fig. 1: Top view (x-y) of centre of the double sided metalized layer. Three<br />
combined electrostatic/magnetic actuation units enable the creation of<br />
lateral and vertical forces. 80º buckled flexures connect the coils with the<br />
bulk vias. The mechanical structuring includes fiducials for the screen<br />
printing process, lamination alignment holes and cut-outs to provide access<br />
to the solder pads on the bottom side layer.<br />
Fig. 1: 3D visualization of the functional elements: The LTCC material at<br />
the centre of the layer has been rendered transparent for clarity. The copper<br />
layer copper is shown in brown. The metallization of the top and bottom<br />
electrode layers are marked in yellow, and the permanent magnet pairs<br />
orange. The LTCC layers of the topside/underside electrodes, as well as the<br />
LTCC bulk, cannot be seen in this picture.<br />
111
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May 2011, Aix-en-Provence, France<br />
<br />
Fig. 3: Actuation principle<br />
the electrostatics, magnetics, [9; 10], magnetostriction [11],<br />
piezoelectricity [12-14] and thermal actuation. The mesoscaled<br />
dimensions have the required surface-to-air gap ratio<br />
to enable moderate electrostatic pull forces and allow the<br />
use of permanent magnets with sufficient field densities to<br />
compensate for the poor scaling of magnetic forces in<br />
MEMS (scaling L 2 [10], compared to L 4 for pure electromagnetic<br />
actuation, where L is the critical dimension). Both<br />
forces are contact-less and allow in reverse capacitive/inductive<br />
feedback measurements using frequencies<br />
much higher than the resonant frequency of the seismic<br />
mass. An approach of combining electrostatic and magnetic<br />
actuation is reported in [15].<br />
Neglecting fringing field effects, the pull force, F z , of one<br />
extended coil (active area 4.1mm x 5mm, gap 25!m) can be<br />
derived from the expression of the potential energy U stored<br />
in a capacitor with the surface A, the gap d, the total charge<br />
Q, the permittivity " 0 " r and the applied voltage V using<br />
equations (1-2).<br />
U =<br />
∫ Q<br />
0<br />
∫ Q<br />
q<br />
V dq =<br />
0 C dq = 1 Q 2<br />
2 C = 1 2 CV 2 = 1 ɛ 0 ɛ r A<br />
V 2 (1)<br />
2 d<br />
The pull force is obtained from the partial derivative in z.<br />
F z = − ∂U<br />
∂z = ɛ 0ɛ r A V 2<br />
2 d 2<br />
A voltage of 200V would generate a pull force of 5.8mN<br />
(deflection=0) for one coil. The maximum applicable voltage<br />
can be obtained from absolute minimum of the Paschen<br />
curve, which describes the breakdown voltage between two<br />
conductors as function of the product of pressure and electrode<br />
distance [16; 17]. According to [18], the curve minimum<br />
for air yields a breakdown voltage of 315V. Hence, a<br />
maximum voltage of 200V can be considered to be safe for<br />
all deflections.<br />
Considering [19] and neglecting the stator field distortions<br />
from the comparatively small rotor fields, the lateral<br />
forces can be calculated from the flux density of the permanent<br />
magnets in the air gap and the current through the six-<br />
(2)<br />
Fig. 4a: y-z cross-section view of the magnetic flux density (B) generated<br />
by the four permanent magnets (5mm x 1.5mm x 1mm, 1.3 T) using the<br />
simulation software package COMSOL TM . The graphic shows a highly<br />
homogeneous field between the magnet pairs in the areas A and B, where<br />
the windings of the coils are placed.<br />
Fig. 4b: Z-component of the magnetic flux density of four permanent magnets<br />
along the x-y plane, determined by a 3D magnetostatic simulation<br />
using the software package COMSOL TM .<br />
112
teen wire segments of a coil, as in (3).<br />
∮<br />
⃗F = I<br />
d ⃗ l × ⃗ B = I( ⃗ L × ⃗ B) (3)<br />
The stator field was obtained from a 3D magnetostatic<br />
simulation from the software package COMSOL TM , as<br />
shown in figure 4. The simulation results were rastered and<br />
imported in Matlab. The mean value of the flux density in<br />
the range from minimum to maximum deflection yielded<br />
value of 0.9T in the z-direction. A driving current of 100mA<br />
would generate therefore a lateral force of 7.2mN.<br />
As depicted in the simplified actuation principle in figure<br />
3, every coil actuation element requires one current source<br />
and two voltage sources. For analogue sources, a floating<br />
potential in the coil has to be taken in to account, which depends<br />
on the driving current. From the sheet resistance of<br />
the TC0302 metallization (< 2m#/sq) and the conductor<br />
geometry of the coils and connections, the total track resistance<br />
is calculated to be 0.71#. Assuming a maximum<br />
current of 100mA, a maximum voltage shift of 71mV is obtained.<br />
As the electrode voltages are above 25V, this shift is<br />
assumed to be negligible. The power consumption is estimated<br />
at 7.1mW.<br />
C. Restoring forces<br />
The restoring forces are generated using single-beam<br />
flexures in the x-y plane because this method enables a feasible<br />
way to interconnect the coils and guarantees a defined<br />
zero-deflection position. Furthermore, the spring forces,<br />
partially compensate for the quadratic dependence of the<br />
pull forces of the electrostatic actuator with respect to the<br />
deflection.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
The mechanical properties of the co-fired LTCC were obtained<br />
from the manufacturers’ data sheets. For the HL2000<br />
ceramic, the flexural strength is specified above 200MPa.<br />
The Young’s modulus, derived from DuPont 941, is assumed<br />
to be 120GPa. Tests with laser-structured doublebeams<br />
(length 12mm, width 120!m / 80!m, height 120!m)<br />
have shown that the flexures break at a deflection between<br />
1.5mm and 2mm, which leads according to the strain/stress<br />
curve of a beam with rectangular cross section to a force between<br />
11mN and 15mN and stresses between 227MPa and<br />
312MPa. Hence, the data determined by the manufacturer<br />
using four-point measurements apply to the aspect ratio of<br />
flexures, and a static maximum actuator stroke of 10!m can<br />
be calculated to result in a stress of 15.5MPa for a vertical<br />
point load. Figure 5 illustrates a flexure of one of the test<br />
structures (80!m sample), which is the smallest LTCC<br />
structure that could be manufactured with the laser.<br />
As described in [5], the strain-stress curve of LTCC<br />
shows nonlinear and strain-speed dependent characteristics.<br />
Thus, in contrast to metals, only small deflections allow the<br />
definition of a spring constant. Hence, the design was configured<br />
to provide the longest possible beams while retaining<br />
the 120º point symmetry of the device. This caused implicitly<br />
a modification of the angle between the two beams<br />
from 90º to 80º, as the via interconnections are fixed. The<br />
spring constant for a single ended beam with the dimensions<br />
l " b " h can be approximated using unified beam theory<br />
[20]:<br />
, k y = E hb3<br />
4 l 3 (4)<br />
Where k z is vertical spring constant, k y the lateral spring<br />
constant and E the elastic modulus. For the inner beams<br />
Fig. 5: Laser cut LTCC beam flexure used for stress test.<br />
Fig. 6: Prototype magnet layers made of acrylic.<br />
113
(2mm " 120!m " 120!m), k z and k y are calculated as 0.78<br />
N/mm, for the outer beams (3.7mm " 120!m " 120!m)<br />
0.12 N/mm. The springs are parallel in z direction providing<br />
a combined spring constant k z for each buckled beam flexure<br />
of 0.10N/mm. For lateral deflections, the direction of<br />
the applied force is relevant. The evaluation of the force<br />
generation model in that direction exceeds the scope of this<br />
article.<br />
D. Manufacturing considerations and feasibility tests<br />
To manufacture the prototype of the integrated actuator,<br />
the Heriot Watt University supports facilities, which include<br />
an automated screen printer with camera alignment (model<br />
DEK Horizon 265), an isostatic press (model KEKO ILS-4),<br />
a cofiring oven (model Nabertherm 30º-3000º), a CO 2 laser<br />
cutter (EPILOG Mini 18"12, 10.6!m wavelength) and a<br />
self-built powder blaster. The design has to comply with<br />
these capabilities and support rapid prototyping methods.<br />
To focus the manufacturing process on the functional layers,<br />
the layers containing the magnets are reusable acrylic<br />
frames, as depicted in figure 6. Via filling and sacrificial<br />
material filling is processed as manual stencil coating step<br />
using laser cut Mylar TM , which is delivered with the unblanked<br />
green sheet as support. The emulsion screen is the<br />
only externally manufactured mask.<br />
The device concept contains two critical aspects that have<br />
to be carefully taken into account, firstly, the structural integrity<br />
and position of the comparatively large actuated area,<br />
including the beam flexures, during the whole manufacturing<br />
process. To prevent deformations and displacements,<br />
the gaps have to be filled with sacrificial material. The<br />
green sheet is in contact with the stencil during automated<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
printing process (and manual stencil filling step for the prototype),<br />
so that the raising stencil causes a contact vacuum<br />
after the print is finished. Thus, mechanical structuring all<br />
features at once would damage the structure. To prevent<br />
this, structuring and printing has to be split in two process<br />
steps, so that the mobile parts remain fixed.<br />
A second manufacturing issue is the accurate top and bottom<br />
metallization of the flexures. The standard LTCC process<br />
printing accuracy depends on the thixotropic viscosity<br />
of the metallization, the metal particle size, the mesh density<br />
and thread diameter of the screen, as well as the alignment<br />
accuracy and configuration of the screen printer. The<br />
used screen printer is specified to guarantee 40!m alignment<br />
accuracy for appropriate fiducial qualities. The internal<br />
c p /c pk statistics show that the machine repeatability is<br />
25!m. The metallization paste is specified to support line<br />
widths of 100!m. Hence, the dimensions of the coils and<br />
counter electrodes are adapted to the size of the magnet<br />
pairs in order to prevent fringing effects in the actuator elements.<br />
This tolerance is configured to be twice the maximum<br />
assumed alignment error (40!m).<br />
Alignment and minimum printing feature size is especially<br />
crucial for the metallization of the flexures if the screen<br />
printing process step follows after the mechanical structuring<br />
and gap filling. The beam widths of 120!m have to be<br />
completely covered to prevent an increased electrical resistance<br />
whilst preventing overprinting to exclude the risk<br />
of short circuits between top and bottom side. To solve this<br />
problem, the mechanical structuring is performed after the<br />
metallization printing process, where the track widths are<br />
increased by twice the maximum alignment error. Thus, the<br />
structuring process removes the LTCC material including<br />
the metallization, and the process becomes self-aligning.<br />
Fig. 7a: Test beam structures manufactured using powder blasting. The masks<br />
were manufactured of 500!m PMMA sheets and clamped on the substrate<br />
during the process. The lamination stacking pins are used for the alignment of<br />
the mask.<br />
Fig. 7b: Powder blasting alignment test. An expired paste was printed for<br />
lower material consumption. In spite of explicitly laser cutting inaccurate<br />
alignment marks causing 70!m misalignment, the metallization is still in<br />
range to process completely covered flexures.<br />
114
As the metallization is too reflective for the wavelength<br />
of the used laser, and increased laser power could cause<br />
melting and wielding of the top and bottom layers, cold mechanical<br />
etching using a powder blaster was applied. Figure<br />
7a shows a test sheet after performing both powder blasting<br />
steps using 9!m alumina powder, 50psi nitrogen pressure<br />
and a nozzle distance of 20mm. Worst-case alignment tests,<br />
as depicted in figure 7b, show that the self-aligning still<br />
works for misalignments higher than the maximum assumed.<br />
The mechanical structures and the use of sacrificial material<br />
have an impact to the lamination and cofiring settings.<br />
Applying high lamination pressure causes a higher risk of<br />
flexure and plane deformation. Information about delamination<br />
was obtained by laminating multiple samples with nano<br />
carbon sheet material (170!m) with reduced pressure and<br />
without an additional sacrificial LTCC frame. The specified<br />
for HL2000 is 10.3MPa (1500PSI) at 75ºC in an isostatic<br />
hot water press. The test has indicated 7MPa and 75ºC can<br />
be applied without effecting observable delamination.<br />
Adapting the cofiring temperature is necessary due to the<br />
evaporation and escaping of the nano carbon material from<br />
the laminated LTCC body [21]. As a large area of sacrificial<br />
material is enclosed with LTCC layers, a high slope of the<br />
temperature can affect layer deformation and cracking. The<br />
cofiring curve provided by the manufacturer recommends a<br />
linear start slope to 440ºC in 6 hours (1.2K/min). The profile<br />
has to be adapted the slope of 0.5K/min, as reported in [22].<br />
The full manufacturing process steps and parameters required<br />
to build the functional layers of the prototype device<br />
will be presented at the conference.<br />
CONCLUSIONS<br />
The described design of an actuator for dynamic alignment<br />
purposes in six degrees of freedom is process compatible<br />
with the standard LTCC process technologies. Via drilling,<br />
via filling and track metallization printing can be performed<br />
with standard processes. Additional process steps<br />
have been used to the underside metallization of the top<br />
electrode layer, structuring of the mechanical features in the<br />
actuated layer, as well as for the gap filling with sacrificial<br />
material. The cofiring temperature curve has to be adapted.<br />
Further investigations are required to analyze the impact of<br />
layer deformations due to the double sided printing on the<br />
electrode gaps in the completed actuator package, as well as<br />
the heat transfer and thermal deformations in the beams<br />
caused by the driving currents.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
REFERENCES<br />
[1] M. Luetzelschwab, D. Weiland, and M. P. Y. Desmulliez, “Adaptive<br />
Packaging Solution for a Microlens Array Placed Over a Micro-UV-LED<br />
Array”, Micro-Assembly Technologies and Applications,<br />
pp. 129-138, 2010.<br />
[2] H. Birol, T. Maeder, C. Jacq, S. Straessler, and P. Ryser, “Fabrication<br />
of low-temperature co-fired ceramics micro-fluidic devices using<br />
sacrificial carbon layers”, Int J Appl Ceram Tec, 2, pp. 364-<br />
373, 2005.<br />
[3] D. Belavic et al., “PZT thick films for pressure sensors: Characterisation<br />
of materials and devices”, Electronics System-Integration<br />
Technology Conference, 2008. ESTC 2008. 2nd, pp. 989 - 994,<br />
2008.<br />
[4] H. Klumbies, U. Partsch, A. Goldberg, S. Gebhardt, U. Keitel, and<br />
H. Neubert, “Actuators to be integrated in Low Temperature Cofired<br />
Ceramics (LTCC) microfluidic systems”, Electronics Technology,<br />
2009. ISSE 2009. 32nd International Spring Seminar on,<br />
pp. 1 - 4, 2009.<br />
[5] Y. Imanaka, “Multilayered low temperature cofired ceramics<br />
(LTCC) technology ”, Springer, ISBN: 0387231307, pp. 229, 2004.<br />
[6] Y. Lacrotte, F. Amalou, W. Yu, and M. P. Desmulliez, “Micro-<br />
Patterning of Green Tape Ceramic Using Powder-Blasting for<br />
LTCC Manufacturing”, IMAPS CICNT, Denver, pp. 1-8, 2009.<br />
[7] D. Andrijasevic, W. Smetana, J. Zehetner, and S. Zoppel, “Aspects<br />
of micro structuring low temperature co-fired ceramic (LTCC) for<br />
realisation complex 3D objects by embossing”, Microelectronic<br />
Engineering, 2007.<br />
[8] H.-J. Kim, Y.-J. Kim, and J.-R. Kim, “An Integrated LTCC Inductor<br />
Embedding NiZn Ferrite”, Magnetics, IEEE Transactions on,<br />
42, pp. 2840 - 2842, 2006.<br />
[9] L. Lagorce, O. Brand, and M. Allen, “Magnetic microactuators<br />
based on polymer magnets”, IEEE Journal of Microelectromechanical<br />
Systems, 8, pp. 2-9, 1999.<br />
[10] B. Wagner, and W. Benecke, “Microfabricated actuator with moving<br />
permanent magnet”, Micro Electro Mechanical Systems, 1991,<br />
MEMS '91, Proceedings. An Investigation of Micro Structures,<br />
Sensors, Actuators, Machines and Robots. IEEE, pp. 27 - 32, 1991.<br />
[11] R. Greenough, M. Schulze, A. Jenner, and A. Wilkinson, “Actuation<br />
with Terfenol-D”, Magnetics, IEEE Transactions on, 27, pp.<br />
5346-5348, 1991.<br />
[12] C. Bolzmacher, K. Bauer, U. Schmid, M. Hafez, and H. Seidel,<br />
“Displacement amplification of piezoelectric microactuators with a<br />
micromachined leverage unit”, Sensors and Actuators A: Physical,<br />
157, pp. 61-67, 2010.<br />
[13] E. Heinonen, J. Juuti, and H. Jantunen, “Characteristics of piezoelectric<br />
cantilevers embedded in LTCC”, Journal of the European<br />
Ceramic Society, 27, pp. 4135-4138, 2007.<br />
[14] L. Golonka et al., “Properties of PZT thick films made on LTCC”,<br />
Microelectronics International, 22, pp. 13-16, 2005.<br />
[15] R. Holzer, I. Shimoyama, and H. Miura, “Hybrid electrostatic/magnetic<br />
microactuators”, Robotics and Automation, 1995. Proceedings.,<br />
1995 IEEE International Conference on, 3, pp. 2941-<br />
2946 vol. 3, 1995.<br />
[16] L. Ledernez, F. Olcaytug, H. Yasuda, and G. Urban, “A modification<br />
of Paschen law for Argon”, 29th ICPIG, July 12-17, 2009,<br />
Cancun, Mexico, Jun 12.<br />
[17] A. J. Wallash, and L. Levit, “Electrical breakdown and ESD phenomena<br />
for devices with nanometer-to-micron …”, Proceedings of<br />
SPIE, 2003.<br />
[18] M. J. Madou, “Fundamentals of microfabrication: the science of<br />
miniaturization ”, CRC Press, ISBN: 0-8493-0826-7, pp. 723, 2002.<br />
[19] D. A. Fleisch, “A student's guide to Maxwell's equations ”, Cambridge<br />
University Press, ISBN-13: 978-0-511-39308-2, pp. 134,<br />
2008.<br />
[20] I. Szabó, “Höhere technische Mechanik”, Springer, ISBN: 3-540-<br />
67653-8, pp. 546, 2000.<br />
[21] L. E. Khoong, Y. Tan, and Y. C. Lam, “Carbon burnout and densification<br />
of self-constrained LTCC for fabrication of …”, Journal of<br />
the European Ceramic Society, 2009.<br />
[22] A. C. Zatarain, “High-Quality CO(2) Laser Machining of LTCC<br />
Structures for Thermal Management of a Group of Single-Emitter<br />
Laser Diodes”, IMAPS Proceedings 2009.<br />
115
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Implementing MEMS resonators in 90 nm CMOS<br />
J. E. Ramstad, J. A. Michaelsen, O. Soeraasen, D. Wisland<br />
Department of Informatics, University of Oslo<br />
P.O. Box 1080 Blindern<br />
N-0316 Oslo, Norway<br />
In the ubiquitous information age of today there is an<br />
increased interest in combining actuating and sensing<br />
technology with the advanced signal processing of the well<br />
established CMOS technology. With that in mind, this work<br />
investigates the possibility of making MEMS resonators in finepitch<br />
CMOS in contrast to making MEMS in more coarsegrain<br />
CMOS processes. This is done by using the metal layers<br />
in the CMOS process both as structural layers and as a mask,<br />
using only a few post-CMOS etch steps. A modern and mature<br />
90 nm CMOS process from ST Microelectronics is used and test<br />
structures are analyzed. A set of design rules have been derived<br />
and a general design guideline is described. As a practical<br />
example, implemented MEMS frequency tunable resonators<br />
are shown that are intended to be used as VCOs in an FDSM<br />
system to show the feasibility of the work.<br />
I. INTRODUCTION<br />
Wireless Sensor Network (WSN) nodes are wireless<br />
devices that monitor the environment and send information<br />
to other nodes through a transceiver. The sensor part and the<br />
transceiver part of such a node typically utilize large bulky<br />
off-chip components to perform these tasks. MEMS devices<br />
can replace some of these off-chip components and can be<br />
implemented directly on-chip by making MEMS out of the<br />
layers offered by the CMOS process (post-CMOS). This will<br />
allow a much more compact and more functional WSN<br />
node.<br />
There is a wide interest of integrating MEMS through<br />
post-CMOS processing for both sensing and transmitting<br />
applications. Examples of such sensor devices and<br />
transceiver components include: resonators [1], inductors<br />
[2], varactors [3], self-assembly devices [4], magnetometers<br />
[5], accelerometers [6], gyroscopes [7] and switches [8]. The<br />
common denominator of these examples are that these<br />
devices are implemented in more coarse-grain CMOS<br />
processes around the 0.35!m technology node. For WSN<br />
nodes it is desirable to make full on-chip systems with<br />
sensors, a transceiver and a signal processing unit with low<br />
power consumption and compact size. In this respect,<br />
investigating the possibility of making MEMS in fine pitch<br />
CMOS (90 nm CMOS or finer) is highly interesting.<br />
II. MEMS IN A 90 NM CMOS PROCESS<br />
A. CMOS-MEMS process features<br />
Our previous implementations of MEMS in CMOS [9,10]<br />
have been realized using a 0.25 !m CMOS process from ST<br />
Microelectronics (STM) where the dies have been postprocessed<br />
by Carnegie Mellon University (CMU) through<br />
the ASIMPS service from Circuits Multi-Projets (CMP).<br />
With the assistance from CMU, a chip in the STM 90 nm<br />
CMOS process has been post-processed which follows the<br />
Silicon substrate<br />
Metal layers 1 to 4<br />
Metal layer 5<br />
Metal layer 6 or 7;<br />
shielding layer<br />
Vias<br />
CMOS circuitry<br />
MEMS resonator structure<br />
(stack of metal-dielectric from M1 to M5)<br />
S<br />
Silicon substrate<br />
(a)<br />
(c)<br />
S<br />
Dielectric layers<br />
CMOS shielded by<br />
the top metal layer<br />
Remaining dielectric layers<br />
after the first etch step<br />
S<br />
Released MEMS resonator<br />
Resulting silicons profile<br />
after the third etch step<br />
Fig. 1. 90 nm CMOS-MEMS process etch steps<br />
same etch steps in order to make MEMS in CMOS as shown<br />
in fig. 1. The top metal layer is used as a mask to define the<br />
MEMS structures. The MEMS structure will contain a stack<br />
of metals and dielectrics that will define the structural<br />
thickness which in turn allows for electrostatic operation<br />
laterally above the wafer surface. As can be seen in fig. 1b)<br />
the dielectric layers are etched away, the silicon is<br />
anisotropically etched in c) and then finally released through<br />
an isotropic etch in d).<br />
In this configuration of the STM 90 nm CMOS process<br />
we used seven metal layers with an extra metal layer<br />
typically used for bonding only. Fig. 2 shows a cross-section<br />
of seven metal layers. As this is a Dual Damascene CMOS<br />
process, all metal layers consist of copper composites except<br />
the bond pad layer which consists of aluminum. The aspectratio<br />
attainable for etching narrow gaps are limited through<br />
the DRIE etch, so the focus is not to tune the process for<br />
small gaps, but instead use self-assembly (SA) structures to<br />
create narrow gaps as shown in section III.<br />
Metal 7<br />
Metal 6<br />
M1-M5<br />
{<br />
Fig. 2. Overview over the metal layers of the process<br />
E2<br />
S2<br />
(b)<br />
(d)<br />
116
B. Implemented die overview<br />
A die containing structures for optical tests, test resonators<br />
and tunable resonators as a VCO-FDSM design (see section<br />
III for the system design details) was implemented using a<br />
standard 90 nm CMOS process as can be seen in fig. 3.<br />
Fig. 3. SEM photo of the implemented 90 nm CMOS-MEMS chip<br />
Through optical measurements, guidelines and tentative<br />
design rules can be established. The primary goal of<br />
characterizing the process is to see what kind of dimensions<br />
and parameters for MEMS structures that are attainable<br />
through a standard post-CMOS process and to see if the<br />
CMOS circuitry becomes affected by the etch steps. The die<br />
layout is seen in fig. 4 where bond pads for ESD protection,<br />
CMOS circuitry and MEMS are highlighted. These pads are<br />
electrically separated but mechanically connected together<br />
through an extra edge. This extra edge prevents the post-<br />
CMOS etch from etching beneath the ESD and CMOS pads<br />
and prevents the circuitry from being etched. The extra edge<br />
also adds mechanical robustness to the die.<br />
Extra edge<br />
to prevent<br />
etch and<br />
for mechanical<br />
support<br />
system design<br />
test resonators<br />
test structures<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
C. Characterizing the 90 nm process<br />
system design<br />
test resonators<br />
test structures<br />
ESD pad<br />
MEMS pad<br />
CMOS pad<br />
Fig. 4. Layout of the implemented 90 nm CMOS-MEMS chip<br />
When characterizing the process, it is important to focus<br />
on what parameters that are required in order to achieve<br />
good MEMS resonator performance. Many MEMS sensors<br />
and RF resonators rely on the electrostatic actuation<br />
principle, operating laterally above a surface. Equation 1<br />
below describes the electromechanical coupling coefficient:<br />
The electromechanical coupling coefficient is related to<br />
the electrostatic force that is used to achieve a certain<br />
resonator displacement which in turn results in an output<br />
current. " can be increased by a long resonator electrode<br />
length (WE), a thick resonator thickness (HR) or a small gap<br />
(g). The polarization voltage VP can be utilized to further<br />
enhance the resonator performance but is not related to the<br />
process development. The metal-dielectric stack in fig. 2<br />
shows metal 1 (M1) up to metal 7 (M7) resulting in a 4 !m<br />
thick structure. In this work, M1 up to M5 is used to make 3<br />
!m thick resonators. M1 to M5 have roughly the same<br />
thickness while M6 and M7 are more than twice as thick.<br />
Compared to coarse grain CMOS processes, making MEMS<br />
in fine-pitch CMOS requires more metal layers in order to<br />
achieve the same structural thickness.<br />
For tunable resonators it is desirable to have long thin<br />
beams with a low spring stiffness (k). Eq. 2 describes the<br />
relationship between resonator stiffness and mass where # is<br />
a topographical scaling factor. A small width (WR) and a<br />
large length (LR) is desirable for obtaining a low spring<br />
stiffness in order to tune the resonance frequency (f0) of the<br />
resonator. A large electrode length WE results in a high<br />
electrical spring stiffness which is subtracted from the<br />
mechanical spring stiffness, thus the resonance frequency<br />
can be tuned by using the polarization voltage VP. Wider<br />
MEMS (large WR) structures have been implemented that<br />
can be used as high frequency filtering components.<br />
f 0 = 1<br />
2π<br />
η = V P<br />
ε 0 W E H R<br />
g 2<br />
<br />
k<br />
m =1.03κ <br />
E<br />
ρ<br />
W R<br />
L 2 R<br />
The 90 nm CMOS process consists of copper composites<br />
instead of aluminum which is standard in older CMOS<br />
processes. This causes the effective Young’s Modulus (E) of<br />
a stack to be lower while the material density ($) will be<br />
larger due to the low-k dielectrics, resulting in a lower E/$<br />
ratio compared to older CMOS processes. A large E/p ratio<br />
is important for high-frequency resonators. The low-k<br />
dielectric at the lower layers will also cause the etch rate to<br />
diminish during the dielectric etch.<br />
Fig. 5 shows a layout of beams that consists of M1-M5,<br />
including polysilicon beneath the beams and a layer known<br />
as the ACTIVE layer. When making CMOS transistors, the<br />
active layer is used to make the thin oxide between the gate<br />
(polysilicon) of the transistor and the doped silicon beneath.<br />
This thin oxide is deposited at a lower temperature compared<br />
to the other dielectrics. By utilizing this thin oxide, the<br />
internal stress between the dielectric and metal layers<br />
becomes reduced, which in turn results in a reduced amount<br />
of curling. Beams with excessive curling is shown in fig 6.<br />
(1)<br />
(2)<br />
117
ACTIVE<br />
layer<br />
Poly<br />
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May 2011, Aix-en-Provence, France<br />
<br />
A<br />
A´<br />
W1<br />
Area<br />
without<br />
ACTIVE<br />
W2<br />
A<br />
gaps. The right part of fig. 7 shows wide aluminum beams.<br />
By using the aluminum layer thicker MEMS structures can<br />
S2 S1<br />
be implemented. However, the surface is more rough as it<br />
reacts more with the etch recipe. Also the aluminum layer is<br />
limited by the CMOS Design Rule Check (DRC) on how<br />
h(14 + n + 1 n ) (3) small the gaps can (4) be and how small the width of such a<br />
Figure 7: NODAS results mechfilter<br />
1<br />
ρ = 24(α 2 − α 1 )(T − T 0 )<br />
layer can be.<br />
Eq. 3 describes the curvature of a bimorph beam where h A Veeco white light optical profilometer has been used to<br />
is the thickness, T is the temperature, T0 is the characteristic characterize the optical test structures. Some of the most<br />
temperature, % is<br />
TCE<br />
the temperature diff =(α 2 −<br />
coefficient<br />
α 1 )<br />
of expansion important measurement (5) results are shown in fig. 8 and fig 9.<br />
(TCE) and n is the difference in Young’s Modulus between Excessive curling occurs when the metal-dielectric stack<br />
the two materials. A small difference in both % and E will becomes less homogenous or has less or no thin oxide<br />
reduce the amount of curling. Due to a rule from the CMOS beneath. Thin oxide can be made by using the active layer<br />
foundry, beams with polysilicon δ = ( L 2 )2<br />
are only partially covered from the CMOS process (6)<br />
2ρ<br />
which will reduce curling as seen in<br />
by the active layer when there is no polysilicon beneath.<br />
R qi = R m<br />
2η<br />
E N =<br />
<br />
Q<br />
q i Q filter<br />
<br />
kT<br />
C<br />
E t = 4kTR∆f<br />
3.0<br />
(9)<br />
Fig. 6. Curling seen from white light interferometer measurements<br />
1.5<br />
When performing the dielectric etch, some test structures<br />
de-attached (delaminated) themselves f2 from the layer beneath<br />
0<br />
and curled upwards IN(in,tot) 2 before the I N(in) etch dω reached the silicon<br />
0 (10) 25 50 75 100<br />
substrate. These delaminated f1 beams greatly dictate the<br />
Length from anchor point [µm]<br />
tentative design rules. Structures that curl upwards or start to<br />
delaminate do that due to stress differences between layers.<br />
Fig. 8. Beams with excessive curling<br />
i o<br />
If the stress is SNR large = enough, 20log( tensile ) forces may cause<br />
(11)<br />
I<br />
structures to curl or delaminate. Between N(in,tot) 1.00<br />
a metal layer and a<br />
M1-M5 active<br />
dielectric layer there is a barrier layer which consists of Ta,<br />
M1-M5 no active<br />
0.75<br />
TaN or TiN. The thickness of this barrier layer is only a few<br />
tenth of nanometers. If this barrier layer is attacked during<br />
etching, the metal may release itself from the layer beneath<br />
due to excessive forces.<br />
A cause of measure to prevent this can be to halve the etch<br />
0.50<br />
0.25<br />
rate to increase the O2 flow and reduce the chamber pressure<br />
0<br />
which might reduce the fluorine reactions with the barrier<br />
0 15 30 45 60<br />
layers [11]. The left part of fig. 7 shows delamination of<br />
1 !m wide and 100 !m long cantilever beams with varied<br />
Length from anchor point [µm]<br />
4<br />
A´<br />
Fig. 5. Layout overview and cross-section for design rules<br />
118<br />
fig. 9. Areas containing the active layer are produced at<br />
lower temperatures, thus reducing the built-in stress between<br />
the metal and the dielectric layers. The M1-M5 stack without<br />
the active layer has (7) about 700 nm curling while the stack<br />
with the active layer has about 300 nm curling.<br />
<br />
Curling [µm]<br />
Curling [µm]<br />
Fig. 7. Left: Beams with varied gaps. Right: Aluminum beams<br />
6.0<br />
4.5<br />
M1-M7<br />
M1-M5 (8) & Poly<br />
Fig. 9. Comparison of beams with or without the active layer
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
D. Guidelines and design rules for 90 nm CMOS-MEMS<br />
III. SYSTEM TEST DESIGNS<br />
Table I shows typical process parameters for the dielectric<br />
etch using a Plasma-Therm 790 parallel-plate RIE system.<br />
As the dielectric etch step is the most challenging step, the<br />
other two etch steps remain the same. For this run the etch<br />
time was slightly longer compared to a coarse-grain post-<br />
CMOS run, and at the end of the process O2 rinsing of the<br />
dies were performed.<br />
TABLE I<br />
Post-process dielectric etch step<br />
Typical<br />
20 CHF3<br />
Gas flow [sccm] 20 CF4<br />
95 O2<br />
Pressure [mT] 100<br />
Power [W] 65<br />
DC bias [V] 270<br />
Time [min] ~120<br />
General observations:<br />
• Beams curling sideways, upwards and even downwards<br />
were observed<br />
• Structures that are long and have sufficiently small widths<br />
will delaminate and start to curl before the release-etch<br />
• Creating uniform spacing in the MEMS area will reduce<br />
the sideways curling phenomenon greatly<br />
• Homogenous material stacks curled less<br />
• Using the active layer will reduce curling<br />
• The thickness of the top metal layer is slightly milled<br />
• Polymer deposition on the sidewalls was negligible<br />
• The top metal layer had clearly defined edges<br />
TABLE II<br />
Tentative 90 nm post-CMOS design rules<br />
Dim. [!m] Rule name Comment<br />
Minimum width 1 W1 Delamination<br />
Maximum width ~10 W2 CMOS rule<br />
Max length fixed-free < 60 L1 Delamination<br />
Max length fixed-fixed < 100 L2 Curling<br />
Max stack thickness ~5 H1 Preliminary<br />
Gap spacing 1.2 S1 Guarantees release<br />
Poly from metal edge 0.6 S2 Prone to etch<br />
Active cover edge 0.3 A1 Reduce curling<br />
Active sep poly ~0.1 A2 CMOS rule<br />
Table II shows general design rules and guidelines for a<br />
general purpose 90 nm CMOS process.<br />
General purpose 90 nm post-CMOS summary:<br />
Pros<br />
• Possible smaller electrostatic gaps<br />
• Less parasitics<br />
• Lower supply voltage for less power consumption<br />
• More intricate routing capabilities<br />
• More in thread with newer CMOS processes<br />
Cons<br />
• Smaller stack thickness possible<br />
• More stringent CMOS design rules, especially the<br />
density rules in the MEMS areas<br />
• Curling and delamination more prominent<br />
In this work, a set of soft-tunable MEMS resonators to be<br />
used as voltage-controlled oscillators (VCO) have been<br />
implemented. Fig. 10 shows an electromechanical equivalent<br />
schematic of the MEMS resonator with the previously<br />
described electromechanical coupling coefficient ". The lz, cz<br />
and rz are related to mechanical and electrical parameters of<br />
the resonator. A signal at the resonance of the resonator<br />
results in the lz and cz reactances becoming equal but with<br />
opposite sign, thus the only part left is the damping part of<br />
the resonator which is rz. When tuning the resonance<br />
frequency with VP, the lz and cz changes values. CP is a<br />
parasitic element from the routing of the MEMS resonator to<br />
the following amplifier.<br />
Fig. 10. Electromechanical schematic representation of the resonator<br />
A soft tunable parallel-plate tuning fork (PPTF) with selfassembly<br />
beams is shown in fig. 11. The part of the<br />
resonator that overlaps the fixed electrodes will have equal<br />
displacement throughout the electrode length in order to<br />
reduce non-linearities. The left part of fig. 11 shows how the<br />
electrostatic gaps are reduced by using self-assembly (SA)<br />
electrodes. The beams are designed to have more metal<br />
layers on one side for each half of the SA length. A lateral<br />
force internally is generated due to built-in sideways stress<br />
during processing, and the SA will move after release and<br />
create a 200 nm small gap between the resonator and the<br />
electrode.<br />
SA (in) PPTF SA (out)<br />
1.2 m<br />
0.6 m<br />
1.2 m<br />
VAC<br />
0.6 m<br />
1 : !<br />
Out<br />
Out<br />
lz cz rz ! : 1<br />
Fig. 11. Left: The self-assembly concept. Right: PPTF-resonator with SA<br />
The lateral displacement of a self-assembly beam is shown<br />
more clearly in fig. 12 where the self-assembly beams are<br />
100 !m long and 2 !m wide. Another MEMS resonator<br />
design with self-assembly beams was implemented as a<br />
clamped-clamped (CC) beam as shown in fig 13.<br />
CP<br />
+<br />
-<br />
119
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May 2011, Aix-en-Provence, France<br />
<br />
102<br />
430<br />
101<br />
100<br />
Simulation<br />
Analytic<br />
425<br />
420<br />
Simulation<br />
Analytic<br />
Curling of this Self-Assembly<br />
electrode is made intentional<br />
Frequency [kHz]<br />
99<br />
98<br />
97<br />
96<br />
Frequency [kHz]<br />
415<br />
410<br />
405<br />
95<br />
400<br />
94<br />
395<br />
93<br />
0.3 0.35 0.4 0.45 0.5 0.55 0.6 0.65 0.7 0.75<br />
V [V] P<br />
(a)<br />
390<br />
0.2 0.4 0.6 0.8 1 1.2 1.4<br />
V [V] P<br />
(b)<br />
)"<br />
3<br />
)'<br />
3<br />
Fig. 12. Self-Assembly beams shows lateral curling<br />
)#<br />
*"<br />
*#<br />
**<br />
*"<br />
+,-./012,-34,113567)"8<br />
*#<br />
!"<br />
!#<br />
+,-./012,-34,113567*#8<br />
*%<br />
*'<br />
!#<br />
!*<br />
#"<br />
!"<br />
##<br />
A B<br />
C"D%#A<br />
A B<br />
C"D*A<br />
!%<br />
!'<br />
A B<br />
C)D$A<br />
A B<br />
C)D#A<br />
Narrow 200<br />
nm gap<br />
$" 3<br />
!" #" $" %" &" '" ("" ((" ()" (*" (!"<br />
90/:;/-?@8<br />
(c)<br />
"# 3<br />
!"# !$# !%# !&# !'# !(# "## ")# "*# "!# ""#<br />
90/:;/-?@8<br />
Fig. 14. Frequency tune range and AC plot for the PPTF and CC resonator<br />
The simulated range of the frequency tuning is shown in<br />
fig. 14 a) and b) while fig. 14 c) and d) show the transfer<br />
function characteristics for the PPTF and CC resonator<br />
respectively. A larger VP results in a reduced resonator loss,<br />
thus reducing the impedance rz that the resonator represents<br />
at the resonance frequency.<br />
(d)<br />
A<br />
Fig. 13. Clamped-clamped beam with Self-Assembly beam<br />
Table III describes the dimensions, parameters and results<br />
of these two soft-tunable MEMS resonators. The PPTF<br />
occupies more space and has less tunability compared to the<br />
CC-beam, however the electromechanical coupling<br />
coefficient is better and the input and output electrode is<br />
clearly separated from each other.<br />
TABLE III<br />
Resonator dimensions and results<br />
PPTF CC<br />
Resonator dimensions<br />
Resonator length, LR [!m]<br />
Resonator width, WR [!m]<br />
LFRAME=100<br />
LCANTILEVER=50<br />
WFRAME=2<br />
WCANTILEVER=1<br />
100<br />
Electrode length, WE [!m] 100 100<br />
Resonator-to-electrode gap, g [nm] 200 200<br />
Resonator results<br />
Nominal resonance frequency [kHz] 102.84 424.73<br />
Frequency tuning range [kHz] 7.12 27.35<br />
Tunability in percentage [%] 6.92 6.43<br />
Effective beam stiffness [N/m] 4.30 4.74<br />
Electromechanical coupling [nN/V] 55.34 35.24<br />
1<br />
x(t)<br />
B FDSM y[n]<br />
V CLK<br />
R<br />
V P<br />
Fig. 15. System overview of resonator in an oscillator loop and with a FDSM<br />
The MEMS resonator is put in a feedback loop with an<br />
amplifier to make it oscillate naturally, shown in fig. 15. The<br />
output of the oscillator is buffered to digital logic levels. A<br />
Frequency Delta Sigma Modulator (FDSM) [12] was<br />
included on-chip in order to show CMOS circuit<br />
compatibility with the MEMS etch, and to demonstrate<br />
CMOS and MEMS interoperability. The FDSM is a<br />
frequency to digital converter, used here to generate a<br />
quantized digital bitstream output from the MEMS oscillator<br />
signal. This bitstream can be post processed offline to<br />
recover the spectral contents of the oscillator output. The<br />
FDSM was designed to operate with sampling frequencies<br />
up to 20 MHz. The amplifier A in the oscillator loop consists<br />
of a common-source amplifier set up in a Pierce amplifier<br />
configuration with a loop gain of more than one in order to<br />
initiate and sustain oscillation.<br />
The clamped-clamped resonator is quite soft and has a<br />
tuning-range of slightly more than 20 kHz using VP from<br />
0.4 V to 1.4 V. For the PPTF resonator, the tuning range<br />
would be from 0.3 V to 0.75 V with a 7 kHz tuning range.<br />
120
IN<br />
CLK<br />
OUT<br />
Fig. 16. Probe setup for FDSM measurements<br />
Before etch<br />
As a part of verifying that the CMOS circuitry survived<br />
the etch, the die was probed as seen in fig. 16. An electrical<br />
test was performed before and after etch as shown in fig. 17,<br />
confirming that the CMOS circuitry is still functional after<br />
the MEMS post-processing. A 1 V supply was used to power<br />
the circuits and the ESD protection circuits in the pad frame<br />
was active. The input of the FDSM was stimulated with 1 V<br />
at 10 kHz while the CLK signal was applied with 1 V at<br />
40 kHz. The output of the FDSM was measured using an<br />
Agilent 54524A oscilloscope.<br />
IV. CONCLUSION<br />
After etch<br />
Fig. 17. Testing CMOS circuitry before and after etch processing<br />
A 90 nm CMOS die has been implemented with a set of<br />
MEMS test structures to characterize the feasibility of<br />
making MEMS resonators in fine-pitch CMOS. The die<br />
contained optical test structures, MEMS test structures and<br />
VCO-FDSM designs. A set of design rules and guidelines<br />
for a general purpose 90 nm CMOS process has been<br />
derived. The important parameters for designing MEMS<br />
resonators were pointed out and examples of two different<br />
tunable MEMS resonators were described. Verification of<br />
the post-CMOS process was performed by testing the<br />
CMOS circuity before and after the etch. Future work<br />
include thorough measurement of the MEMS resonators and<br />
the total system performance.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
REFERENCES<br />
OUT<br />
IN<br />
CLK<br />
[1] J. Teva et al, “From VHF to UHF CMOS-MEMS resonator<br />
monolithically integrated in a standard 0.35 !m CMOS<br />
technology”, in Proc. IEEE 20th Int. Conf. MEMS, pp.779-782,<br />
2007<br />
[2] S.-H. Tseng et al, “A 5.8-GHz VCO with CMOS-compatible<br />
MEMS inductors”, Sensors and Actuators, Vol. 139, pp. 187-193,<br />
2007<br />
[3] M. Bakri-Kassem, S. Fouladi and R. R. Mansour, “Novel High-Q<br />
MEMS curled-plate variable capacitors fabricated in 0.35!m<br />
CMOS technology”, Microwave Theory and Techniques, IEEE<br />
Transactions on, Vol. 56, No. 2, pp. 530-541, 2008<br />
[4] A. Jain, H. Qu, S. Todd and H. Xie, “A thermal bimorph<br />
micromirror with large bi-directional and vertical actuation”,<br />
Sensors and Actuators, Vol. 122, No. 1, pp. 9-15, 2005<br />
[5] F. Tejada, A. G. Andreou, D. K. Wickenden and A. S.<br />
Francomacaro, “Surface micromachining in Silicon on Sapphire<br />
CMOS technology”, Circuits and Systems, 2004. ISCAS ’04.<br />
Proceedings of the 2004 International Symposium on, Vol 4, pp.<br />
IV-920-3, May 2004<br />
[6] H. Qu and H. Xie, “Process Development for CMOS-MEMS<br />
Sensors with robust electrically isolated bulk silicon<br />
microstructures”, Journal of Microelectromechanical Systems, Vol.<br />
16, No. 5, pp. 1152-1161, October 2007<br />
[7] H. Xie and G. K. Fedder, “Fabrication, characterization and analysis<br />
of a DRIE CMOS-MEMS Gyroscope”, Sensors Journal IEEE, Vol.<br />
3, No. 5, pp. 622-631, October 2003<br />
[8] C.-L. Dai et al, “Modeling and fabrication of a<br />
microelectromechanical microwave switch”, Microelectronics<br />
Journal, Vol. 38, No. 4-5, April 2007<br />
[9] O. Soeraasen and J. E. Ramstad, "From MEMS Devices to Smart<br />
Integrated Systems", Journal of Microsystem Technologies, Vol. 14,<br />
No. 7, pp. 895-901, Springer-Verlag, 2008<br />
[10] J. E. Ramstad, K. G. Kjelgaard, B. E. Nordboe and O. Soeraasen,<br />
"RF MEMS front-end resonator, filters, varactors and a switch<br />
using a CMOS-MEMS process”, Design, Test, Integration &<br />
Packaging of MEMS/MOEMS, Symposium on, pp. 170-175, April<br />
2009<br />
[11] X. Zhu, S. Santhanam, H. Lakdawala, H. Luo and G. K. Fedder,<br />
“Copper interconnect low-K dielectric post-CMOS<br />
micromachining”, in Proc. of 11th International Conference on<br />
Solid-state sensors and actuators digest of technical papers, pp.<br />
1548-1551, 2001<br />
[12] M. Hovin, A. Olsen, T. S. Lande, C. Toumazou, “Delta-Sigma<br />
modulators using frequency modulated intermediate values”,<br />
Journal of Solid-State Circuits, vol. 32, no. 1, pp.13-22, 1997<br />
ACKNOWLEDGEMENTS<br />
The authors would like to thank Suresh Santhanam from<br />
Carnegie Mellon University for post-processing the dies.<br />
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<br />
The Influence of Adhesive Materials on<br />
Chip-On-Board Packing of MEMS Microphone<br />
Cheng-Hsin Chuang *1 , Yi-Hsuan Huang 1 and Shin-Li Lee 2<br />
1 Department of Mechanical Engineering, Southern Taiwan University<br />
No. 1, Nantai St., Yung-Kang City, Tainan, Taiwan, ROC.<br />
2 Micro System Technology Center, Institute Technology Research Institute Southern, Tainan, Taiwan, ROC.<br />
*E-mail Address:chchuang@mail.stut.edu.tw, Tel:+886-6-3010081 and Fax:+886-6-2425092<br />
Abstract- Adhesive material is commonly used for attaching die<br />
onto the printed wiring board (PWB) in the Chip-on-Board (COB)<br />
packaging of MEMS devices. However, the polymer-based<br />
adhesive usually possesses large difference in the coefficient of<br />
thermal expansion (CTE) between silicon chip and PWB.<br />
Therefore, the mismatch of CTE could lead to the<br />
thermally-induced stress and coupling deformation of multilayer<br />
structure in the reflow process as surface-mount technology<br />
(SMT) on printed circuit board (PCB). In this study, three<br />
different adhesive materials, namely 2025D, 3140RTV and<br />
SDA6501, and two different cap materials, namely liquid crystal<br />
polymer (LCP) and nickel (Ni), were evaluated the influences on<br />
the thermally-induced stress in the ploy-silicon diaphragm of<br />
MEMS microphone based on Finite Element Analysis (FEA).<br />
According to the results, we obtained the following two findings:<br />
(1) The CTE mismatch of LCP cap and the metal (Ni) cap caused<br />
different type of thermal deformations of PWB and lower<br />
thermally-induced stress and deformation were found in the case<br />
of LCP cap, however, different cap materials less affected the<br />
thermal stress in the diaphragm. (2) Soft adhesive materials<br />
(3140RTV and SDA6501) have better mechanical isolation of<br />
PWB thermal deformation due to the buffer layer effects. On the<br />
contrary, hard adhesive material (2025D) could be affected by<br />
PCB thermal deformation when the thickness of adhesive was less<br />
than 30μm, thus, a lower stress in the diaphragm existed due to<br />
the stress compensation by PWB thermal deformation. In<br />
general, present study provides the basis of selection of adhesive<br />
material for COB MEMS packaging.<br />
Keywords: Diaphragm, Adhesive, Die attach, Thermal analysis<br />
I. INTRODUCTION<br />
Chip-on-Board (COB) packaging technology is directly<br />
bonding a device chip to a second level substrate with<br />
adhesive material. Currently, COB packaging has been<br />
adopted by semiconductor manufacturing as well as MEMS<br />
foundry due to multiple advantages including low thermal<br />
resistance, high cost efficiency, ideal design flexibility, etc.<br />
However, the mismatch of coefficients of thermal expansion<br />
(CTE) between multi-laminated materials may introduce<br />
thermal stress and deformation when the COB device is<br />
further mounted onto a printed circuit board (PCB) surface by<br />
surface-mount technology (SMT). During SMT process, a<br />
heating process of solder reflow is necessary to produce a<br />
high quality of solder joint between COB device and PCB.<br />
The soldering processing involves four steps such as preheat,<br />
activation, reflow and cool down, the highest temperature<br />
during reflow usually gets up to 260℃ and the total time is<br />
about 360 seconds from oven entrance to the end of reflow<br />
stage. Therefore, it’s necessary to evaluate the thermal<br />
influence on COB devices as assembling by SMT process.<br />
Several researchers have investigated the thermal influences<br />
of IC packaging based on COB method [1-3]. Tom Tuhus and<br />
Are Bjomeklett [1] indicated a soft adhesive material could<br />
bring the stress relaxing effect but may lead to fatigue of the<br />
adhesive layer as repeat cyclic temperature changes. Qing’an<br />
Huang, et al.[2], proposed a 2D theoretical model of COB<br />
packaging for evaluating the coupling deformation and stress<br />
under thermal load. They found less thermal influence when<br />
the silicon die attached on a ceramic substrate instead of an<br />
organic substrate. Andrew A. O. Tay and K. Y. Goh [3]<br />
addressed the delamination phenomenon might occur during<br />
solder reflow in the COB packaging device. As we knows, the<br />
packaging of MEMS devices is quite different with regards to<br />
IC packaging due to internal moving parts and external<br />
environmental exposure for sensing purposes. The moving<br />
parts in an MEMS device usually are the most important and<br />
fragile structures relevant to its performance, e.g., sensing or<br />
actuating; therefore, the thermally-induced stresses and<br />
distortions of the moving parts could affect overall<br />
performance after reflow process. Recently, COB packaging<br />
has already been used in MEMS devices and found<br />
significant influence on sensor performance after adhesive<br />
curing or reflow process. Zhigno Sun et al. [4] experimentally<br />
revealed the residual stress after adhesive curing could be tens<br />
of MPa for a piezoresistive pressure sensor packaged by COB.<br />
Furthermore, the offsets of pressure sensor output varied with<br />
different kinds of adhesive materials and adhesive thickness<br />
have been investigated by several studies [5-7]. Consequently,<br />
the selection of adhesive material and its thickness could play<br />
an important role for reduction of the thermal influence under<br />
thermal load. In this study, we tried to numerically evaluate<br />
two different cap materials and three adhesive materials with<br />
different Young’s modulus and CTE for a silicon MEMS<br />
microphone attached on printed wiring board (PWB) as<br />
shown in the Fig. 1. Two cap materials are nickel and liquid<br />
crystal polymer (LCP), and three commercial adhesive<br />
122
11-13 <br />
May 2011, Aix-en-Provence, France<br />
materials are 2025D (Ablestik), 3140RTV and DA6501<br />
<br />
convergence tests.<br />
(Dow Corning). Based on these material properties, the better<br />
solution for reduction of thermal influence on MEMS<br />
microphone can be provided by using an actual size 3D model<br />
in the numerical analyses.<br />
Fig. 2: A FEA model of CMOS Microphone packing and diaphragm thickness<br />
is 0.002mm.<br />
Fig.1: Silicon microphone chip packaged by COB packing,(a) Metal cap<br />
packing ,(b) LCP cap packing , (c) The cross-section of chip.<br />
II.<br />
SIMULATION MODEL<br />
A 3D model of COB packaging for a backside-etched<br />
MEMS microphone chip with a 2μm-thick polysilicon<br />
diaphragm attached onto a PWB substrate by adhesive<br />
material was illustrated in Fig. 2. The dimensions of COB<br />
packaging and all the material properties were indicated in the<br />
Fig. 1 and listed in Table 1, respectively. There were three<br />
kinds of adhesive materials utilized for die attachment, first<br />
adhesive material, 2025D, possesses higher Young’s modulus<br />
but relative low CTE, the other two silicone adhesive<br />
materials, 3140RTV and DA6501, have very low Young’s<br />
modulus but relative high CTE. In addition, an external cap<br />
was bonded on PWB by epoxy material for shielding<br />
electromagnetic wave and pollutions. Two kinds of cap<br />
material, nickel and LCP, were investigated to decrease the<br />
thermal deformation of PWB for further reduction of the<br />
thermal influence on diaphragm. For heating process, we<br />
adopted the heating temperature of the solder reflow process<br />
which took a total of 400 seconds to raise room temperature<br />
from 23 ℃ to 260 ℃ and then dropped back to room<br />
temperature, as shown in Fig. 3. In order to evaluate the<br />
effects of adhesive thickness, three kinds of thickness were<br />
employed for each adhesive material, 20, 30 and 50 μm, in the<br />
simulation models. All these numerical models were solved<br />
by commercial finite element analysis software, ABAQUS. In<br />
the boundary condition setting, only one corner at the bottom<br />
of PWB substrate was set as no displacement in X, Y and Z<br />
directions, i.e., U X =U Y =U Z =0, but the other three corners<br />
were set as no displacement in the Y direction, i.e., U Y =0.<br />
Therefore, the model has free constraint in the X and Z<br />
direction as thermal expansion and there is no rotational<br />
constraint in this model so that we can evaluate the warpage<br />
of packaging cap as well as PWB substrate. The element type<br />
used in the 3D model was 8-node cubic element and the total<br />
element number was about 79000 to 84000 based on<br />
Fig. 3: The heating history for simulation of reflow process as COB device<br />
was surface mounted on PCB in the SMT process.<br />
III.<br />
(a) Influence of packaging cap<br />
RESULTS AND DISCUSSIONS<br />
The thermal deformations of the 3D model as maximum<br />
reflow temperature 260 ℃ for different packaging caps and<br />
adhesives were indicated in Fig. 4. In the Fig. 4(a) and 4(b),<br />
the thermal deformations of LCP cap were larger than PWB<br />
substrate due to a higher CTE value of LCP than PWB. Thus,<br />
a convex thermal deformation can be seen in the PWB<br />
substrate. However, a concave thermal deformation of PWB<br />
substrate can be found in the Fig. 4(c) and 4(d) due to a lower<br />
CTE value of nickel than PWB. Namely, the deform direction<br />
are opposite for different cap materials. Furthermore, the<br />
thermal warpage of PWB in the nickel cap case is larger than<br />
the value in the LCP cap due to the nickel cap constrains the<br />
PWB to deform in the out-of-plane direction. The maximum<br />
thermal stress for nickel cap case occurred at the metal cap<br />
near the bonding region and the Von-Mises stress values were<br />
about 246.2 ~ 246.9 MPa as shown in the Fig. 4(c) and 4(d),<br />
which is larger than the stress values in the LCP cap cases,<br />
192.5 ~ 197.7 MPa, happened at the bonding paste epoxy<br />
material. Therefore, LCP cap for COB packaging can provide<br />
less thermal deformation of PWB and lower thermal stress in<br />
the bonding paste to reduce the thermal influence on MEMS<br />
123
11-13 <br />
May 2011, Aix-en-Provence, France<br />
chip and improve the reliability of cap bonding.<br />
<br />
30<br />
Max stress on membrane (MPa)<br />
20<br />
10<br />
LCP Cap<br />
2025D<br />
3140RTV<br />
DA6501<br />
Fig. 4: The deformation and stress distribution of 3D model with 20um-thick<br />
adhesive at 260°C, all the deformations are magnified by 10 times. (a) and (b)<br />
are LCP cap with different adhesives 2025D and 3140RTV, respectively. (c) and<br />
(d) are Ni cap with different adhesives 2025D and 3140RTV, respectively.<br />
(b) Influence of adhesive material and thickness<br />
As in the COB packaging, only adhesive material was used<br />
to attach the chip and the PWB together, therefore, the<br />
thickness and material properties of the adhesive material<br />
play a critical role to reduce the thermal stress in the<br />
membrane during reflow process. The maximum stress on the<br />
membrane at 260 ℃ for different adhesive material and<br />
thickness in the cases of LCP cap and nickel cap were shown<br />
in the Fig. 5(a) and 5(b), respectively. As the results, no<br />
matter packaged under LCP cap or nickel cap, the difference<br />
of thermal stress in the membrane is small. Moreover, the<br />
softer adhesive material (3140RTV and DA6501) shows a<br />
rising trend of maximum thermal stress in the membrane<br />
when the thickness increases owing to the thicker adhesive<br />
leads to larger thermal expansion in the adhesive layer, which<br />
induces larger stress in the membrane. In addition, the stress<br />
values in the cases with softer adhesive were higher than the<br />
values in the cases with harder adhesive (2025D). This can be<br />
attributed to the CTE of 2025D is only about half value of<br />
3140RTV or DA6501. Another interesting phenomenon in<br />
the Fig. 5 is that when the thickness of 2025D adhesive ranges<br />
between 20μm and 30μm, its stress first declined before it<br />
increased, showing a different trend from the other two soft<br />
adhesive materials. This is due to the high elasticity of 2025D<br />
adhesive material; therefore, when the thickness of 2025D<br />
adhesive is thin the thermal deformation of PWB could affect<br />
to the silicon chip. However, when the thickness of 2025D<br />
adhesive material is thicker above 30μm the thermal<br />
deformation of PWB is less influence to the silicon chip due<br />
to the stress relaxation and energy absorption by adhesive<br />
material, so called the buffer layer effect. In contrast, the<br />
buffer layer effect starts to play at the very beginning on an<br />
adhesive material with lower Young’s modulus, so the silicon<br />
chip is not strongly affected by the PWB deformation at the<br />
bottom for a soft adhesive case.<br />
Max stress on memebrane (MPa)<br />
0<br />
30<br />
20<br />
10<br />
0<br />
20 30 40 50 60<br />
Adhesive thickness (um)<br />
(a)<br />
Metal Cap<br />
2025D<br />
3140RTV<br />
DA6501<br />
20 30 40 50 60<br />
Adhesive thickness (um)<br />
(b)<br />
Fig. 5: The thermal stresses in the diaphragm for various cases with different<br />
adhesive material and different thickness, (a) Material of cap is LCP and (b)<br />
Material of cap is nickel.<br />
(c) Discussion on buffer layer effect<br />
As indicated in the Fig. 5, a different phenomenon between<br />
soft and hard adhesive when the thickness blew 30μm. Hence,<br />
the elasticity of adhesive material has significant influence on<br />
the thermal stress in the membrane. When a hard adhesive<br />
material (2025D) with 20μm to 30μm in thickness, the<br />
inadequate thickness fails to bring obvious buffer layer effect<br />
so that the thermal stress of the PWB at the bottom goes<br />
upwards. Fig.6 (a) shows how the membrane stress changes<br />
as the heating temperature described in Fig. 3. The membrane<br />
stress increases when the 2025D adhesive material is<br />
thickened. In the stress history of the 20μm-thickness case,<br />
the stress increased with temperature at beginning in a tensile<br />
range, but the stress turned to compressive value as the<br />
temperature went to 260℃. The similar trend can be observed<br />
in the case of 30μm thickness, but the compressive stress is<br />
smaller. As the results, the thermal stress in the membrane<br />
with 2025D adhesive material declined first as the thickness<br />
is between 20 to 30μm. From the comparison of thermal<br />
deformations between PWB and membrane for 2025D<br />
124
adhesive material with 30μm thickness illustrated in the Fig.<br />
6(b), the thermal deformation in the PWB is in the shape of<br />
upward concave, so that the PWB deformation could affect<br />
the silicon chip in a compressive way if the adhesive as a<br />
stress-transmission layer. Therefore, the tensile thermal stress<br />
in the membrane could be compensated by the external<br />
compressive stress from the bottom PWB substrate. However,<br />
when the thickness is increased, the buffer layer effect<br />
improves to reduce the transmission of the external stress<br />
from the bottom, so that the compressive stress in the<br />
membrane gradually declines in the 30μm-thickness case of<br />
2025D adhesive.<br />
In Fig.7 (a), as the soft adhesive material demonstrates<br />
buffer layer effect from the beginning, as the temperature<br />
rises, the thermal stress of the chip also increases but there is<br />
little difference between the case of 20μm thickness and<br />
30μm thickness. When the thickness of adhesive is thin, the<br />
thermal stress in the membrane mainly comes from the CTE<br />
mismatch between poly-silicon and silicon substrate in the<br />
range of dozen MPa. As the thickness is over 30μm, thermal<br />
expansion of adhesive directly affects the diaphragm<br />
therefore the stress in the diaphragm increase rapidly. In<br />
addition, from the comparison of thermal deformation<br />
between PWB and membrane in the 3140RTV adhesive<br />
material as indicated in the Fig. 7(b), the deformation of PWB<br />
is in an outward convex shape, therefore, the external stress<br />
from bottom PWB to membrane can be regarded as a tensile<br />
way. So that, we cannot find the stress compensation in a soft<br />
adhesive material.<br />
Stress on membrane (MPa)<br />
25<br />
20<br />
15<br />
10<br />
5<br />
0<br />
-5<br />
LCP Cap 2025D<br />
2025D 20um<br />
2025D 30um<br />
2025D 50um<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
0<br />
Displacement (mm)<br />
-0.002<br />
-0.004<br />
-0.006<br />
LCP Cap - Adhesive 2025D<br />
2025D-30um-Mem<br />
2025D-30um-PWB<br />
-0.008<br />
0 0.4 0.8 1.2 1.6<br />
Distance (mm)<br />
(b)<br />
Fig. 6: (a) The time history of thermal stress in the membrane as COB<br />
packaging with LCP cap and difference thickness of 2025D adhesive; (b) The<br />
deformation of membrane and PWB at 260°C as COB packaging with LCP cap<br />
and 30 μm thickness of 2025D adhesive.<br />
Stress on membrane (MPa)<br />
30<br />
25<br />
20<br />
15<br />
10<br />
5<br />
0<br />
-5<br />
-10<br />
0 50 100 150 200 250<br />
0.006<br />
0.004<br />
LCP Cap 3140RTV<br />
3140RTV 20um<br />
3140RTV 30um<br />
3140RTV 50um<br />
Time (sec)<br />
(a)<br />
LCP Cap - Adhesive 3140RTV<br />
3140RTV-30um-Mem<br />
3140RTV-30um-PWB<br />
-10<br />
-15<br />
0 50 100 150 200 250<br />
Time (sec)<br />
(a)<br />
Displacement (mm)<br />
0.002<br />
0<br />
-0.002<br />
0 0.4 0.8 1.2 1.6<br />
Distance (mm)<br />
(b)<br />
Fig. 7: (a) The time history of thermal stress in the membrane as COB<br />
packaging with LCP cap and difference thickness of 3140RTV adhesive; (b)<br />
The deformation of membrane and PWB at 260°C as COB packaging with<br />
LCP cap and 30 μm thickness of 3140RTV adhesive.<br />
125
IV.<br />
CONCLUSIONS<br />
In this paper, we evaluated three kinds of adhesive material<br />
as well as two cap materials for COB packaging of MEMS<br />
microphone. According to the simulations results, the<br />
thickness of adhesive material is the critical parameter for<br />
controlling the thermal stress in the diaphragm during heating<br />
process, and the 2025D adhesive material provides lower<br />
thermal stress in the diaphragm due to the lower CTE<br />
mismatch between silicon and adhesive. However, owing to<br />
the high Young’s modulus of 2025D adhesive material, the<br />
influence of thermal deformation from bottom PWB cannot<br />
be neglected as thin thickness of 2025D. In contrast, a low<br />
Young’s modulus adhesive materials, such as 3140RTV and<br />
DA6501, effectively isolates the silicon chip with less<br />
influence of thermal deformation from bottom PWB as the<br />
buffer layer effect. However, the thickness of soft adhesive<br />
needs to be controlled in a low level to prevent the thermal<br />
stress in the diaphragm rising with thickness. Consequently, a<br />
soft adhesive material with low CTE value is better for COB<br />
packaging of MEMS device with diaphragm, such as<br />
microphones, pressure sensors and so on.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Nanotechnology. In 2004, he joined the Electronics Research<br />
Organization and Service (ERSO) at ITRI, where he conducted<br />
development of the MEMS microphone and SAW based<br />
biosensor. In 2005, he was recruited by the Department of<br />
Mechanical Engineering and Institute of Nanotechnology at the<br />
Southern Taiwan University as an Assistant Professor. Now he<br />
is an Associate Professor and leads the Micro and Nano Sensing<br />
Technology Lab (MANST Lab). His research interests focus on<br />
flexible tactile sensors, Roll-to-Roll imprinting technology, and<br />
DEP chips for single-cell-based biosensors. He has published<br />
over 80 papers in different international journals and<br />
conferences and has owned 10 patents in biosensor and tactile<br />
sensor technology.<br />
Dr. Chuang won two Special Awards of HIWIN Thesis<br />
Award in 2007 and 2008 as well as two best conference paper<br />
awards of 3 rd IEEE NEMS in 2008 and Taiwan automation<br />
conference in 2010.<br />
REFERENCES<br />
1. Tom Tuhus, Are Bjomeklett, ”thermal Cycling Reliability of Die<br />
Bonding Adhesives”, Proceedings of IEEE/IRPS, pp.204-208, 1993<br />
2. M. Li, Q. Huang, J. Song, J. Tang, F. Chen, ”Theoretical and<br />
Experimental Study on the Thermally Induced Packaging Effect in<br />
COB Structures”, Proceedings of Electronic Packaging Technology.<br />
ICEPT '06. 7 th , pp1-5, 2006.<br />
3. Andrew A. O. Tay, K. Y. Goh, ”A Study of Delamination Growth in the<br />
Die-Attach Layer of Plastic IC Packages Under Hygrothermal Loading<br />
During Solder Reflow”, Proceedings of IEEE Transactions on Device<br />
and Materials reliability, vol. 3, no. 4,pp.144-151, December 2003<br />
4. Z. G. Sun, W. D. Huang, Y. Jiang, L. Luo, ”Evolution of<br />
Residual-Inplane Stress during Adhesive Curing and Recuring in<br />
Chip-on-Board Packages”, Journal of ELECTRONIC MATERIALS,<br />
Vol. 31, No. 8,pp.887-894, 2002<br />
5. M. S. Zarnik, D. Rocak, and S. Macek, “Residual stresses in a<br />
pressure-sensor package induce adhesive material during curing: a case<br />
study”, Proceedings of Sensors and Actuators A, Vol. 116, pp. 442-449,<br />
2004.<br />
6. J.B. Xu, Y.L. Zhao, Z.D. Jiang, ”Analysis of the Packaging Stresses in<br />
Monolithic Multi- Sensor”, Proceedings of the 2nd IEEE International<br />
Conference on Nano/Micro Engineered and Molecular Systems,<br />
pp.241-244, 2007.<br />
7. Z. Y. Zhang, Z. Wan, C. Liu, G. Cao, Y. Lu and S. Liu , “Effects of<br />
Adhesive Material on the Output Characteristics of Pressure Sensor ”,<br />
Proceedings of 11th International Conference on Electronic Packaging<br />
Technology & High Density Packaging, pp.657-660, 2010.<br />
Biography:<br />
Cheng-Hsin Chuang (M’04) received his<br />
B.S. degree and Ph.D. degree from the<br />
National Cheng Kung University in 1995<br />
and 2002, respectively, both in Civil<br />
Engineering. He then held the Postdoctoral<br />
research scholarship with the Center for<br />
Micro/Nano Science and Technology at<br />
NCKU, where he held the lead position in<br />
the core facilities for MEMS fabrication and<br />
126
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Table1. Material properties for simulation.<br />
2025D DA 6501 3140RTV Epoxy Substrate Membrane PWB LCP cap Metal cap<br />
Material Silica Silicone Silicone Epoxy Silicone Poly-si FR4 LCP Ni<br />
Conductivity<br />
0.4 0.18 0.14 0.92 124 e-6 30 e-6 3 e-7 1.6 e-6 607 e-6<br />
(W/mK)<br />
Density<br />
2e-006 1e-006 1.03e-006 1.8e-006 2.329 e-6 2.33 e-6 1.9 e-6 2.7e-3 8.8 e-3<br />
(Kg/mm 3 )<br />
Elasticity<br />
Young’s<br />
modulus<br />
(MPa)<br />
Poissson<br />
Ratio<br />
Expansion<br />
(ppm/℃)<br />
Specific Heat<br />
(J/Kg-℃)<br />
410 (25℃)<br />
60 (100℃)<br />
40 (150℃)<br />
70 (200℃)<br />
120 (250℃)<br />
120 (300℃)<br />
0.88 2.918 [6] 3 e4 1.31 e5 1.50 e5 1.6 e4 68.8 [6] 2.07 e5<br />
0.3 0.2 [6] 0.24 [6] 0.3 0.28 0.22 0.28 0.31 [6] 0.31<br />
0(0℃)<br />
48(42℃)<br />
140(43℃)<br />
140(260℃)<br />
674 674<br />
(Assume<br />
same with<br />
2025D)<br />
300 315 9 (20℃)<br />
9 (135℃)<br />
35 (136℃)<br />
35 (260℃)<br />
674<br />
(Assume<br />
same with<br />
2025D)<br />
2.7 2.33 16 58.5873 0.131<br />
1000 702 710 1369 1000 460<br />
127
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May 2011, Aix-en-Provence, France<br />
<br />
Model of a voltage driven capacitive coupled micro<br />
electro-mechanical RF Switch<br />
P. Heeb a , W. Tschanun b , R. Buser a<br />
a Interstate University of Applied Sciences of Technology Buchs; Institute for Micro- and Nanotechnology, Buchs, Switzerland<br />
b Reinhardt Microtech AG, Wangs, Switzerland<br />
ABSTRACT<br />
A comprehensive and completely parameterised model is<br />
proposed to determine the related electrical and mechanical<br />
dynamic system response of a voltage driven capacitive coupled<br />
micro mechanical switch. An analytical approach is used<br />
throughout the modelling, providing representative coefficients<br />
in the set of two coupled time-dependent differential<br />
equations. The model also describes all the transferred energies,<br />
as e.g. the dissipated energy at the switch contacts and<br />
the re-feeded electrical power to the bias line. The determined<br />
switching dynamics is confirmed by experimental measurements,<br />
showing the validity of the model. The developed ohmic<br />
contact RF MEMS switch shows high isolation in the offstate<br />
and low insertion loss of in the onstate<br />
up to frequencies as high as . The presented<br />
model is intended to be integrated into standard circuit simulation<br />
software, allowing circuit engineers to design the switch<br />
bias line, minimizing induced currents and contact bouncing,<br />
as well as to find the needed dimensions of the mechanical<br />
structure, for a desired switching time and actuation voltage<br />
wave-form. Moreover, process related design rules can be<br />
automatically verified.<br />
I. INTRODUCTION<br />
RF MEMS are expected to allow for new circuit designs<br />
and higher integration density, enabling for a new generation<br />
of RF communication electronics. In particular, the RF<br />
switch excels in its high linearity and high intermodulation<br />
performance, resulting in an element for high sensitive and<br />
low noise circuitry, operating in the time-domain. Moreover,<br />
its architecture and implementation within a microstrip<br />
or coplanar waveguide offers the realization of lowloss<br />
matched circuits with a high cut-off frequency [1]. Especially,<br />
planar RF circuitry with integrated RF switches on<br />
alumina substrates has the potential to overcome the shortages<br />
of state-of-the art monolithic integrated RFIC’s, based<br />
on power consuming silicon technologies or expensive and<br />
heavy coaxial technology.<br />
On board RF switches allow passive networks to change<br />
their transfer characteristic, in order to match the impedance<br />
for highest transmit power, to tune the phase shift in<br />
an scanning antenna array or to shift filter band edges in reconfigurable<br />
filter for high isolation switches [2] [3] [4] [5].<br />
For the circuit design, modelling of these switches is<br />
compulsory. However, as compared to electronical<br />
switches, the characteristics of such micromechanical<br />
switches are determined by the coupling of their electrical<br />
and mechanical behaviour, which makes modelling much<br />
more complex. As compared to literature [6] [7], our model<br />
consists of lumped parameters and allows calculating the<br />
dynamics of movement and energy flows.<br />
II. THE EM-SYSTEM APPROXIMATION<br />
In order to understand and predict the time-dependent<br />
behaviour of a MEMS switch, a mathematical-physical<br />
model is presented.<br />
Figure 1: Underlying model of the electro-mechanical transducer<br />
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11-13 <br />
May 2011, Aix-en-Provence, France<br />
The model is represented by the mechanical and the The transformation factor is primarily valid for the<br />
electrical equations, concatenated by the coupling terms, static case. Nevertheless, is an appropriate approximation<br />
and returns the time-dependent solutions for any physical in the time-dependent case, as long the cantilever bending<br />
quantity of interest. Figure 1 shows the underlying model moment and shape is not significantly altered by the damping<br />
of the electro-mechanical transducer. The RF switch discussed<br />
action.<br />
in this paper consists of a free-standing cantilever At the time , when the contact surfaces at the tip<br />
made of gold. The bias line impedance comprises a touch each other, the velocity at the tip changes its sign, reproducing<br />
series resistor and a series capacitor , whereas the<br />
the bouncing effects, well-known from mechani-<br />
variable capacitance is formed by the two electrodes,<br />
cal relays. Attention has paid to these changing boundary<br />
of the moveable cantilever, as one electrode, and the conditions while solving (1). The non-harmonic parametric<br />
bias line electrode used to actuate the switch, separated by conditions (3) for the differential equation (1), describing<br />
the gas and eventually a dielectric thin film. A voltage drop the movement of the tip, are used to reproduce bouncing, in<br />
across the gap<br />
causes the lever to bend case when a vibration of the cantilever itself after contact is<br />
downward, and enables this way mechanical and electrical not relevant. This turns out to be true after the calculations,<br />
contact at . During the switch off, a current is fed because the vibration frequency is much lower than the<br />
back to the bias line.<br />
bouncing frequency.<br />
III. THE MECHANICAL SUBSYSTEM<br />
For the modelling of the time-dependence of the MEMS<br />
switch, a 1-dimensional model with concentrated quantities<br />
is aimed. Therefore a representative point was defined.<br />
In order to calculate the lumped coefficients ,<br />
, and the lumped force , an analytical<br />
approach was applied. References on this topic can be<br />
found throughout various literature [8] [9] [10] [11]. From<br />
the derived lumped parameters , , and<br />
the force , it becomes possible to formulate the<br />
differential equation of motion (1).<br />
(1)<br />
The solutions of (1) for returns the timedependent<br />
displacement of the cantilever at the point . In<br />
order to determine the displacement at the cantilever tip,<br />
has to be scaled to . Hence, we define a transformation<br />
factor (2), which links the displacement and<br />
velocity along the -direction at to the displacement and<br />
velocity at the tip of the cantilever and vice versa.<br />
(2)<br />
(3a)<br />
(3b)<br />
In case of a vibrating cantilever, bouncing is not expected,<br />
and the cantilevers will resonate around the steadystate<br />
equilibrium position.<br />
IV. THE ELECTRICAL SUBSYSTEM<br />
Besides its mechanical part, the RF switch consists also<br />
of an electrical counterpart. The complete modelling of<br />
both parts, including forward and backward coupling, allows<br />
studying the influence of different actuation waveforms<br />
in more detail. The bias line treated in this model<br />
comprises the two serial components: a bias line resistor<br />
and a capacitor . The voltage at the actuation electrode<br />
is given by .<br />
Whereas the current in the bias line is the sum of two<br />
components (4), on one hand introduced by the reduction of<br />
the gap<br />
and thus, by the change of the capacitance<br />
, and on the other hand by a change of the electrode<br />
potential .<br />
(4)<br />
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May 2011, Aix-en-Provence, France<br />
(5) chosen set of materials, the results are presented and exemplified<br />
in the following section.<br />
After substitution of into (5) and solving for the The actuation test signal applied to the system has a rising<br />
second derivative of the voltage a differential equation is<br />
edge of , a pulse width of and a trail-<br />
gained, containing the backward coupling, attributed to ing edge of .<br />
, and . represents<br />
the stacked dielectric, e.g. alumina with thickness adjoining<br />
A. Dynamics of Movement<br />
the air gap.<br />
Figure 2 shows the displacement curve of the cantilever<br />
at the position , whereas exhibits<br />
mechanical contact at<br />
. The trigger delay<br />
(6)<br />
of the trailing edge is set to . The velocity depicted<br />
in figure 3 shows two different resonating modes of<br />
the system, one in the down-state position, damped by the<br />
energy absorption by the contact and the squeeze film, and<br />
the second in the up-state position, damped by the squeeze<br />
V. ENERGY BALANCE<br />
film.<br />
The energy balance compares the total energy entering<br />
the system boundary with the sum of the energy<br />
components stored or dissipated in the system. Stored energy<br />
comprises the kinetic and potential component of the<br />
mechanical resonant structure , as well as the energy<br />
stored in the bias line capacitor and the moving<br />
variable capacitor . Dissipation terms are the squeezefilm<br />
damping , the bias line resistor and the absorbed<br />
energy by the contact .<br />
(7)<br />
can be calculated by integration of the absorbed<br />
momentum (31). With the momentum reflection coefficient,<br />
defined as<br />
, and the momentum<br />
absorption coefficient .<br />
(8)<br />
VI. SIMULATION RESULTS<br />
The mathematical model, implemented in Simulink, describes<br />
the switch dynamics by a 1-dimensional model. The<br />
model consists of two coupled subsystems, one describing<br />
the mechanical resonator, and the other representing the<br />
electrical bias line. For a specific switch geometry and a<br />
displacement z(l3,t) [um]<br />
1.4<br />
1.2<br />
1<br />
0.8<br />
0.6<br />
0.4<br />
0.2<br />
0<br />
-0.2<br />
-0.4<br />
-0.6<br />
0 10 20 30 40 50 60 70<br />
time [us]<br />
Figure 2: Displacement versus time of the cantilever contact tip.<br />
velocity dz(l3,t)/dt [um/us]<br />
0.5<br />
0.4<br />
0.3<br />
0.2<br />
0.1<br />
0<br />
-0.1<br />
-0.2<br />
-0.3<br />
-0.4<br />
-0.5<br />
0 10 20 30 40 50 60 70<br />
time [us]<br />
Figure 3: Velocity of the cantilever contact tip, showing oscillation in the<br />
down-state and the up-state, including damping caused by the squeezefilm<br />
and the energy absorption by the contact.<br />
130
B. Dynamics of Energy Flows<br />
For a layout designer of switching networks, figure 4<br />
provides a quantitative prediction of the energy consumed<br />
and released by the switch. Immediately after the driving<br />
voltage is applied, a major part of energy is stored in the<br />
electrical field of the variable capacitor. The energy dissipation<br />
along the bias line resistor, caused by the induced<br />
current, is around fifty times smaller than the energy dissipated<br />
by the squeeze-film damping and therefore negligible.<br />
During the transition from up-state to down-state, the<br />
potential and kinetic energy increase with increasing cantilever<br />
deflection and velocity. At the same time, energy is<br />
dissipated by the squeeze-film damping.<br />
int[Vext(t)*i(t)]dt [microWs ]<br />
3<br />
2<br />
1<br />
0<br />
0 10 20 30 40 50 60 70<br />
time [us]<br />
Figure 4: Total energy delivered to the system and fed back.<br />
From the mechanical point of view, the system can be<br />
optimised either, to provide fast switching, to dissipate<br />
minimal energy, to minimise the actuation voltage or to reduce<br />
the stored energy in the on-state of the system. From<br />
the electrical point of view, the electrical signal can be preprocessed<br />
offering a desired wave-form in order to control<br />
the re-feeded power level, or to minimize the oscillation<br />
amplitudes and contact bouncing.<br />
VII.<br />
x 10 -4<br />
MEASUREMENTS<br />
The developed ohmic contact RF MEMS switch shows<br />
high isolation in the off-state and a low insertion<br />
loss in the on-state up to frequencies as high as<br />
[12].<br />
All processing was carried out in industrial fabrication<br />
line. The measurement results validate the theoretical<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Figure 5: fabricated ohmic contact MEMS switch in CPW configuration.<br />
model presented in the previous sections. All measurements<br />
are performed with switches fabricated within the same<br />
batch (figure 5).<br />
In order to ensure the initial conditions and the assumptions<br />
used in the model, concerning geometrical aspects,<br />
scanning electron microscopy is used. Scanning electron<br />
microscopy allows identifying cantilever bending and verifying<br />
geometric dimensions.<br />
The accessible electrical triggering characteristics are<br />
described by the bias threshold voltage resulting in an electrical<br />
through-connection of the signal path, and the restoring<br />
voltage, disconnecting the signal path. When performing<br />
the measurement in ambient environment, possible interaction<br />
with humidity and contamination of organic compounds<br />
on the contact surfaces can provoke an excess of<br />
force to overcome the action of an adhering passivation<br />
layer.<br />
Finally, the model is validated by the dynamic system<br />
response: the transition time, defined as the time passed between<br />
the supply of the bias potential and the event of first<br />
charge transferred by the electrical contacts, and the switching<br />
time, defined as the time passed between the supply of<br />
the bias potential and the event of continuous charge transfer<br />
by the electrical contacts.<br />
The test setup uses a Keithley 2400 voltage source, powering<br />
the collector path of a 2SC2911 npn-transistor form<br />
SANYO. The transistor gate is controlled by a HP 3312A<br />
function generator, providing a frequency variable square<br />
signal of positive half-waves with an amplitude set to<br />
131
, ,<br />
. The rise time of the voltage, provided at the drain of <br />
the npn-transistor, biasing the DUT, is below 1 .<br />
The signal path of the DUT is excited by a sine wave<br />
with an amplitude of<br />
at a frequency of 1MHz, oscillating<br />
around ground potential. The sine signal is supplied<br />
by a HAMEG 8030. At the switched end of the signal line,<br />
the voltage is probed by a TDS 2012 oscilloscope.<br />
Transient electrical measurements (figure 6) were conducted<br />
with respect to the transition time, contact bouncing<br />
and the switching time. The probed signal line is charged<br />
due to the induced current driven by the voltage rising<br />
edge. Since the time constant of the measurement set-up is<br />
long compared to the switching time, the open signal line<br />
remains on potential, unless the switch rests in the downstate<br />
and makes ohmic contact. The transmitted sine signal<br />
indicates a transition time of and a first bouncing period<br />
of , which compares to the calculated theoretical<br />
values. The switching time is 22 . During the bouncing<br />
sequence, the signal transmission is interrupted two times.<br />
The contact time is approximately .<br />
The static electrical measurements provide information<br />
about the micro mechanical structure and the contact altering<br />
during the first 10 th of cycles. Figure 7 shows the bias<br />
threshold voltage (*) and the restoring voltage (o) of two<br />
different switches. Thereby, the switches have passed a<br />
80<br />
voltage [V]<br />
120<br />
110<br />
100<br />
90<br />
80<br />
70<br />
60<br />
50<br />
40<br />
0 5 10 15 20 25<br />
number of cycles<br />
Figure 7: Bias threshold voltage (*) and restoring voltage (o) after annealing<br />
steps.<br />
thermal annealing of at after 4 cycles and<br />
again after 11 cycles. Following the cycle number 17, the<br />
switches are stored for 3 days at ambient conditions. It can<br />
be seen, that the second annealing step has no significant<br />
effect on the bias threshold voltage and the restoring voltage.<br />
Nevertheless, the storage at ambient can essentially affect<br />
the actuation behaviour.<br />
A decrease of the bias threshold voltage correlates with a<br />
decreasing contact force necessary to achieve electrical<br />
break trough. A decreasing hysteresis between bias threshold<br />
voltage and restoring voltage is an indication of lower<br />
attractive forces at the contacts, for example caused by capillary<br />
forces.<br />
bias line voltage [V]<br />
signal path voltage [mV]<br />
60<br />
40<br />
20<br />
0<br />
-10 -5 0 5 10 15 20 25 30 35<br />
time [us]<br />
60<br />
40<br />
20<br />
0<br />
-20<br />
-40<br />
-60<br />
-10 -5 0 5 10 15 20 25 30 35<br />
time [us]<br />
Figure 6: Actuation voltage, and transmitted signal along the switched<br />
path.<br />
VIII. DISCUSSION<br />
The theoretical model is in good agreement with the experimental<br />
data, when comparing the dynamics of movement,<br />
figure 2 to figure 6.<br />
An interaction between the ambient and the contact surfaces<br />
has been proven by their impact on the actuation behaviour.<br />
Supposing the adsorption of humidity or organic<br />
compounds on top of the contact surfaces demands for a<br />
certain static contact force [8], in order to squeeze the unwanted<br />
molecules and enable electrical contact. Thus, a<br />
higher bias voltage, compared to the so called pull-in voltage,<br />
and hence, a higher velocity and bouncing momentum<br />
is expected. In the reverse process, when restoring the<br />
structure and interrupting electrical contact, adhesive forces<br />
132
11-13 <br />
May 2011, Aix-en-Provence, France<br />
have to be surmounted, settle the restoring voltage little below<br />
workers in the European project SMARTIS (smart thin<br />
the bias threshold voltage.<br />
films on alumina<br />
substrates).<br />
IX.<br />
CONCLUSIONS<br />
REFERENCES<br />
The comprehensive mathematical-physical model introduced,<br />
demonstrates clearly the dynamic mechanism of the<br />
RF switch fabricated.<br />
The bias threshold voltage can significantly differ from<br />
the theoretical pull-in voltage in presence of a surface passivation<br />
layer, formed by organic contaminations, or due to<br />
contact alloy oxidation. From the small hysteresis, we can<br />
see, that the contact force has a significant influence on the<br />
bias threshold voltage and the restoring voltage. Additionally,<br />
a bias threshold voltage, well above the theoretical<br />
pull-in voltage, corroborates this hypothesis.<br />
Nevertheless, it can’t be eliminated yet, that the real<br />
stiffness coefficient differs from the theoretical value. Considering,<br />
that in terms of failure estimation, a variation of<br />
the cantilever thickness by 3%, would lead to a 9% deviation<br />
of the stiffness constant.<br />
In summary, our model fits well the dynamics of the fabricated<br />
switch. However, no exact comparison can be made<br />
due to the presence of humidity and contaminations, which<br />
essentially can affect the dynamics and static characteristics<br />
of the fabricated MEMS switch.<br />
Future work will consider contact conditions by hermetic<br />
packaging of the RF switch.<br />
ACKNOWLEDGEMENTS<br />
The authors acknowledge the support by the innovation<br />
promotion agency CTI of Switzerland for its financial contribution.<br />
The authors also wish to thank all project co-<br />
[1] G. M. Rebeiz and J.B. Muldavin, IEEE Microwave Magazine, 2 (4)<br />
59-71 (2001).<br />
[2] A. Pothier et al., Low-Loss 2-Bit Tunable Bandpass Filters Using<br />
MEMS DC Contact Switches, IEEE Trans. Microwave Theory and Tech.<br />
53, (2005).<br />
[3] E. R. Brown, RF-MEMS Switches for Reconfigurable Integrated Circuits,<br />
IEEE Transactions on Microwave Theory and Techniques, Vol. 46,<br />
No. 11, November 1998.<br />
[4] Vijay K. Varadan, RF MEMS and their applications, Wiley-<br />
Interscience 2003.<br />
[5] A. van Bezooijen, A GSM/EDGE/WCDMA Adaptive Series-LC<br />
Matching Network Using RF-MEMS Switches, IEEE Journal of Solid-<br />
State Circuits, vol.43, no.10, October 2008.<br />
[6] Z. J. Guo, N. E. McGruer and G. G. Adams, Modeling, simulation<br />
and measurement of the dynamic performance of an ohmic contact, electrostatically<br />
actuated RF MEMS switch, Journal of Micromechanics and<br />
Miroengineering, 2007.<br />
[7] S. Halder, C. Palego, Z. Peng, J. C. M. Hwang, D. I. Forehand and C.<br />
L. Goldsmith, Compact RF Model for Transient Characteristics of MEMS<br />
Capacitive Switches, IEEE Transactions on Microwave Theory and Techniques,<br />
Vol. 57, No. 1, January 2009.<br />
[8] G. M. Rebeiz, RF MEMS Theory, Design, and Technology, John<br />
Wiley & Sons (2003).<br />
[9] I. Szabó, Höhere Technische Mechanik, Springer, Berlin 2001<br />
[10] J. B. Starr, Squeeze-Film Damping in Solid-State Accelerometers,<br />
Tech. Digest, IEEE Solid State Sensor and Actuator Workshop, 44-47<br />
(1990).<br />
[11] P. G. Steeneken et al., Dynamics and squeeze film damping of a capacitive<br />
RF MEMS switch, Journal of Micromechanics and Microengineering<br />
15, 176-184 (2005).<br />
[12] M. El Khatib, A. Pothier, P. Blondy, Packaging of RF MEMS<br />
Switching Functions on Alumina Substrate, DTIP of MEMS & MOEMS,<br />
Stesa, Italy, 26-28. April 2006.<br />
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<br />
A Closed-Loop Micromachined Accelerometer<br />
Based on Thermal Convection<br />
Alexandra Garraud, Philippe Combette, Benoît Charlot, Pierre Loisel* and Alain Giani,<br />
Institut d’Electronique du Sud – UNIVERSITE MONTPELLIER 2 – CNRS UMR 5412<br />
Place E. Bataillon, 34095 Montpellier, France.<br />
*SAGEM D.S. 72-74 Rue de la Tour de Billy, BP 72, 95101-ARGENTEUIL<br />
Abstract- In this work, we present the frequency<br />
analysis of a micromachined thermal accelerometer<br />
based on convection. Open-loop block diagram<br />
representation is first introduced to explain the sensor<br />
behavior. New sensor architecture is imagined to<br />
enhance sensor characteristics: a closed-loop<br />
configuration is designed by addition of two resistors<br />
closed to detectors. Effects on thermal sensitivity and<br />
bandwidth are investigated.<br />
I. INTRODUCTION<br />
In recent years, a new concept of accelerometer based on<br />
thermal exchanges has been intensively studied. The<br />
physical principle is based on a hot gas bubble acting as a<br />
proof mass. Under acceleration, free-convection transfers<br />
are modified and induce the bubble motion. Like other<br />
transduction mechanisms, such as piezoelectricity,<br />
piezoresistivity or capacitive sensing [1], it converts<br />
acceleration into electrical signal. But the absence of<br />
mechanical clamping between the gas and the chip induces<br />
no stress concentration. This leads to higher shock<br />
reliability. In addition, its simple structure allows low<br />
fabrication costs and competitive performances [2]–[3].<br />
Previous studies have focused on the influence of several<br />
parameters on both sensitivity and frequency bandwidth of<br />
the thermal accelerometer: nature and pressure of gas,<br />
cavity volume, detectors’ dimensions [4–6]. The highest<br />
– 3 dB bandwidth of a thermal accelerometer was 120 Hz<br />
with a standard gas filled cavity configuration [7] and more<br />
recently 320 Hz with a helium-filled cavity configuration<br />
[6]. But these high bandwidths go with low sensitivities<br />
due to a constant sensitivity-bandwidth product.<br />
This present article will present a way to improve<br />
bandwidth without reducing sensitivity by introducing a<br />
closed-loop configuration in the sensor functioning. In a<br />
first time, we will present the open-loop system and then<br />
we will explain how to improve the system with feedback.<br />
II. SENSOR PRINCIPLE<br />
The thermal accelerometer, described in Fig. 1, is based<br />
on natural free convection in a closed chamber containing a<br />
gas. It contains a heating resistor suspended over a cavity<br />
etched on silicon, providing thermal isolation. When the<br />
resistor is supplied by an electrical current, it heats up the<br />
surrounding gas creating a symmetrical temperature<br />
distribution.<br />
Fig. 1. Schematic diagram of the micromachined thermal accelerometer.<br />
When no acceleration is applied, the system is balanced<br />
so that two temperature detectors placed on either side give<br />
the same value, as shown in Fig. 2 by the straight line.<br />
When the sensor is subjected to an acceleration Γ, the<br />
temperature profile shifts, as can be seen in Fig. 2, and the<br />
balance in free-convection heat transfer is modified. The<br />
two detectors don’t measure the same temperature anymore<br />
and this temperature difference δT is correlated to the<br />
acceleration by the sensitivity S equal to δT/Γ (°C/g).<br />
Fig. 2. Schematic diagram of the micromachined thermal accelerometer.<br />
Fig. 3 shows a SEM image of the device. We notice the<br />
suspended wires standing over the micromachined cavity.<br />
The device contains two pairs of suspended bridges on<br />
each side of the heater resistor. The cavity is obtained by<br />
KOH wet anisotropic etching of the silicon (100) oriented<br />
and measures typically 1000 μm x 2000 μm for a depth of<br />
800 μm. The heater (100 μm wide) and detectors (20 μm<br />
wide) are made of a 300 nm thick platinum layer (including<br />
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11-13 <br />
May 2011, Aix-en-Provence, France<br />
Cr adhesion layer) deposited on a 500 nm thick low stress<br />
<br />
silicon nitride membrane (SiN x ) [8]. The SiN x layer has<br />
been chosen for its low stress level allowing flat standing<br />
structures. More details concerning the manufacturing<br />
process may be found in [5].<br />
Fig. 5. Typical Bode diagram of an open-loop configuration.<br />
Fig. 3. Scanning Electron Microscope (SEM) image of a thermal<br />
accelerometer, with two pairs of detectors, each of them having a different<br />
use.<br />
III. FROM AN OPEN-LOOP CONFIGURATION…<br />
The accelerometer principle can be modelled with an<br />
open-loop block diagram, as described in Fig. 4. In this<br />
case, only detectors wires as described in Fig. 3 are<br />
required, that is to say the ones closest to the heater.<br />
An acceleration Γ applied on the sensor leads to a<br />
modification of convection heat transfers among the gas.<br />
Then detectors are subjected to a temperature variation<br />
which induces a variation of their electrical resistance. As<br />
detectors are assembled in a Wheatstone bridge, it results in<br />
an output voltage variation.<br />
Fig. 4. Block Diagram of the open-loop system.<br />
With this configuration, we obtain the frequency<br />
response as the one obtained in Fig. 5 and characteristics of<br />
thermal accelerometer presented in section 2 are the<br />
following:<br />
- sensitivity in bandwidth: S 0 = 100 mV/g or<br />
0.0445 °C/g;<br />
- cut-off frequency at -3 dB: F c = 66 Hz.<br />
IV. …TO A CLOSED-LOOP SYSTEM<br />
In [6], we established that sensitivity-bandwidth product<br />
is constant for a given cavity size: it is impossible to have<br />
both a large sensitivity and a large bandwidth. A way to<br />
improve frequency bandwidth without reduction of thermal<br />
sensitivity is to adopt a closed-loop configuration. Two<br />
solutions can be designed.<br />
The first one is to include a typical thermal<br />
accelerometer with two detection wires and a sigma-delta<br />
modulator which assures the feedback [9]. But no<br />
experimental results are presented in this reference.<br />
The second solution is the one chosen here: inverse<br />
feedback is directly implemented in the sensor itself by<br />
addition of two other platinum suspended bridges, placed<br />
closed to detection wires, as shown in Fig. 3 and named<br />
feedback resistors. Their aim is to maintain the temperature<br />
difference between the two detectors, δT, equal to zero<br />
when an acceleration is applied to the sensor. This<br />
technical solution has led to two patents: [10] and [11].<br />
A current is injected in each feedback resistor to hardly<br />
increase their temperature and as a result the temperature of<br />
the closest detector. When the sensor is submitted to an<br />
acceleration, one detector gets warm while the other cools<br />
off. Thus less current is injected in the feedback resistor<br />
placed closed to the hottest: the detector temperature<br />
decreases. In the other feedback resistor, the same amount<br />
of current is added to the equilibrium current to warm the<br />
nearby detector. As a consequence the two detector<br />
temperatures are maintained in their equilibrium value.<br />
Block-diagram description of the feedback loop and its<br />
effect on detectors is presented on Fig. 6. The feedback<br />
resistor temperature is modified by the variation of the<br />
nominal injected electrical power. This induces a<br />
modification of heat-transfer around these suspended<br />
bridges which is then caught by detectors with a<br />
modification of their global temperature, δT(δP).<br />
135
Fig. 6. Block diagram of the feedback loop.<br />
Synoptic view is resumed in Fig. 7 with the complete<br />
block diagram of the closed-loop system. Abbreviations<br />
mentioned are the ones used in Fig. 4 and Fig. 6.<br />
Henceforth, output voltage is a function of the electrical<br />
power required to maintain a zero temperature difference<br />
between the two detectors.<br />
Fig. 7. Block diagram of the closed-loop system.<br />
With this closed-loop configuration, we perform a<br />
frequency analysis, visible in Fig. 8, of the accelerometer<br />
presented in section 2. A sensitivity S 1 = 76 mV/g or<br />
0.034 °C/g closed to the open-loop configuration one is<br />
obtained while the cut-off frequency has increased by more<br />
than a factor 15 with a value F c = 1025 Hz. In this case, a<br />
resonance appears that would disappear with better settings<br />
on the PID controller. Nevertheless, this bandwidth result is<br />
the larger never obtained with a thermal accelerometer. It is<br />
ten times larger that could be found in the literature.<br />
Fig. 8. Closed-loop frequency response.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
V. CONCLUSION<br />
This work investigates the behavior improvement<br />
brought by a closed-loop configuration when using a<br />
MEMS thermal accelerometer. This sensor was<br />
micromachined by micro-electronics techniques. In a first<br />
time, we studied the open-loop configuration with typical<br />
characteristics. A closed-loop structure has been conceived<br />
to improve bandwidth with no sensitivity reduction.<br />
Feedback loop is directly included in this sensor<br />
architecture by addition of two resistors close to detectors.<br />
Modulation of the electrical power injected allows<br />
temperature profile to remain symmetric. Experimental<br />
measures prove that we can achieve a large bandwidth for a<br />
system based on thermal exchanges with no modification<br />
of thermal sensitivity.<br />
REFERENCES<br />
[1] J. Fraden, Handbook of Modern Sensors. Woodbury, New York:<br />
American Institute of Physics, 1998.<br />
[2] A.M. Leung, J. Jones, E. Czyzewska, J. Chen and B. Woods,<br />
“Micromachined accelerometer with no proof mass,” Digest Tech.<br />
Papers International Electron Devices Meeting, Conference,<br />
Washington, DC, USA, December 7–10, 1997, pp. 899–902.<br />
[3] X.B. Luo, Y.J. Yang, F. Zheng, Z.X. Li and Z.Y. Guo, “An<br />
optimized micromachined convective accelerometer with no proof<br />
mass,” J Micromech Microeng, vol. 11, no. 5, pp. 504–8, 2001.<br />
[4] F. Mailly, A. Martinez, A. Giani, F. Pascal-Delannoy and A.<br />
Boyer “Effect of gas pressure on the sensitivity of a<br />
micromachined thermal accelerometer,” Sensor Actuat A-Phys,<br />
vol. 109, pp. 88–94, 2003.<br />
[5] J. Courteaud, P. Combette, N. Crespy, G. Cathebras and A. Giani,<br />
“Thermal simulation and experimental results of a micromachined<br />
thermal inclinometer,” Sensor Actuat A-Phys, vol. 141, pp. 307–<br />
313, 2008.<br />
[6] A. Garraud, P. Combette, F. Pichot, J. Courteaud, B. Charlot and<br />
A. Giani, “Frequency response analysis of an accelerometer based<br />
on thermal convection,” J. Micromech Microeng, 21 (2011)<br />
035017.<br />
[7] J. Courteaud, N. Crespy, P. Combette, B. Sorli and A. Giani,<br />
“Studies and optimization of the frequency response of a<br />
micromachined thermal accelerometer,” Sensor Actuat A-Phys,<br />
vol. 147, pp. 75–82, 2008.<br />
[8] P. Temple-Boyer, C. Rossi, E. Saint-Etienne and E. Scheid,<br />
“Residual stress in low pressure chemical vapor deposition SiNx<br />
films deposited from silane and ammonia,” J Vac Sci Technol A,<br />
vol. 16, no. 4, pp. 2003-2007, 1998.<br />
[9] O. Leman, L. Latorre, F. Mailly and P. Nouet, “A Closed-Loop<br />
Architecture with Digital Output for Convective Accelerometers,”<br />
Proceedings of IEEE Computer Society Annual Symposium on<br />
VLSI, Montpellier, France, April 07-09, 2008, pp. 51-56.<br />
[10] J. Dido, P. Loisel, A. Renault, P. Combette, J. Courteaud and A.<br />
Giani, "Thermal cell system for measuring acceleration", United<br />
States Patent 7469587, to Sagem Defense Securite (Paris, FR),<br />
2008.<br />
[11] C. Gervais, A. Renault, B. Varusio, A. Boyer and A. Giani,<br />
“Thermal measure of acceleration, speed, position or inclination<br />
uses a predetermined volume of heated fluid, compensates for<br />
reduction in temperature due to acceleration by using auxiliary<br />
heaters”, FR2832802 - 2003-05-30<br />
136
11-13 May, Aix-en-Provence, France<br />
<br />
PANEL DISCUSSION<br />
TEXTILE MICROSYSTEMS<br />
INTRODUCTION: ELECTRONICS MEETS TEXTILES - CHALLENGES AND OPPORTUNITIES<br />
Erik JUNG, Fraunhofer IZM, Germany<br />
PASTA-BRIDGING THE GAP BETWEEN TEXTILE AND ELECTRONICS<br />
Dominique VICARD, CEA-LETI, France<br />
PIEZOELECTRIC CHARGING FOR SMART FABRIC APPLICATIONS<br />
Ross HACKWORTH, R. MAXWELL, R. KOTHA, A.Arturo AYON, U. of Texas at San Antonio, USA,<br />
J. R. MORIERA, U. of California, Santa Barbara, USA<br />
HIGH VOLUME, LOW COST: TEXTILE INTEGRATION FOR RFID<br />
Christine KALLMAYER, Fraunhofer IZM, Germany<br />
METER-SCALE SURFACE CAPACITIVE TYPE OF TOUCH SENSORS FABRICATED BY WEAVING CONDUCTIVE-<br />
POLYMER-COATED FIBERS<br />
Seiichi TAKAMATSU1, Takeshi KOBAYASHI1,2, Nobuhisa SHIBAYAMA1, Koji MIYAKE1,2 and Toshihiro ITOH1,2<br />
1. Macro BEANS Center, BEANS Laboratory, Tsukuba, Ibaraki, Japan<br />
2. UMEMSME, National Institute of Advanced Industrial Science and Technology, Tsukuba, Ibaraki, Japan<br />
SMALL VOLUME, HIGH TECH: INTELLIGENT TEXTILES FOR HEALTH AND WELLNES<br />
Guofu ZHOU, Philips<br />
137
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May 2011, Aix-en-Provence, France<br />
<br />
Piezoelectric Charging for Smart Fabric<br />
Applications<br />
1<br />
R. Hackworth , J. R. Moriera, R. Maxwell, R. Kotha, and A.A. Ayon, Member, IEEE<br />
Abstract— Our current research includes an innovative design<br />
and fabrication method for a wearable piezoelectric power<br />
generating fabric. Rather than building a flexible piezoelectric<br />
device and then applying it onto clothing as reported by other<br />
groups, our approach is to integrate the devices directly with the<br />
fabric while requiring no processing temperatures over 150 ˚C.<br />
The device discussed here simply consists of a piezoelectric layer<br />
encapsulated between a bottom electrode/wearable fabric and a<br />
top layer electrode/conductive film layer.<br />
Index Terms—Charge generation and storage, polyvinylidene<br />
fluoride (PVDF), piezoelectric materials, smart fabrics.<br />
I. INTRODUCTION<br />
e report on the feasibility of employing flexible PVDF<br />
Wpiezoelectric membranes to be used to generate<br />
electrical charge for powering portable electronics. By<br />
converting some of a person’s naturally expended mechanical<br />
energy into useful electrical energy, the batteries currently in<br />
use in portable electronic devices may be minimized, as well<br />
as made efficiently green. Ideally, these devices will be<br />
wearable and light-weight. The energy generation will<br />
originate from the exploitation of the piezoelectric effect of<br />
certain materials such as PVDF due to their naturally<br />
occurring deformation as shown on Fig. 1.<br />
a commercially available fabric enable greater comfort for the<br />
end-user in a wearable energy harvester. Research has largely<br />
focused on PVDF in cyclopentanone as a solvent due to its<br />
natural flexibility compared with that of any other<br />
piezoelectric materials we tested so far (which include zinc<br />
oxide, barium titanate, et al..). For optimal piezoelectric<br />
effects the PVDF film needs to be in the β-crystalline phase<br />
[1]. PVDF is usually observed in one of the four main phases,<br />
but only the β-phase is expected to possess strong d 31 ,<br />
properties which are critical requirements in building a<br />
practical prototype. The beta phase is generally formed by<br />
poling the polymer at a high voltage while stretching it. When<br />
the alpha phase is poled the dipoles align and for the delta<br />
phase. Upon stretching the bonds in the chains will reorient<br />
themselves into an all trans configuration. To convert the<br />
largely α-phase PVDF into β-phase requires aligning the<br />
polymer chains into the all trans phase structure as in Fig. 2.<br />
To achieve this, the cured PVDF can be simultaneously<br />
stretched and poled [2, 3] while at elevated temperature to<br />
induce more of the bulk membrane into the stronger<br />
piezolelectric β- phase [4].<br />
Fig. 2. β-phase PVDF.<br />
Fig. 1. The piezoelectric effect.<br />
Multiple options for fabric substrates, bottom and top<br />
electrodes, piezoelectric materials, as well as with the methods<br />
by which they were fabricated were investigated. Our<br />
observations indicate that the most successful bottom<br />
electrode and fabric combination found so far consists of a<br />
commercial polyester fabric coated through an electroless<br />
deposition of nickel followed by copper. The characteristics of<br />
Manuscript received April 8, 2011.<br />
Thanks go to: Army Research Office<br />
UTSA MEMs lab<br />
II. MATERIALS<br />
We start with the discussion of multiple options for fabric<br />
substrates, bottom and top electrodes, and piezoelectric<br />
materials, along with the methods [5] in which they are<br />
fabricated.<br />
Many commonly available fabrics and their melting points,<br />
flash points, and decomposition temperatures were evaluated.<br />
These fabrics were initially studied due to their ability to<br />
withstand temperatures in excess of 150˚C, which is close to<br />
the upper temperature range that any PVDF processing should<br />
require. The curie temperature and melting point of PVDF are<br />
close to each other. The currently accepted answer is a curie<br />
temperature of about 150˚C and a melt temperature of 175˚C.<br />
According to the manufacturer (Solef), devices made from<br />
PVDF typically have a maximum usage temperature of about<br />
138
80˚C. Additionally, we looked into military fatigues and their<br />
fabric compositions as well as a cotton polyester fabric coated<br />
with nickel and copper. The military fatigues come in a variety<br />
of compositions which might convolute testing, but the<br />
FlecTron® Nickel/Copper fabric boasts strong characteristics<br />
such as a high surface conductivity of less than 0.1 Ohms/sq.,<br />
and can withstand temperatures up to 200°C as shown in<br />
Table 1. The fabric is also flexible, lightweight, breathable,<br />
durable, washable, and tear resistant. Since the fabric is coated<br />
with an electroless metal layer that is conductive it also serves<br />
as the bottom electrode for the piezoelectric circuit as well as<br />
the substrate/wearable garment.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
PVDF/TrFE (from Solef, see spec sheet in Appendix A) in<br />
cyclopentanone was chosen for its flexibility and durability as<br />
well as for its piezoelectric properties and required processing<br />
temperatures. Other materials tested such as zinc oxide were<br />
brittle and cracked, thus failing any fatigue testing. Others<br />
were too rigid to deform in such a way that would make them<br />
practical in this application. The cyclopentanone has a<br />
relatively high boiling point compared with other solvents, as<br />
well as a strong ability to dissolve PVDF.<br />
III. FABRICATION<br />
To prepare the PVDF solution enough powdered<br />
PVDF/TrFE) is added to a cyclopentanone solvent to produce<br />
a 15 wt% PVDF solution. The mix is then heated to 70˚C<br />
while being magnetically stirred in a sealed container until the<br />
PVDF is completely dissolved (~4 hrs). For testing purposes,<br />
membranes were created by spin coating the PVDF solvent<br />
onto a silicon wafer and allowing to cure at 70˚C for 4 hours,<br />
which produces a ~6um membrane when the spin coater is set<br />
to 500 rpm.<br />
2<br />
Table 1. FlecTron properties (from manufacturer, 2010).<br />
There are four possible crystalline phases for PVDF: alpha,<br />
beta, and gamma are the most common [6]. Of these, only the<br />
beta phase is piezoelectric. The beta phase is formed when alltrans<br />
polymer chains pack themselves into an orthorhombic<br />
crystalline lattice. Because all the fluorine atoms are on the<br />
same side of the backbone the lattice lacks a center of<br />
symmetry causing net dipole moment. Upon crystallization<br />
out of the melt the most common phase is the alpha phase. The<br />
chains in the alpha phase are of the form trans-gauche-transgauche.<br />
They also pack into an orthorhombic lattice because<br />
the fluorine atoms alternate sides of the backbone therefore<br />
exhibiting no net dipole, and thus not piezoelectricity.<br />
Fig. 4. Gold coated PVDF membrane in flexing test jig.<br />
To minimize any bubbles in the film, the just spin coated<br />
wafer was placed in a vacuum desiccators for 2 minutes prior<br />
to curing. Finally, 100 nm gold was sputter coated on the<br />
membrane at room temperature to form the top electrode as<br />
shown in Fig. 4. The bottom electrode can be sputter coated<br />
onto the silicon wafer prior to the spin coat of PVDF if<br />
desired, as it will adhere to the PVDF when the membrane is<br />
separated from the silicon substrate. The electrodes were then<br />
connected using copper wire bonded to the gold electrodes<br />
using silver conductive epoxy (cross section shown in Fig. 5).<br />
Fig. 3. Solef brand PVDF crystalline phase diagram.<br />
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11-13 May 2011, Aix-en-Provence, France<br />
<br />
Fig. 6. Schematic diagram of test set up.<br />
3<br />
Fig. 5. FlecTron conductive fabric coated with PVDF and<br />
gold upper electrode.<br />
IV. POLING<br />
To aid in the transition from α-phase to β-phase, our cured<br />
PVDF membrane was placed in a Lloyd Instruments tensile<br />
tester and stretched in length 5%. Poling plates were placed on<br />
each side of the 6um membranes. A poling voltage of greater<br />
than 25V/um is desired. The sample was then stretched, an<br />
enclosure put in place, and the environmental temperature of<br />
the enclosure was raised to 80˚C. The poling voltage was then<br />
applied for 90 minutes, before reducing the temperature and<br />
allowing the environment return to ambient lab conditions<br />
before removing the poling voltage.<br />
V. TEST FIXTURES<br />
The test set up shown in Figs. 6 & 7 consists of a linear<br />
stepper motor that cycles back and forth at microprocessor<br />
controlled frequency to cause a deformation in the test sample.<br />
The frequency is set to 1 Hz for these tests. We then start the<br />
motor deforming the 6 um thick membrane (2 inch diameter)<br />
of PVDF (combined with 100 nm sputtered gold electrodes on<br />
the top and bottom surfaces to allow for the electrode<br />
contacts) which causes a charge separation in the piezoelectric<br />
membrane that can be recorded and/or used to charge a<br />
battery. The contacts were connected to one of the following<br />
at a time: a data collection device, a rechargeable battery, an<br />
energy harvesting circuit [7], or a capacitor as required for<br />
energy storage and evaluation.<br />
The basic energy harvesting circuitry is shown in Fig. For<br />
evaluation purposes the electrodes were connected to a DAC<br />
system to determine and record the voltage output. X-Ray<br />
diffractometry (XRD), Fourier transform infra-red (FTIR), and<br />
scanning electron microscopy (SEM) [8] tools were used to<br />
evaluate the crystalline orientation for phase information, as<br />
well as thickness and consistency of the membranes.<br />
Additionally, an optical microscope was used to determine if<br />
there were any large imperfections in the polymer films.<br />
Fig. 7. Actual test set up.<br />
VI. RESULTS<br />
As shown in figure 3, the membrane generates a peak to peak<br />
voltage of approximately 200 mV, with a background noise<br />
level of ≤40 mV. Composition was verified using EDX, FTIR,<br />
and XRD as shown in fig. 8. Thickness was measured on a<br />
Rudolph ellipsometer and verified by measuring a gold<br />
sputtered cross-section of a PVDF membrane.<br />
800<br />
750<br />
700<br />
Intensity (cps)<br />
Intensity (cps)<br />
650<br />
600<br />
550<br />
500<br />
450<br />
400<br />
350<br />
300<br />
250<br />
200<br />
150<br />
100<br />
50<br />
0<br />
18 19 20 21 22<br />
116<br />
66<br />
16<br />
-34<br />
-84<br />
[3]<br />
-134<br />
18 19 20 21 22<br />
2-theta (deg)<br />
Fig. 8. XRD showing a 2θ peak of > 20.1˚.<br />
140
The XRD shows a peak above the 2θ = 20˚ angle associated<br />
with the α-phase, but well lower than the 2θ = 21˚ associated<br />
with the β-phase Clearly the PVDF hasn’t made the complete<br />
transition from alpha phase to beta phase, but it may have<br />
started the transition. In a number of samples we experienced<br />
PVDF precipitates or other imperfection, such as small voids<br />
which disqualified the samples from further testing. As for<br />
FTIR measurements no clear interpretation has been obtained<br />
between unpoled and poled PVDF for our test samples. The<br />
test sample of gold coated PVDF membrane has withstood<br />
over an hour of testing in the flexing test jig and shows no<br />
signs of fatigue yet. Ultimately, the details for applying PVDF<br />
to the fabric substrate require waiting until we identify the best<br />
method for poling the PVDF solvent while curing and in situ.<br />
One of last test suggests dominant d 33 [9] properties as it<br />
stretched in the test jig showing piezoelectricity, as opposed to<br />
the d 31 parameter that we are looking for. This is important<br />
due to the method of deformation. What is occurring is<br />
stretching as opposed to bending as would be natural with a<br />
fabric. The membrane is clearly piezoelectric, as can be seen<br />
in Fig. 8, but not yet in the β-phase required for better power<br />
generation.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
APPENDIX<br />
David Elam<br />
Greg Collins<br />
James Benson<br />
Dr. C.L. Chen<br />
Dr. A. Chabanov<br />
ACKNOWLEDGMENT<br />
4<br />
Fig. 8. Voltage generated from membrane at 1 Hz frequency.<br />
VII. CONCLUSION<br />
The membrane generated a charge due to its piezoelectricity.<br />
The in-situ poling and curing show promise largely because<br />
PVDF shrinks during curing, causing stress, strain. We<br />
continue to test new methods such as developing better in situ<br />
preparation of poling, curing, and heating simultaneously. We<br />
hope to complete the α-phase to β-phase process transition as<br />
efficiently as possible and apply that technique to the FlecTron<br />
fabric. Further research activities will consist of thin film<br />
optimization to maximize charge generation as well as<br />
extended testing to determine the useful life of the membranes<br />
produced. This is just the beginning of research into<br />
piezoelectric materials and applying what we learn towards<br />
improving our quality of life.<br />
REFERENCES<br />
[1] I. Elashmawi, N. Elsheshtawi, H. Abdelkader, N. Hakeem, Cryst. Res.<br />
Tech 2007, pp. 157–163<br />
[2] A. Kumar, and M. Perlman, J. Phys. D: Appl. Phys. 1992, pp.469-475<br />
[3] Y. Huan, Y. Liu, and Y. Yang, Polymer Eng. And Science 2007,<br />
pp.1630-1633<br />
[4] V.Sencadas, S. Lanceros-Mendez, J. Mano, Thermochimica Acta 2004,<br />
pp.201-207<br />
[5] V. Kochervinskii, V. Volkov, and K. Dembo, Physics of Solid State<br />
2006, pp. 1083-1085 (any par)<br />
[6] W. Yu, Z. Zhao, W. Zheng, B. Long, Q. Jiang, G. Li, X. Ji, Polymer<br />
Engineering and science 2009, pp.491-498<br />
[7] Y. Liu, G. Tian, Y. Wang, J. Lin, Q. Zhang, and H. Hoffmann, J.<br />
Intelligent Mat. Systems and Structures 2009, pp. 575-585<br />
[8] W. Ma, J. Zhang, X. Wang, and S. Wang, Applied Surface Science 2007,<br />
pp. 8377-8388<br />
[9] V. Kochervinskii, Crystallography Reports 2003, pp. 649- 675<br />
[10] Z. Wang, and J. Miao, J. Phys D: Appl. Phys. 2008, pp. 41-47<br />
141
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Meter-scale surface capacitive type of touch sensors<br />
fabricated by weaving conductive-polymer-coated<br />
fibers<br />
Seiichi Takamatsu 1 , Takeshi Kobayashi 1,2 , Nobuhisa Shibayama 1 , Koji Miyake 1,2 and Toshihiro Itoh 1,2<br />
1. Macro BEANS Center, BEANS Laboratory, Namiki 1-2-1, Tsukuba, Ibaraki 305-8564, Japan<br />
2. UMEMSME, National Institute of Advanced Industrial Science and Technology, Namiki 1-2-1, Tsukuba, Ibaraki<br />
305-8564, Japan<br />
phone: +81-29-868-3883, e-mail; stakamatsu@beanspj.org<br />
Abstract-We report on surface capacitive type of touch sensor<br />
fabric for large-area electronic devices. The fibers on which<br />
conductive and dielectric polymers were coated, were woven as<br />
wefts and warps in the pitch of 5 cm. The woven fabric sensed<br />
surface capacitances between the conductive polymer-coated<br />
fibers and human fingers and then the touched point was<br />
detected. To process long fibers (>100 m), we developed the<br />
die-coating technology applied to plastic fibers for coating<br />
conductive polymer of PEDOT:PSS and dielectric one of<br />
UV-curable adhesive. The resultant fibers were woven with<br />
automatic looming machines, forming meter-scale devices (1.2<br />
m × 3 m). The fabricated sensor fabric was examined on the<br />
detection of human touch. Then, the sensor observed surface<br />
capacitance of about 0.5 pF by touching sensors with a human<br />
finger. Therefore, our sensor will lead to meter-scale touch<br />
sensors and input devices for various electronic devices.<br />
Keywords: Large area electronics, die-coating, weaving,<br />
PEDOT:PSS, capacitive sensor<br />
I. INTRODUCTION<br />
In recent times, electronic textiles (e-textiles) that integrate<br />
sensors, actuators, antennas, and computers into fabrics have<br />
gained considerable attention because they have the<br />
advantage of instantly obtaining and providing information<br />
to humans [1-5]. In previous studies [1-3], touch sensors<br />
were integrated into clothing and they functioned as fabric<br />
keyboards of wearable computers. Another study [3]<br />
reported that push sensors were embedded into a carpet to<br />
provide humans with a sense of position, and computers and<br />
antennas were integrated into the fabric to enable<br />
communication. In addition to the advantage of e-textile’s<br />
functionality, textiles can cover extremely large-area (> 1<br />
m 2 ) because meter-scale fabric is woven by automatic<br />
looming machines. On the other hand the present MEMS<br />
sensor fabrication process can be applied for several inch<br />
wafers, but meter-scale wafer can not be processed because<br />
of the vacuum chamber size of manufacturing tools. The<br />
technology for printed circuit board can not be applied for<br />
meter scale devices due to its size of processing tools. The<br />
large area processing machines for liquid crystal display<br />
Figure 1. A conceptual sketch of large area touch sensors for detecting<br />
human motion. The sensor consists of conductive polymer and<br />
dielectric polymer-coated nylon fibers. The fibers are woven, forming<br />
sensor array sheet.<br />
offers highly integrated devices including pixel element and<br />
switching transistors, but its fabrication cost is extremely<br />
high in comparison with the fabrication of simple structures<br />
like touch sensors. Therefore, e-textiles are preferable for<br />
fabrication of large area sensors.<br />
Among these e-textiles, touch sensors play a key role<br />
because they function as human interface devices and<br />
human monitoring sensors in computers. In previous studies<br />
[1-3], many touch sensors have been developed by weaving<br />
copper wires as wefts and warps because of their simple<br />
structure and fabrication process. This type of touch sensors<br />
detect the electronic connection between wires under the<br />
applied pressure. However, the sensitivity of wire-woven<br />
sensors to touch input is not stable because the gaps between<br />
142
fibers which define the electric contact of touch sensors were<br />
formed in deformable fabric structure and were easily<br />
changed. Thus, more stable touch sensor has been required.<br />
Moreover, for the fabric type of sensor, copper wires are<br />
relatively hard to weave and more flexible material is<br />
preferable. In this study, we propose large-area and surface<br />
capacitive type of touch sensors for stable sensing<br />
mechanism (figure 1). The sensors detect the induced<br />
capacitances between fibers and human fingers and the gap<br />
between fibers does not affect the sensitivity (figure 2). The<br />
surface capacitance detection method is utilized for iphone<br />
or other portable systems for detecting human input with<br />
fingers. As a sensor structure, we proposes the conductive<br />
polymer of PEDOT:PSS and dielectric film of UV-curable<br />
adhesive are coated on the fibers and the resultant fibers are<br />
woven for forming sensor fabric. Since conductive polymer<br />
has the advantage of high flexibility, it is highly compatible<br />
to fabric. To solve the problem of the conventional MEMS<br />
process on the large area and high cost, a die-coating method<br />
to coat PEDOT:PSS and UV-curable adhesive on fibers is<br />
proposed, as a continuous coating process for long fibers.<br />
The solutions containing PEDOT:PSS and UV curable<br />
adhesive are put in the die and fibers travels through the die,<br />
coating the polymers on the fibers. And finally, coated fibers<br />
are woven with the automatic looming machines. Then, the<br />
fabricated sensor is examined on the sensitivity to sense the<br />
touch input.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Figure 3. Fabrication process of large area touch sensors. (1)<br />
die-coating of PEDOT:PSS layer on nylon fiber, (2) Die-coating of<br />
UV-curable adhesive, (3) Plain weaving of resultant fibers with<br />
automatic looming machines<br />
Ⅱ. SENSOR DESIGN AND FABRICATION PROCESS<br />
Figure 2. Structure of fabric touch sensors and mechanism of surface<br />
capacitance type of touch sensors. Sensors detect capacitance change<br />
between fibers and fingers because human fingers are conductive and<br />
work as an electrode. The capacitor is formed between conductive film<br />
on fibers and human fingers.<br />
The sensor fabrication consists of reel to reel coating process<br />
of conductive and dielectric polymer with die-coating and<br />
weaving the resultant fibers (figure 3). The sensor structure<br />
where conductive polymer and dielectric films is coated<br />
onto the fibers and they are woven as warps and wefts. The<br />
fibers we used are 470 μm-diameter nylon, which is<br />
commercially available fishing line. Conventional fibers are<br />
manufactured with a standard of g/km and are not easy to<br />
apply to certain diameter dies, but fishing line is<br />
manufactured with a standard diameter and it is easy to fit<br />
the diameter of the die to the diameter of the fibers<br />
(Standards of the Japan Fishing Tackle Manufacturers<br />
<strong>Association</strong>). Conductive polymer of PEDOT:PSS is coated<br />
on the fiber and its thickness is around 300 nm because most<br />
of PEDOT:PSS electrodes are coated with the thickness of<br />
100 nm. The dielectric layer is 10 μm thick UV-adhesive<br />
polymer. The touch sensors’ spatial density is a pitch of 5 cm.<br />
The areas except for the PEDOT:PSS-coated fibers are filled<br />
with 205 μm diameter pristine nylon fibers to retain the<br />
shape of the fabric. The fabricated fibers are woven in the<br />
manner of plain weaving whose structure is simplest. The<br />
fibers are crossed alternately. The method to determine the<br />
pushed point is based on the detection of capacitance<br />
changes of all fibers. After detecting the pushed vertically<br />
placed fiber and horizontally placed fiber, the crossed point<br />
143
was the pushed place.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Ⅲ. CHARACTERIZATION OF FABRICATION PROCESS<br />
The die-coating system and weaving system were<br />
characterized. In the die-coating system, we established a<br />
system of die-coating to form PEDOT:PSS and UV-curable<br />
adhesive-coated fibers that conformed to two specifications<br />
including length and film thickness. The system we<br />
developed, consisted of a winding machine to continuously<br />
transfer fibers, a die to coat PEDOT:PSS, and a heater or UV<br />
light to dry the solvent. The system was required to be tested<br />
to check whether it could form fibers that met the<br />
specifications. First, the die’s diameter for defining the film<br />
thickness is needed to be tuned. The thickness of<br />
PEDOT:PSS should be several hundreds of nm while that of<br />
UV-curable adhesive should be several micrometers.<br />
Second, to make fibers that were hundreds of m-long, the<br />
traveling speed of the fiber had to be considered in terms of<br />
throughput. If the traveling speed was as slow as the speed<br />
with sputtering, which is 30–40 centimeters of chamber size<br />
per 4–5 h for one process duration, it took hundreds of<br />
processes and thousands of hours to coat 100 m-long fibers.<br />
Therefore, a traveling speed of several m/min. was required<br />
to complete a long fiber. In fast-speed deposition, a speed<br />
that was too fast did not enable the wet PEDOT:PSS solution<br />
to follow the fiber. The maximum speed to form a uniform<br />
film was defined, tested, and confirmed so that it satisfied<br />
requirements. Finally, long fibers were actually fabricated<br />
under the defined conditions for the diameter of the die and<br />
the maximum traveling speed.<br />
The die-coating system which was used is shown in figure<br />
4. The machine at both sides is a winding machine<br />
(Factory-Automation Electronics Inc.) for moving fibers<br />
from left to right. To prevent fibers from hanging loosely,<br />
they are transferred under a certain tension by using bobbins<br />
with pressure sensors. The machine is operational at speeds<br />
of up to 50 m/min. The machine in the middle has a die and a<br />
heater. The die consists of a PEDOT:PSS or UV-curable<br />
adhesive reservoir and a nozzle for coating them onto the<br />
fibers. Because the die surrounds fibers with a certain gap,<br />
the wet film was coated on all surfaces of the fiber. The die<br />
had holes in the cap and nozzle. The diameter of the hole in<br />
the cap was 1 mm while those on the nozzle, which were<br />
larger than the 470-μm fiber diameter and ranged from 500<br />
to 980 μm were prepared through machining provided by the<br />
Daiichi-Daiya Company. The resultant gap between fibers<br />
and dies ranges from 0-300μm. The film thicknesses of the<br />
wet PEDOT:PSS and UV-curable adhesive are defined by<br />
the gap between the diameter of the nozzle and that of the<br />
fiber. The thickness of PEDOT:PSS was tuned from 0 to 500<br />
nm. The thicknesses were proportional to the gaps, but the<br />
large gaps induce unevenness of the coated PEDOT:PSS<br />
film. Therefore, the thickness below 200 nm is preferable for<br />
offering even film. On the other hand, UV curable polymer<br />
was also coated with die-coating system by changing the<br />
dies diameters. The diameters were chaged form 486 to 680<br />
Figure 4. Die-coating system. The both sides are yarn winding<br />
machines which travel fibers from left to the light. In the middle, dies<br />
and heaters or UV light were placed for coating PEDOT:PSS or UV<br />
curable adhesive. In the photograph, PEDOT:PSS was coated. The<br />
dies consists of coating nozzle and reservoir of the coating solution.<br />
Figure 5. The structure of the dies. The coating film thickness is defined by the<br />
gap between nozzle diameter of dies and the diameter of Nylon fiber. The<br />
coated wet film was dried by heater or UV light and the dried thickness was<br />
reduced by evaporating solvent.<br />
μm. The film thickness were tuned from 0 to 45 μm. The<br />
thickness was enough thick because the required thickness is<br />
5 μm. The thicknesses were larger than that of PEDOT:PSS<br />
because the PEDOT:PSS solution contained 99 % of water<br />
solvent and the solid content of PEDOT:PSS was only 1 %<br />
while the solution of UV-curable polymer contains almost<br />
100% solid content.<br />
Fast throughput for coating PEDOT:PSS films is essential<br />
to make long electrode-coated fibers. For fibers that are<br />
more than 100-m long, a traveling speed of at least several<br />
m/min is required. However, a fast traveling speed induced<br />
unevenness in PEDOT:PSS and UV curable adhesive film,<br />
resulting in defects. Defects in the electrodes were<br />
problematic because insensitive areas were formed in the<br />
sensors. To avoid these defects, the maximum traveling<br />
speed was experimentally evaluated. The PEDOT films<br />
were coated with the 680 μm die by changing traveling<br />
speeds in a range from 5 to 50 m/min. The photos of the<br />
films coated in the different traveling speeds were taken. It<br />
was confirmed that the films were even until the speed of 50<br />
144
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Figure 6. The relation ship between the gaps and PEDOT:PSS thicknesses. The<br />
thickness of the PEDOT:PSS was tuned from 0 to 500 nm by changing the gap<br />
from 0 o 300 μm.<br />
Figure 9. Automatic looming machine. The pristine fiber and conductive and<br />
dielectric film coated fibers were woven in the structure of plain weaving.<br />
Figure 7. The relation ship between dies diameters and UV-curable polymer<br />
thicknesses. The thickness of the UV curable adhesive was tuned from 0 to 45<br />
μm by changing the dies diameter from 485 to 680 μm.<br />
Figure 10 Woven fabric touch sensors. The PEDOT:PSS and UV-curable<br />
adhesive-coated fibers were woven in the pitch of 5 cm. The sensor width and<br />
length were 1.2 m and 3m, respectively.<br />
Figure 8. The relationship between the traveling speed of fibers and<br />
PEDOT:PSS thickness. Even if the traveling speed increased upto 50 m/min,<br />
the thicknesses of PEDOT:PSS were almost same.<br />
m/min. In case of UV curable adhesive, it was also<br />
confirmed that the films were even until 50 m/min.<br />
Therefore, the throughput of the fiber processing met the<br />
requirements for the meter scale fabric sensors.<br />
Ⅳ. WEAVING THE SENSOR FABRIC<br />
The processed nylon fiber with PEDOT:PSS and<br />
UV-curable adhesive was woven with automatic looming<br />
machine, forming touch sensor sheet. The automatic<br />
looming machine was made for looming the processed fibers.<br />
The fibers were very week for friction during the weaving.<br />
When the wefts were threaded in the warps, the wefts were<br />
easily fiber got stuck with the warps, resulting in the<br />
separation of the coated films. To solve the problems, the<br />
machine had the linear actuator to move the wefts between<br />
warps (figure 9).<br />
The fibers were woven in the manner of plain weaving by<br />
using the developed weaving machines. The coated fibers<br />
were placed in the pitch of the 5 cm. The width of the fabric<br />
is 1.2 m and the length was 3 m. Figure 10 shows the<br />
fabricated fabric and the size is meter scale.<br />
Ⅴ. CHARACTERIZATION OF FABRICATED TOUCH SENSOR<br />
The fabricated sensor detects the capacitance change<br />
between fiber and human finger. A human finger works as<br />
an electrode for the capacitor formed with conductive<br />
polymer-coated fibers because human is large-scale<br />
conductor. In the capacitance measurement, the potential<br />
was applied to the fiber and the flown current between fiber<br />
and human was detected for estimating formed capacitors.<br />
To detect the capacitance change between fiber and finger,<br />
the capacitance measurement circuit was connected to the<br />
fibers. Recently, capacitance measurement circuit was<br />
embed on the conventional MCU for touch sensor on<br />
portable electronic devices like i-phones. In our experiment<br />
145
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Silicon laboratories C8051F700 was used for capacitance<br />
inversely proportional to the width of the objects.<br />
Figure 11. Experimental setup of touch sensor characterization. Woven fabric<br />
touch sensor was connected to the Silicon laboratories MCU8051F700. MCU<br />
measured the capacitance of fibers and transfer it to PC through UART.<br />
Figure 12. Response of capacitance change under touch input. The capacitance<br />
increased when the fiber was touched with human finger.<br />
Figure 13. The relationship between width of object and capacitance change.<br />
The capacitance change was proportional to the width of the objects.<br />
measurement circuit. Figure 11 shows the experimental set<br />
up. The capacitance measurement pin was connected to the<br />
PEDOT:PSS electrode on fibers by removing the UV<br />
curable adhesive and connecting the PEDOT:PSS layer with<br />
conventional copper wire. The MCU measured the<br />
capacitance and transferred the data to PC through UART<br />
port. The figure 12 shows the measured capacitance when<br />
the fiber was touched by human finger. The capacitance was<br />
changed from 34 pF to 36 pF. The capacitance change was<br />
about 1-2 pF. The capacitance change of 2 pF is<br />
conventional among the capacitive type of touch screens<br />
used for cell phone or portable electronic devices.<br />
The response of capacitive sensor was defined by the area of<br />
the touched sensor area. Because the diameter of the fiber<br />
was fixed to be 475 μm, the area was defined by the width of<br />
the object. Therefore, the different width of the objective<br />
electrode was fabricated with conductive rubber and placed<br />
on the fabricated sensor. Figure 13 shows the relationship<br />
between width of objects and capacitance change. The<br />
chapacitance was proportional to the width of the objects.<br />
The capacitance change ranges from 0.9 to 2.0 pF by<br />
changing the width from 5 to 50 mm. Since the human<br />
finger size is 20 mm, the sensor can detect human touch<br />
input. On the other hand, because the PEDOT:PSS is high<br />
electric resistance in comparison with conventional metals,<br />
the detected capacitance was decreased by the length of the<br />
fibers. Therefore, we detected capacitance change in the<br />
different lengths between pointed place and the<br />
measurement circuit. Figure 14 shows the relation ship<br />
between length of the fiber and the detected capacitance<br />
change. The capacitance change was inversely proportional<br />
to the length. Therefore, large area sensors are required for<br />
the low electric resistance and the fabricated sensors can<br />
detect the touched point with the length of 40 cm because the<br />
capacitance change of the touched point with the length of<br />
50 cm is very small.<br />
Finally, we demonstrated key board system with fabricated<br />
touch sensor fabric. Key board system consisted of 3 x 9<br />
sensor fabric, MCU and PC (figure 15). The 3 x 9 sensor<br />
array was used as a key board. Layout of the keyboard is<br />
standard QWERTY. Figure 16 shows the demonstration of<br />
the keyboard input. Typed alphabet was displayed on the<br />
PC.<br />
Figure 14. The relationship between the length between touched point and<br />
measurement circuit and the capacitance change. The capacitance change was<br />
146
Figure 15. Key board input system. The System consists of 3 x 9 touch sensor<br />
array, MCU and PC.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
[5] Abouraddy, A. F.; Shapira, O.; Bayindir, M.; Arnold, J.; Sorin, F.;<br />
Hinchzewski, D. S.; Hoannopoulos, J.; Fink, Y. Large-scale optical-field<br />
measurements with geometric fibre constructs. Nature materials 2006, 5,<br />
532-536.<br />
Figure 16. Demonstration of key board input. The typed alphabet was displayed<br />
on the PC.<br />
Ⅵ. CONCLUSIONS<br />
In summary, we proposed surface capacitive type of fabric<br />
touch sensor for large-area electronic devices. In the sensor<br />
structure, PEDOT:PSS and UV-curable adhesive-coated<br />
fibers were woven as wefts and warps. The die-coating of<br />
PEDOT:PSS and UV-curable adhesive was developed to<br />
continuously form functional material on fibers. The<br />
weaving with automatic looming machine was employed for<br />
constructing meter-scale sensor fabric in continuous manner.<br />
Then, the 1.2 m × 3 m sensor fabric was woven. The sensors<br />
could detect human touch by measuring surface capacitance<br />
between human fingers and fibers. The values of capacitance<br />
change under touch input was 1-2.0 pF which is easy to<br />
detect by conventional capacitance meters that were<br />
integrated in MPUs. The developed sensor structure and<br />
fabrication process will lead to large area touch sensors in<br />
various electronic devices.<br />
ACKNOWLEDGMENT<br />
The research “Development of Continuous<br />
Nano/Micromachning and Integration Process for Fiber<br />
Substrates” has been being conducted as one of the research<br />
items of New Energy and Industrial Technology<br />
Development Organization (NEDO) project “Development<br />
of Manufacturing Technologies for Hetero Functional<br />
Integrated Devices “(BEANS project).<br />
REFERENCES<br />
[1] Marculescu, D.; Marculescu, R.; Zamora, N; Stanley-Marbell, P.; Khosla,<br />
P.; Park, S.; Jayaraman, S.; Jung, S.; Lauterbach, C.; Weber, W.; Kirstein, T.;<br />
Cottet, D.; Grzyb, J.; Troster, G.; Jones, M.; Martin, T.; Nakad, Z. Electronic<br />
textiles: a platform for pervasive computing. Proc. of IEEE 2005, 91,<br />
1995–2018.<br />
[2] Post, E.; Orth, M.; Russo, P.; Gershenfeld, N. E-broidery: design and<br />
fabrication of textile-based computing. IBM system journal 2000, 30, 840–860.<br />
[3] Gould, P. Textiles gain intelligence. Materials Today 2003, 38-43.<br />
[4] Catrysse, M.; Puers, R.; Hertleer, C.; Van Langenhove, L.; Van Egmond H.;<br />
Matthys, D. Towards the integration of textile sensors in a wireless monitoring<br />
suit. Sensors and Actuators A 2004, 114, 302-311.<br />
147
11-13 May 2011, Aix-en-Provence, France<br />
On-Wafer-Packaging of Crystal Quartz Based<br />
<br />
Devises Using Low-Temperature Anodic Bonding<br />
Y. Zimin, T. Ueda<br />
Graduate School of Information, Production and Systems, Waseda University<br />
2-7 Hibikino, Wakamatsu-ku, Kitakyushu-shi, Fukuoka 808-0135, Japan,<br />
Email: zimin-yura@fuji.waseda.jp<br />
Abstract- Low-temperature bonding of crystalline quartz and<br />
silicon wafers is described. The bonding has a big potential for<br />
MEMS applications because it could integrate the processing and<br />
packaging in a single high-tech process. In this work, strong<br />
bonding of silicon and crystal quartz wafers close to the<br />
mechanical strength of the initial materials has been achieved.<br />
Tensile test shows a disruptive stress of the samples at about 35<br />
MPa. High bonding strength is associated with minimization of<br />
the residual stresses, optimization of surface activation, and<br />
application of an electric field during annealing. Lowest possible<br />
annealing temperature and the optimum thickness ratio of<br />
silicon and quartz layers have been used in order to minimize the<br />
residual stresses.<br />
I. INTRODUCTION<br />
Wafer bonding is coming into wide use in MEMS<br />
technology. One of the most important candidates for bonding<br />
is a pair of silicon-crystalline quartz. Quartz is widely used for<br />
generators, high frequency filters, gyroscopes and<br />
microbalances because its physical properties are extremely<br />
stable. Conventional fabrication of devices based on quartz<br />
consists of a high tech processing in the very crystal with<br />
electrodes and subsequent manual assembling to the package.<br />
The manual assembling could be eliminated through<br />
integration of the processing and packaging in a single<br />
high-tech process by means of silicon/crystal quartz bonding.<br />
The integration could also provide a miniaturization and<br />
significantly improve parameters and quality of ready-made<br />
devices. High-temperature direct bonding is well known and<br />
provides a strong coupling and low level of residual stresses<br />
for materials with identical thermal expansion coefficients.<br />
High temperature increases a mobility of atoms across the<br />
interface that largely determines a strong bonding. When<br />
bonded structure consists of materials with different thermal<br />
expansion coefficients, excessive internal stresses may arise at<br />
the interface as result of high annealing temperature. Silicon<br />
and quartz are requiring the processing temperature as low as<br />
possible because the thermal expansion coefficient mismatch<br />
is quite large. Moreover, preprocessed wafers should not be<br />
exposed to high temperature in order to avoid the damage of<br />
the structures. The preprocessed structure could be also<br />
sensitive to residual stresses that can lead to subsequent<br />
degradation of the structure. Therefore, the development of a<br />
low-temperature technology is a key requirement of the strong<br />
bonding of dissimilar materials such as silicon and quartz pair.<br />
The thorough preparation of the surfaces for each specific<br />
pair of dissimilar materials can be an alternative to<br />
high-temperature annealing. The most promising results are<br />
achieved when the surface preparation includes a plasma<br />
treatment [1-3]. Even such dissimilar materials as crystalline<br />
silicon and lithium niobat show relatively strong bonding at<br />
room temperature as result of surface activation [1].<br />
Low-temperature technology can essentially reduce the<br />
residual stresses, but does not completely eliminate them for<br />
materials with different thermal expansion coefficients.<br />
Operating conditions of MEMS devices should include a<br />
temperature range as wide as possible. In this connection,<br />
internal stresses distribution must be given proper weight in<br />
designing the bonded structure. This work aims to produce a<br />
strong bonding of silicon-quartz structures with the lowest<br />
possible residual stresses. The experiment was performed by<br />
plasma-assisted activation of silicon and quartz surfaces, with<br />
further annealing in the electric field. Strong bonding, close to<br />
the mechanical strength of the initial materials, has been<br />
achieved.<br />
II. RESIDUAL STRESS IN BILAYER SYSTEM<br />
Stoney’s [4] and Timoshenko’s [5] formulas are often used<br />
to calculate the residual stresses in layered structures. Stoney<br />
analyzed the model of a thin film deposited on thick substrate.<br />
Timoshenko's approach looks the most appropriate for the<br />
bonding because it imposes no restrictions on the thickness of<br />
the layers. This model was originally developed for analysis<br />
of operation of a bimetal strip thermostat and based on the use<br />
of the radius of curvature ρ of a structure which is curved as<br />
result of a difference ∆α of the thermal expansion coefficients<br />
of the layers. The model is also appropriate for description the<br />
residual stresses under bonding of the plates of dissimilar<br />
material at elevated temperature because the bonded wafers<br />
usually have comparable thicknesses in the range between 0.1<br />
mm and 1 mm. In the case of the bonding, ∆T means a<br />
difference between annealing temperature and room<br />
temperature, or more precisely, concrete operating<br />
temperature of the bonded structure.<br />
Let h 1 and h 2 be the thicknesses of bonded plates, E 1 and E 2<br />
are their Young’s modulus, and ∆T is the difference between<br />
annealing temperature and operating c temperature of the<br />
bonded structure. Then the radius of curvature of the strip of<br />
unit width will be [5]<br />
148
ρ =<br />
h<br />
2∆α∆T +<br />
+ E 1h 3 1 +E 2 h 3 2 (1⁄ E 1 h 1 +1/E 2 h 2 )/6(h 1 +h 2 )<br />
∆α∆T<br />
<br />
<br />
. (1)<br />
Using (1), the residual stresses can be calculated according<br />
to the condition that, on the interface, the unit elongation<br />
occurring in the longitudinal fibers of both materials must be<br />
equal.<br />
α 1 ∆T +<br />
P<br />
+ h 1<br />
= α E 1 h 1 2ρ<br />
2∆T −<br />
P<br />
− h 2<br />
. E 2 h 2 2ρ (2)<br />
11-13 May 2011, Aix-en-Provence, France<br />
For the case of E 1 =E 2 , the principal residual stresses on the<br />
external faces of the structure and at its interface are<br />
calculated in [6]. As is seen in Fig. 1 (b), the stresses on both<br />
sides of the interface σ 1 and σ 2 individually depend on the<br />
thicknesses of layers. At the same time, the total stress<br />
∣σ 1 −σ 2 ∣ at the interface does not depend on the thickness<br />
(Fig. 1(b), straight dashed line). In Fig. 1, the tensile stresses<br />
are considered positive (arrows point to the right),<br />
compressive stress are considered negative (arrows point to<br />
the left).<br />
The general case of arbitrary E 1 and E 2 has been considered<br />
in [7]. The stresses on the both sides of the interface for a pair<br />
of quartz/silicon are shown in Fig. 1(c). The curves σ 1 , σ 2 , and<br />
∣σ 1 −σ 2 ∣ on Fig. 1(c) are plotted for modified Young’s<br />
moduli E * 1 = E 1 /(1-ν 1 ) and E * 2 = E 2 /(1-ν 2 ) instead of E 1 and<br />
E 2 , where ν 1 and ν 2 the Poisson ratios, because precisely the<br />
modified moduli should be used for calculations of plate<br />
deformation [5]. The principal difference between Fig. 1 (a)<br />
and (b) is the dependence of the total stress at the interface<br />
∣σ 1 −σ 2 ∣. Contrary to Fig 1(b), the total stress ∣σ 1 −σ 2 ∣<br />
depends on the thickness ratio for the case when E1<br />
≠E2 (Fig. 1(c)). The latter gives an additional opportunity to<br />
reduce the residual stresses. For example, if h 1 /h 2 =0,2, the<br />
residual stresses on the interface will be approximately 20%<br />
lower than in the case when both wafers are of equal in<br />
thickness. But this opportunity takes place exclusively due to<br />
the inequality of E * 1 and E * 2, as it take place for silicon/quartz<br />
and many other pairs.<br />
III. EXPERIMENTAL<br />
In the experiment, silicon wafers 0.5 mm thick and crystal<br />
quartz wafers 0.1 mm thick were used (h 1 /(h 1+ h 2 ) =1/6). The<br />
structures were fabricated by standard photolithography and<br />
wet etching. Prior to plasma activation, the silicon wafers<br />
were cleaned in two stages. In the first stage, organic<br />
contaminations were removed successively with acetone,<br />
isopropyl alcohol, and in an ultrasound bath with water.<br />
Afterwards, the wafers were dried in N 2 gas. The quartz<br />
wafers were cleaned in a mixture H 2 SO 4 : H 2 O 2 = 3:1 at 110 o C,<br />
distillated water at 80 o C, rinsed in DI water, and dried in N 2<br />
Fig.1. (a)- schematic diagram of bilayer bonded structure; (b)- the<br />
principal residual stresses on the interface of bilayer structure for<br />
E1=E2=185 GPa, ∆T=100 o C, and ∆α=2.05•10 -6 / 0 C as a function of<br />
normalized thickness h 1/(h 1+h 2); (c)- the principal residual stresses on the<br />
interface of quartz/silicon structure for ∆T=100 o C, E* 1=86.4 GPa, E* 2=256.9<br />
GPa, ν 1=0.17, ν 2=0.28, ∆α=2.05•10 -6 / 0 C as a function of normalized<br />
thickness h 1/(h 1+h 2).<br />
gas. Plasma exposure of both silicon and crystalline quartz<br />
plates was held in a reactive ion etcher (SAMCO RIE-10NR)<br />
with the RF generator 200W maximum power. Oxygen<br />
plasma was chosen because of its effectiveness in the<br />
activation of the surface [8]. Immediately after plasma<br />
exposure, Si and quartz wafers were brought into contact, then<br />
clamped between two stainless steel plates with well-polished<br />
surfaces and placed in a heater for annealing. Identically<br />
prepared samples were annealed under pressure of about 25<br />
KPa and temperature 130 0 C during 8 hours. In addition, DC<br />
voltage of about 300V was imposed across the specimen<br />
during annealing, as is shown in Fig. 2.<br />
149
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
Fig.2. Schematic drawing of the setup for annealing in electric field.<br />
After annealing, the bonded pairs are diced into 12x12 mm<br />
pieces for tensile strength measurements. To produce a<br />
specimen for the tensile test, two socles were glued to both<br />
sides of the bonded pair by epoxy resin. Prior to gluing, socle<br />
surfaces were sand blasted and cleaned in an ultrasonic bath<br />
with acetone. The schematic diagram of the tensile test is<br />
shown in Fig. 3. The samples were loaded gradually until they<br />
burst.<br />
IV. RESULTS AND DISCUSSION<br />
The bonding process consisted of the following successive<br />
stages: cleaning of the wafer, surface plasma activation,<br />
connecting the surfaces at room temperature and annealing.<br />
Taking into account [8], an attempt was made to rinse the<br />
wafers with DI water immediately after the plasma activation.<br />
However, this rinsing drastically reduced bonding strength in<br />
our experiment. Therefore, the rinsing was excluded from the<br />
bonding process in the following experiment. Next, the<br />
influence of etching regimes was examined: flow rate and the<br />
pressure of the oxygen inside of ion chamber. It has been<br />
found that bonding strength is not changed in the 40 to 150<br />
milliliters per second range of oxygen flows for the ion<br />
camera used. The oxygen pressure affects the bonding more<br />
significantly than the flow rate. The strongest bonding can be<br />
achieved in a narrow interval of oxygen pressure at<br />
approximately 5Pa. Moreover, testing shows that the samples<br />
activated at this plasma pressure keep bonding when the<br />
epoxy glue brakes.<br />
In addition to conventional technology of silicon and quartz<br />
bonding, the electric field was utilized over the course of<br />
annealing. Annealing in the electric field (anodic bonding)<br />
provides a very high bonding strength, close to the<br />
mechanical strength of the initial bulk materials. Anodic<br />
bonding efficiency is associated with the high mobility of<br />
alkali ions at a high temperature, the movement of which is<br />
controlled by an electric field. However alkali atoms are<br />
absent in silicon and quartz wafers. High temperature is<br />
Fig. 3. Schematic drawing of tensile test.<br />
unacceptable because it causes an internal stress in the bonded<br />
materials with a significant difference in the thermal<br />
expansion coefficient. Nevertheless, we have used the electric<br />
field and have verified that the electric field has a favorable<br />
effect on the bonding strength. This result could be due to the<br />
fact that the surface, immediately after activation by plasma,<br />
represents a loose structure with weak or even broken<br />
interatomic bonds. Therefore, the atoms become highly<br />
mobile and actively react to the electric field. And vice versa,<br />
when the bonds are intensified as a result of an external action,<br />
the reaction of the atoms to the electric field becomes weaker.<br />
The latter explains the negative role of rinsing the specimen<br />
after plasma activation, as exposure, which saturates the<br />
interatomic bonds.<br />
The tensile test shows that the samples are mainly keep<br />
bonding up to 35 MPa when epoxy glue brakes. To<br />
qualitatively compare the bonding strength and strength of the<br />
bulk material, the samples were specially cleaved and a<br />
fracture surface was examined through a microscope. As is<br />
seen in Fig. 4, the fracture surface intersects the silicon and<br />
quartz crystals and does not include a bonding interface plane.<br />
The cleaved surface passes through the bonded sandwich<br />
structure as if it is a homogeneous bulk medium. That proves<br />
that the bonding is of a high strength, close enough to that of<br />
the initial bulk material.<br />
In addition to flat wafers, a structured silicon wafer was<br />
bonded to the crystal quartz. Bonding of structured wafer<br />
would be able to radically improve and simplify MEMS<br />
technology. The complicated structure could be prepared on<br />
the open surface by using conventional technology and then<br />
150
11-13 May 2011, Aix-en-Provence, France<br />
strength between the structured and the initial plates with flat<br />
<br />
surfaces<br />
<br />
V. CONCLUSIONS<br />
In this work, a strong low-temperature plasma assisted<br />
bonding of crystalline quartz/silicon wafers has been<br />
developed. High bonding strength, which is close to the<br />
mechanical strength of the starting materials was achieved<br />
through optimization of plasma activation and minimize the<br />
residual stresses. It is found that the electric field applied<br />
during annealing also contributes to the bonding strength. A<br />
possible mechanism of influence of the electric field on the<br />
bonding process was discussed. Similar strength has been<br />
achieved for the bonding of the crystal quartz wafer and<br />
structured silicon with micro cavities. The results demonstrate<br />
the unique potentials of the bonding technique in fabrication<br />
of MEMS devices based on the crystal quartz.<br />
Fig. 4. Fracture surface intersects the bonding plane<br />
be bonded to the quartz wafer. In addition, further pre<br />
processing of the bonded quartz could be performed on the<br />
open surface too. Therefore, the complicated multilayer<br />
structure could be fabricated in the framework of simple<br />
combinations of conventional surface technology and the<br />
bonding technique.<br />
The structure with cavities of 1x1mm in size and 0.1mm in<br />
depth was prepared with standard technology of<br />
photolithography and wet etching on the surface of the silicon<br />
wafer. The structured wafer has been subjected to the same<br />
processing as the initial silicon plate with a flat surface. The<br />
tensile test shows no significant differences in bonding<br />
REFERENCES<br />
[1] H. Takagi,R. R. Maeda, N. Hosoda, and T. Suga. “Room-temperature<br />
bonding of lithium niobat and silicon wafers by argon-beam surface<br />
activation”, Appl. Phys. Lett., Vol. 74, 2387-2389 (1999)<br />
[2] S. Bengtsson, and P. Amirefeiz. “Room temperature wafer bonding of<br />
silicon, oxidized silicon, and crystalline quartz,” Journal of Electronic<br />
Materials, vol. 29, No. 7, 2000, pp. 909-915<br />
[3] T. Suni, K. Henttinen, I. Suni, and J. Makkinen. “Effect of plasma<br />
activation on hydrophilic bonding of Si and SiO 2,” Journal of<br />
Electrochemical Society, Vol. 149, No 6, 2002, G348-G351.<br />
[4] G. G. Stoney. “The tension of metallic films depositd by electrolisis,”<br />
Proc. R. Soc.,London, Ser. A, 82, pp.172-175.<br />
[5] S. Timoshenko, “Analysis of bi-metal thermostate”, J. Optical Soc.<br />
America and Review Scientific Instruments, 11, 233 (1925)<br />
[6] A. V. Dobrynin. “On the aplication of Stoney Formula for calculating<br />
stresses in thick filmes and coatings,” Pis’ma Zn. Teck. Phiz. 23, 32-36<br />
(1997)<br />
[7] Y. Zimin, T. Ueda, “Bonding of silicon and crystal quartz wafers for<br />
MEMS application,” unpublished.<br />
[8] A. Weinert, P. Amrfeiz, and S. Bengtsson. “Plasma assisted room<br />
temperature bonding for MST,” Sensors and Actuators, VOl. A 92,<br />
2001, pp. 214-222.<br />
151
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
A Novel Self-Powered Method for Pipe Flow<br />
Measuring<br />
1 SONG HAO WANG, 1 RONALD GARCIA, 2 PEI HUA CHANG<br />
1 Department of Mechanical Engineering, Kun Shan University<br />
2 Department of Electronics Engineering, Kun Shan University<br />
949 Da Wan Road, Young Kang City, Tainan, 710 Taiwan<br />
Abstract: This paper presents a novel “Self-Powered<br />
Pipe Flow Metering System (SPPFM)” with the functions<br />
of measurement, analysis and display incorporated with<br />
self-powering. The features of vortex flow metering and<br />
pipe flow generator are adopted simultaneously to<br />
develop this integrated system.<br />
In this system, the pipe flow drives the rotor of a generator to<br />
power the measuring unit itself. Thanks to the system, only<br />
small part of the flow power is converted into electricity to<br />
obsolete the replacement of batteries or to eliminate the lengthy<br />
electric supply wires all together. The reading of SPPFM was<br />
adjusted with an accredited commercial flow meter for<br />
the purpose of background calibration.<br />
Major parameters including power needed and power<br />
generated under different flow rate are investigated. Based on<br />
the results of the experiments, feasibilities of the system are<br />
discussed. A commercial flow meter was used to calibrate this<br />
SPPFM system and the results of such a calibration are<br />
presented in the paper. A technique to achieve both goals using<br />
single generator hardware is explained. It was seen from the<br />
experiment that the SPPFM system would produce more electric<br />
power than the basic needs for an embedded digital measuring<br />
unit, leaving rooms for growing features such as wireless<br />
communication, graphic display and data storage.<br />
Keywords: Pipe flow, Metering, Generator, Self Power<br />
I. INTRODUCTION<br />
Flow metering devices are one of the most important<br />
apparatus to measure/control fluid flows in pipelines,<br />
including industries such as chemical/petroleum plants, as<br />
well as residential/municipal facilities.<br />
For example, water loss is an extremely important issue<br />
for human beings. The control of water losses has been an<br />
activity associated with water distribution as early as the<br />
earliest systems were built [1]. Since Roman times many<br />
advances have been made but even in the newest distribution<br />
system, leakage occurs and today leakage engineers require a<br />
variety of equipment and techniques to measure, control and<br />
reduce leakage on water supply networks[2].<br />
Water leakage can be divided into two categories,<br />
background leakage and burst leakage (Fig. 1). However,<br />
even today, water leakage can still be as high as 25% in<br />
developed countries and 50% in developing countries.<br />
The concept of “Water footprint” has been introduced<br />
recently to trace the clean water usage through human<br />
activities and development of products. To efficiently trace<br />
water leakage and usage, and water flow metering is a<br />
necessity.<br />
For deep sea oil drilling, the monitoring of the pipe flow is<br />
obviously a key issue (Fig. 2). During the Gulf of Mexico oil<br />
field disaster in 2010, the big pipe leakages had not been<br />
detected until after seven days of explosion, causing almost<br />
unrecoverable loss and damages to the United States<br />
economy and the world’s environment.<br />
Fig. 1 Water supply and leakages<br />
Fig. 2 Pipe lines in deep sea oil drilling<br />
152
In recent years, the pipe flow monitoring systems have<br />
advanced into digital era with the development of science and<br />
technology. However, at some sites it is troublesome to<br />
accomplish the task of maintain continue power supply to<br />
flow metering systems at either local or remote areas. The use<br />
of solar panels to power a regular flow meter is inefficient due<br />
to the lack of security and limited, if any, accessibility for<br />
maintenance of the equipment. Here is where SPPFM comes<br />
in handy [4], allowing the communication system to power<br />
itself with enough power to send signals using wire/wireless<br />
communication technology. Some examples of remote areas<br />
where SPPFM could be used are water pipes at mountains<br />
where the water is collected and extracted for domestic use or<br />
at water bodies where water is pumped through pipes for land<br />
irrigation or other usages.<br />
In urban areas, flow meters do not face the same issues as<br />
in remote areas. Electricity is available almost everywhere<br />
and communications networks are well spread out. However,<br />
as the cities grow so the needs for more flow meters. Two<br />
main sources power these flow meters, electricity from the<br />
power grid or batteries. By using SPPFM there will be no<br />
need to connect to the power grid anymore. Since SPPFM is<br />
powered by itself, it could be installed in pipes underneath the<br />
streets, and not only use them as flow meters but also as<br />
sensors to detect leakages in the pipelines. The usage of<br />
SPPFM goes beyond remote areas and cities. It can also be<br />
used at hazardous places where chemicals or liquids at<br />
extreme temperature are being handled.<br />
Oil pipelines at the desert and geothermal pipelines<br />
systems are among the numerous dangerous sites which can<br />
be beneficiated by the use of SPPFM (Fig. 3 and Fig. 4). High<br />
temperature water, going above 160ºC in geothermal plants,<br />
and crude oil are some of the substances human being can not<br />
be exposed to, still flow of these substances need to be<br />
controlled.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Fig. 4 Pipe lines in geothermal plant<br />
One of the most commonly used device in the industry are<br />
the vortex flow meters [3] which can work on extreme<br />
environment and with low or free maintenance. However,<br />
they still require an input power ranging from 13 to 32 VDC<br />
to operate. Powering these devices is troublesome in some of<br />
the cases. Therefore, SPPFM is an advantageous device in<br />
these types of environments. The application of the SPPFM<br />
does not end here. A more futuristic application would be that<br />
after miniaturization, this invention can even be integrated<br />
into embedded medical self-powered device inserted into a<br />
human body. This would enable physician to monitor/control<br />
the micro device and constantly analyzing and<br />
communicating the results.<br />
The Self-Powered Pipe Flow Metering System or SPPFM<br />
was created to minimize the maintenance needs and still have<br />
an accurate water flow measurement. This study was made<br />
practically and the measurements were compared with an<br />
accredited commercial flow meter, the Mini-wheel flow<br />
meter W-116, by TOKYO KEISO CO., LTD.<br />
. MECHANISM OF SPPFM WATER TURBINE<br />
Started with a small pipe line, this SPPFM uses Pelton<br />
Wheel design for its water turbine, the geometry and the<br />
quantity of impellers was analyzed and optimized with the<br />
help of CFD software and Rapid Prototyping technology [4,<br />
5].<br />
A prototype of SPPFM is shown in Figure 5. The system<br />
consists of an impeller, a data acquisition/analyzing unit and a<br />
power/pulse generator connected to the water/liquid pipe.<br />
For medium size pipes, the coils of the stator can be<br />
located outside the wall of the pipe (Fig. 6). For large pipes,<br />
the impeller and the generator can be inserted into the<br />
mainstream flow (Fig. 7). [15]<br />
Fig. 3 Pipe lines in remote area<br />
<br />
153
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Fig. 5 A prototype of SPPFM for small and big pipes<br />
Fig. 6 Water turbine for medium size pipes<br />
ELECTRONIC DESIGN FOR MEASURING UNIT<br />
In this study, the SPPFM measuring unit uses the ATMEL<br />
AVR Micro Controller Unit (MCU). This MCU uses from<br />
1.6V to 5.5V, and in active mode it only consumes 3.6mA.<br />
Therefore, the whole unit, including its display, only<br />
consumes about 0.1W to operate in active mode. On the<br />
contrary, most of the commercial system such as the<br />
Mini-Wheel W-116 uses 12V DC and 1.2A to operate, thus it<br />
is very difficultly to install them in areas where electric grid<br />
or other external source of energy is not present.<br />
As an example, at 10 L/min this generator produces 5.00<br />
VDC and 0.016A, supplying a total power of 0.08W. At this<br />
flow rate the generator is supplying for 80 percent of the<br />
power required. However, at 20 L/min, this generator<br />
achieves 11.20 VDC, supplying a total power of 0.41W,<br />
providing three times more than the total power required.<br />
While the power generated exceeds power required the extra<br />
energy is stored in ultra-capacitors which, fully charged, can<br />
provide the system with enough energy to measure the flow<br />
for five hours.<br />
A voltage regulator was needed to protect the circuit from<br />
burning out and keep a constant voltage of 3.3VDC. This<br />
arrangement allows the SPPFM to only use the power of the<br />
generator to operate safely without the need for batteries or<br />
any external power source.<br />
IV. SINGAL PROCESSING<br />
Fig. 7 Water turbine for large pipes<br />
EXPERIMENTAL SETUPS<br />
The experiment was carried on with a recirculation<br />
pumping test bench. It consists of a water tank, a pump and<br />
a loop of pipes (Fig.8).<br />
The SPPFM has a small generator which converts the<br />
mechanical force of flow into electrical energy. In this<br />
experiment, the inside diameter of the flow pipe is 18mm.<br />
The reading of flow rate is obtained from a commercial<br />
flow meter, “Mini-Wheel W-116” from TOKYO KEISO<br />
Co. LTD.<br />
The SPPFM system measures the frequency generated by a<br />
separated set of coil attached to the stator of the generator.<br />
The measuring system uses the analog comparator feature of<br />
the Micro Controlling Unit (MCU) to get an accurate<br />
frequency measurement.<br />
The analog comparator on AVR microcontrollers is a<br />
hardware feature that allows the comparison of two voltages,<br />
when the voltage of the positive analog comparator is higher<br />
that then voltage of the negative one a high pulse is<br />
generated. This feature was used to create a frequency out of<br />
the sine waveform of the alternating current received as<br />
signal from the generator.<br />
In order to process the signal received from the generator,<br />
hardware as well as software filtering were configured. The<br />
following figure shows the hardware filter used to control<br />
unwanted noise. However, if still noisy signals pass through,<br />
the MCU software detects and omits them by comparing the<br />
time past between each pulse.<br />
Fig.8 Experimental setup<br />
Figure 9. Filter for AC signals<br />
154
As the speed of the rotor increases, the frequency of the<br />
second set of coil increases as well. This second set of coil<br />
produced a voltage, high enough to enable the circuit to<br />
measure the rotor’s speed without an external amplifier or<br />
buffer, and without supplying any power to the signal..<br />
Therefore, the system is able to accurately measure the<br />
whole required range of frequency efficiently.<br />
This signal was compared with a hall sensor and it was<br />
ensured that the frequency of the AC signal generated by the<br />
second set of coil matches exactly the signal generated by<br />
the hall sensor, which is a commonly used method to<br />
measure the rotational speed of commercial flow meters.<br />
The circuit that measures this frequency was simulated<br />
using PC simulation software from Proteus Company. ISIS<br />
software was used to simulate the circuit as shown in figure<br />
10. The upper part of the figure is a simulation of the<br />
seven-segment display that SPPFM measuring system uses<br />
and the bottom part is a Virtual Signal Generator (VSG).<br />
This allows the unit to simulate the frequency generated by<br />
SPPFM, and so giving evidence of the accuracy of the<br />
measurement. In this study, the SPPFM is configured to<br />
update its value every five seconds. The unit of measurement<br />
shown on the display is in Hz. Therefore, this is 54Hz<br />
generated by the VSG and the same value is shown on the<br />
seven-segment display.<br />
The MCU that SPPFM uses has an accuracy of ± 3% of<br />
nominal frequency. This accuracy can only be achieved at<br />
voltage between 2.7 and 5.5 VDC and room temperatures of<br />
25ºC. For application under higher or lower temperatures,<br />
calibration could be executed to compensate and achieve<br />
higher precision.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Fig.11 Different flow rates measurement before calibration<br />
. EXPERIMENTAL RESULTS<br />
The electric power generated under different flow<br />
conditions is presented in Table 1 and plotted in Fig. 12 and<br />
Fig. 13.<br />
Table 1 Electricity from the pipe flow generator<br />
Fig. 10 ISIS software simulation for SPPFM measuring system (Hz)<br />
Contrary to Mini-Wheel W-116, which according to its<br />
data-sheet, has ± 5% of the nominal frequency. In order<br />
to calibrate SPPFM, data samples between SPPFM and<br />
W-116 were taken to obtain the relationship between Hz<br />
and the flow rate (L/min). The output of the SPPFM<br />
shows a linear relationship between the speed of the flow<br />
and the frequency of the signal. Similar conclusion was<br />
drawn between the speed of the flow and the flow rate, for<br />
Mini-Wheel W-116. The relationships are shown in figure<br />
11. Such direct linear relationship between the two<br />
systems allowed the SPPFM to be calibrated to give a<br />
very close approximation to the Mini-Wheel W-116 unit<br />
of measurement.<br />
Fig. 12 Power and voltage vs. flow rate (18 mm pipe)<br />
Fig.13 Power and voltage vs. water velocity (18 mm pipe)<br />
155
Since the power required by the present SPPFM<br />
measuring unit is about 0.1Watt, the idea of “self-power” is<br />
proved to be feasible under most of the flow conditions,<br />
even for this 18mm diameter pipe flow condition.<br />
Pictures on figure 14 show the operation of the system<br />
after calibration. The left side hand picture is at flow rate<br />
about 33 L/min, while the right side picture is about 47L/min.<br />
Please take a note that this is the operation without battery or<br />
outside electric supply.<br />
.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
publication on Journal of Advance Meterials, 2011<br />
8. Aleksandr Nagorny, Ph.D, “High Speed Permanent Magnet<br />
Synchronous Motor/Generator Design for Flywheel Applications”,<br />
Report for NASA Glenn Research Center.<br />
9. James W. Nilson, Susan a. Riedel, “Electric Circuit”, ISBN<br />
978-986-412-554-8, page 69<br />
10. http://www.national.com/ds/LP/LP2987.pdf<br />
11. http://www.atmel.com/dyn/resources/prod_documents/2486s.pdf<br />
12. http://www.tokyokeiso.co.jp/english/products/download/tg/W-100_TG-<br />
ES821E.pdf<br />
13. http://www.racinefed.com/RWL.pdf<br />
14. G.L. Pong, “Fluid Mechanics in Civil Engineering”, New<br />
Wun Ching Developmental <strong>Publishing</strong>,<br />
15. Song-Hao Wang, Ronald José Doblado Perez, Ronald<br />
García, and Jiacheng Chen, “Development of Pipe Flow<br />
Generators”, International Conference on Chemical<br />
Engineering and Advanced Materials (CEAM 2011)<br />
Fig. 14 Left hand side W-116 in L/Min vs Right hand side SPPFM measuring<br />
system in L/Min, calibrated<br />
. CONCLUTION AND OUTLOOKS<br />
The feasibility of the SPPFM concept has been proved<br />
through the study, achieving self powering and measurement<br />
accuracy. Under most flow conditions, only small amount of<br />
energy in the flowing fluid need to be extracted and<br />
transformed into electricity, to power the measuring unit. In<br />
this study, the electronic unit requires 0.1W to operate while<br />
the system reaches self-power at flow rate of 15L/min.<br />
Although this is a stand-alone system, there are<br />
alternatives for applications. It could also be a combination of<br />
the pipe flow generator with other existing metering<br />
methodologies such as ultrasonic or pressure based metering.<br />
In addition to flow rate, other characteristics of the flow<br />
such as temperature, pressure, leakage through acoustical<br />
signal processing, and PH value could be measured with<br />
self-power.<br />
To optimize the system, a Permanent Magnet Generator<br />
(PMG) for lower flow speed is under development. Moreover,<br />
the work of reducing power consumption of electronic unit is<br />
also underway.<br />
REFERENCES<br />
1. Richard Pilcher, “Leak Detection Practices & Techniques - A Practical<br />
Approach”, Water21,2003<br />
2. John Morrison, “Managing Leakage by District Metered Areas”,<br />
Water21, 2003<br />
3. http://www.atmel.com/ad/picoPower/<br />
4. Songhao Wang, Ronald Garcia, Xinyin He, Jiacheng Chen,”<br />
Development of a Self-Powered Pipe Flow Metering System”,15 th Flow<br />
Measurement Conference (FLOMEKO), 2010 Taipei, Taiwan<br />
5. S. Wang, C. F. H. Porres, M. Zuo, W. Xiao, “Study of Impeller Design<br />
for Pipe Flow Generator with CFD and RP”, Conference Proceedings of<br />
American Institute of Physics, AIP, (http://www.aip.org/), ISSN<br />
0094-243X.<br />
6. S. Wang, X. He, J. Ke and Z. Yang, “Optimization of A Pipe Flow<br />
Generator”, E12-005, Proceedings of 26 th CSME, 11/2009, Taiwan<br />
7. Songhao Wang, Ronald José Doblado Perez, Ronald García, and<br />
Jiacheng Chen, “Development of Pipe Flow Generators”, Accepted for<br />
156
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
A Microfluidic Chip with Single-particle-based Arrays<br />
Using Electroosmotic Flow<br />
Chun-Ping Jen and Ju-Hsiu Hsiao<br />
Department of Mechanical Engineering,<br />
National Chung Cheng University,<br />
Chia Yi, Taiwan, R.O.C.<br />
Abstract- Microfabrication technologies achieving precise<br />
manipulation of biological cells provide the potential for<br />
individual characterization, detection and assay to cells at the<br />
single-cell level. The main purpose of the present study was to<br />
develop a microfluidic chip with microwells for<br />
single-particle-based positioning by using electroosmotic flow.<br />
Therefore, the process could not only be reliable, but also simple<br />
without a syringe pump. A biocompatible material of<br />
polydimethylsiloxane (PDMS) was adopted as a structure in the<br />
microfluidic chip for single-particle-based array. The sample of 6<br />
μL with latex particles (17 μm in diameter) was suspended in the<br />
sucrose medium with a concentration of 10 6 particles/mL and<br />
dropped into the microchannel for micropatterning. The DC<br />
(direct current) voltages for electroosmotic flow were set as 10, 15<br />
and 20 volts, respectively. The velocity of electroosmotic flow<br />
increased with the applied voltages. The occupancy of particles<br />
decreased with voltages applied for both the microfluidic chips<br />
containing 20 or 30-μm microwells, which implied that the higher<br />
velocity of electroosmotic flow caused lower particulate<br />
occupancy. Furthermore, there was only one single particle<br />
within the individual microwell in most of occupied microwells<br />
with 20 μm in diameter, which was much higher than that for the<br />
30-μm-diameter microwells. Micropatterned latex particles in<br />
microwells were successfully achieved in this preliminary study.<br />
The microfluidic chips with microwells with different diameters<br />
were fabricated herein, which was suitable for measurements at a<br />
single-cell level.<br />
Keywords: microarray, single-particle, electroosmotic flow.<br />
I. INTRODUCTION<br />
Microfabrication technologies achieving precise<br />
manipulation of biological cells or microparticles provide the<br />
potential for individual characterization, detection and assay<br />
to cells at the single-cell level. It also has stimulated research<br />
to understand the fundamental cell biology and<br />
pharmaceutical analysis by exposure of cells to drugs and<br />
environmental perturbations [1]. Numerous methods, such<br />
as microcontact printing, microfluidic patterning and<br />
photolithography have been employed to create<br />
micropatterned surfaces containing adhesive and<br />
non-adhesive regions for cells, as reviewed previously [2-5].<br />
There approaches are limited to adherent cells and additional<br />
surface chemistry procedures are often required. Alternative<br />
methods including dielectrophoresis [6], optical tweezers [7]<br />
and selective dewetting [8] are adopted for trapping single<br />
cells and do not require that the cells are adherent. However,<br />
these methods are not suitable for high-throughput<br />
applications. The approach that cells are confined inside<br />
microwells passively becomes attractive because of its<br />
simplicity and easy-handling [9]. However, the injection by<br />
a syringe pump is still required for introducing the particles<br />
into the closed microchannel in this approach. Otherwise, a<br />
suspension of cells is pipetted onto the surface of the chip<br />
with opened microwells immersed in the medium in a culture<br />
dish [10], which required manually handling and was not<br />
reliable. The main purpose of the present study is to develop<br />
a microfluidic chip with microwells for single-particle-based<br />
positioning by using electroosmotic flow. Therefore, the<br />
process could not only be reliable, but also simple without a<br />
syringe pump.<br />
II. EXPERIMENTAL SECTION<br />
A biocompatible material of PDMS was adopted as a<br />
structure in the microfluidic chip for single-cell-based arrays,<br />
as illustrated in Fig. 1. The main channel formed on the top<br />
PDMS was 15 mm wide, 160 μm in height and 26 mm long.<br />
The main channel is divided into four microchannels with<br />
800 μm wide and 10 mm long at the center region. Each<br />
microchannel contains six 10×10 microwells with 20 μm or<br />
30 μm in diameter and 20 μm in depth on the substrate. The<br />
mold masters were fabricated by spinning SU-8 (SU-8 50,<br />
MicroChem Corp., Newton, MA, USA) on the silicon wafer<br />
to define the microwells and microchannel, respectively.<br />
The mold master of microfluidic channels (around 160 μm in<br />
height) were fabricated by spinning SU-8 at 500 rpm for 20<br />
seconds and then at 800 rpm for 35 seconds on the silicon<br />
wafer. The resist was soft baked on a hotplate at 65 °C for 10<br />
minutes and then at 95 °C for 30 minutes. The resist was<br />
then allowed to cool to room temperature. The SU-8 was<br />
exposed to ultraviolet (UV) radiation at a dose of 200 mJ/cm 2 .<br />
The post-exposure baking was done at 65 °C for 3 minutes<br />
and 95 °C for 10 minutes. The exposed samples were<br />
developed with the SU-8 developer for 5 minutes. Moreover,<br />
the mold master of microwells (around 20 μm in height)<br />
were fabricated by spinning SU-8 at 500 rpm for 20 seconds<br />
and then at a higher spin speed of 4500 rpm for 35 seconds on<br />
the silicon wafer. The resist was developed with the SU-8<br />
developer for about 2 minutes after baked and exposed to UV<br />
radiation under the same conditions mentioned above. The<br />
PDMS prepolymer mixture (Sylgard-184 Silicone Elastomer<br />
Kit, Dow Corning, Midland, MI, USA) was poured and<br />
cured on the mold masters to replicate the patterned<br />
structures. After peering off the PDMS replica with the<br />
microchannel, the inlet and outlet ports were made by a<br />
157
puncher. The two PDMS replicas were bonded after<br />
treatment of the oxygen plasma in the O 2 plasma cleaner<br />
(model PDC-32G, Harrick Plasma Corp. Ithaca, NY, USA).<br />
The electric field generating the electroosmotic flow was<br />
applied by inserting electrodes close to the ports<br />
aforementioned and the distance between these two<br />
electrodes was about 16 mm. The sample of 6 μL with latex<br />
particles (17 μm in diameter) was suspended in the sucrose<br />
medium with 8.62 wt% and 10 -4 M of KCl (σ=6.5×10 -3 S/m)<br />
and dropped into the microchannel for micropatterning. The<br />
concentration of particles was 10 6 particles/mL. The DC<br />
(direct current) voltages for electroosmotic flow were set as<br />
10, 15 and 20 volts, respectively. The detailed procedures<br />
were illustrated in Fig. 2.<br />
III. RESULTS AND DISCUSSION<br />
The images of micropatterned latex particles in 30 and<br />
20-μm microwells under different applied voltages were<br />
revealed in Fig. 3. The velocities of electroosmotic flow for<br />
10, 15 and 20 volts were measured as 3.0, 4.5 and 5.9 μm/s,<br />
respectively. The results of micropatterned in Fig. 3<br />
qualitatively showed that the occupancy of particles<br />
decreased with the voltage applied due to the increase of<br />
velocity. The statistical data for particle occupancy in the 20<br />
μm- and 30-μm-diameter microwells for different applied<br />
voltages were plotted in Fig. 4. The experimental data were<br />
based on manual counts of particles in three arrays of 10×10<br />
microwells by an inverted microscope. Each experimental data<br />
point represents the average value, and the error bar depicts<br />
the standard error from the mean. The occupancy of particles<br />
decreased with voltages applied for both the microfluidic<br />
chips containing 20 or 30-μm microwells, which implied that<br />
the higher velocity of electroosmotic flow caused lower<br />
particulate occupancy. For the case of applying 10 volts, the<br />
occupancy of particles on the microchip with 30 μm<br />
microwells was up to 93.67 %, which was higher than that<br />
obtained on the chip with 20 μm microwells (approximately<br />
85.16 %). The data for the particulate occupancy in the<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
individual microwells were revealed in Fig. 5 to investigate<br />
the performance of the single-particle level. The results<br />
indicated apparently that there was only one single particle<br />
within the individual microwell in approximate 97 % of<br />
occupied 20-μm microwells, which was much higher than<br />
that for the 30-μm-diameter microwells (62.48 %).<br />
IV. CONCLUSIONS<br />
The method of micropatterning latex particles in<br />
microwells at single-particle level was successfully achieved<br />
in this preliminary study. The microfluidic chips with<br />
microwells were fabricated herein, which was suitable for<br />
measurements at a single-cell level. However, the<br />
experimental demonstration of micropatterning biological<br />
cells is required in future work. Microchips with microwells<br />
proposed herein could be used for cellular micropatterning.<br />
The technique has the potential to realize single cell analysis<br />
and to acquire a population of data based on high-throughput<br />
and parallel processing.<br />
ACKNOWLEDGMENT<br />
The authors would like to thank the National Science<br />
Council of the Republic of China for its financial support<br />
under contract No. NSC-99-2923-E-194-001-MY3.<br />
REFERENCES<br />
[1] K. Yoshimoto, M. Ichinoa and Y. Nagasaki, Lab Chip, 9, 1286<br />
(2009).<br />
[2] A. Folch and M. Toner, Annu. Rev. Biomed. Eng., 2, 227 (2000).<br />
[3] T. H. Park, M. L. Shuler, Biotechnol. Prog., 19, 243 (2003).<br />
[4] D. Falconnet, G. Csucs, H. M. Grandin, M. Textor, Biomaterials, 27,<br />
3044 (2006).<br />
[5] J. Y. Lim, H. J. Donahue, Tissue Eng., 13, 1879 (2007).<br />
[6] J. Voldman, M. L. Gray, M. Toner, M. A. Schmidt, Anal. Chem. 74,<br />
3984 (2002).<br />
[7] A. Ashkin, Proc. Natl. Acad. Sci. 94, 4853 (1997).<br />
[8] N. Klauke, G. L. Smith, J. M. Cooper, Biophys. J., 85, 1766 (2003).<br />
[9] J. Y. Park, M, Morgan, A. N. Sachs, J. Samorezov, R. Teller, Y.<br />
Shen, K. J. Pienta, S. Takayama, Microfluid. Nanofluid., 8, 236<br />
(2010)<br />
[10] J. R. Rettig and A. Folch, Anal. Chem., 77, 5628 (2005).<br />
158
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Particles occupance(%)<br />
100<br />
90<br />
80<br />
70<br />
60<br />
50<br />
40<br />
30<br />
The sample of latex particles: 6 μL; 10 6 particles/mL<br />
20 μm microwells<br />
30 μm microwells<br />
Figure 1: Schematic diagram of the microfluidic chip for single-particle-based<br />
microarray.<br />
20<br />
10<br />
5 10 15 20 25<br />
Applied voltage (volts)<br />
Figure 4: Particle occupancy of latex particles in the microwells of 20-μm and<br />
30-μm diameter for different applied voltages.<br />
Particle occupance (%)<br />
100<br />
90<br />
80<br />
70<br />
60<br />
50<br />
20 μm microwells<br />
Applied voltage = 10 V<br />
Applied voltage = 15 V<br />
Applied voltage = 20 V<br />
Applied voltage = 10 V<br />
Applied voltage = 15 V<br />
Applied voltage = 20 V<br />
30 μm microwells<br />
40<br />
Figure 2: Experimental procedures for particle positioning.<br />
30<br />
20<br />
10<br />
0<br />
0 1 2 3<br />
Number of particles inside a microwell<br />
Figure 5: Particle occupancy of latex particles within the individual microwell<br />
of 20-μm and 30-μm diameter for different applied voltages.<br />
(a)<br />
(b)<br />
Figure 3: The images of micropatterned latex particles in the microwells of (a)<br />
30 and (b) 20 μm in diameter, respectively.<br />
159
11-13 May 2011, Aix-en-Provence, France<br />
<br />
A novel full range vacuum pressure sensing<br />
technique using free damping decay of micro-paddle<br />
cantilever beam deflected by electrostatic force<br />
Guan-Lan Chen, Chi-Jia Tong, Ya-Chi Cheng, Yu-Ting Wang, Ming-Tzer Lin *<br />
Graduate Institute of Precision Engineering,<br />
National Chung Hsing University, Taichung, Taiwan 402<br />
Abstract- We report here a novel full range vacuum pressure<br />
sensing technique. The technique, using the free damping decay<br />
of micro-cantilever beam, gives us a full range pressure sensing<br />
capacity ranging from 0.2 to 1 x 10 -8 torr. The method<br />
demonstrated in this study allows researchers and engineers to<br />
observe the vacuum pressure according to the free decay of the<br />
deflected MEMS structure responding to electrostatic loads. The<br />
sensing results show the free decay of the deflected beam is linear<br />
proportion to the vacuum pressure. This can be performed at<br />
various vacuum pressures and the measurements can be<br />
achieved at frequency rates of up to 500 Hz.<br />
I. INTRODUCTION<br />
With rapid development of MEMS technology and industry<br />
has increased not only the response speed of the devices but<br />
also the actuation or driving method in products. Moreover, it<br />
enhanced the miniaturization of products. For the VLSI and<br />
MEMS process applications, there are many equipments using<br />
vacuum. However, the vacuum working environment are not<br />
the same. For example, it is usually using 10 -4 torr in general<br />
heat treatment while plasma etching system required vacuum<br />
pressure lower than 1 x 10 -6 torr due to its anisotropic etching<br />
by ion collision. Moreover, Molecular Beam Epitaxy (MBE)<br />
processes need to be performed in the ultra-high vacuum<br />
environment thus to avoid epitaxy growth chamber and<br />
substrate pretreatment contacting with atmosphere during the<br />
manufacturing process [1]. As a result, it required various<br />
vacuum environments for MEMS industry and an accurate<br />
vacuum pressure measurement tools are essential.<br />
In general, vacuum gauge can be characterized from direct<br />
pressure measurement and indirect pressure measurement<br />
according to its sensing response with gas. At present, there<br />
are three types of vacuum gauge that are commonly used:<br />
capacitive diaphragm gauge (1 x 10 3 torr - 1 x 10 -1 torr),<br />
thermal conductivity gauge (1 x 10 3 torr - 1 x 10 -3 torr), and<br />
ionization gauge (1 x 10 -2 torr - 1 x 10 -10 torr). However, they<br />
are more or less with some drawbacks. For example,<br />
temperature effects on zero stability and the resolution of<br />
capacitive diaphragm gauge are difficult to avoid. In addition,<br />
geometric design of sensor elements in thermal conductivity<br />
gauge has to meet the flow of gas molecules in very sensitive<br />
state. Moreover, ionization gauge is only limited for high<br />
vacuum measurement. Furthermore, the vacuum gauges<br />
described above through these three methods have limited<br />
range for pressure measurement and there are no sensors that<br />
are capable to measure pressure with full range from<br />
atmosphere to high vacuum.<br />
Previously, Beams and Young [2] designed a method used a<br />
spinning metal rotor in vacuum to determine the<br />
environmental pressures. This vacuum gauge based on the air<br />
viscosity decelerate rotor speed is the first wide-range vacuum<br />
gauge. The measurement principle is to accelerate the rotor<br />
served as a vacuum gauge to a certain speed using the<br />
magnetic field. After the rotor reached a proper speed then the<br />
magnetic field is turned off. At the same time, the speed of the<br />
rotor will be decelerated with respect to the viscosity of the<br />
surrounding gas. As a result, the environmental pressures can<br />
be calculated according to the changes of the rotating speed.<br />
Despite its application for the wide-range vacuum pressure<br />
measurement, this vacuum gauge has experienced some<br />
disadvantages as well, example such as the high production<br />
cost and the long measurement time. In addition, the<br />
dimensions of the gauge are also difficult to be reduced.<br />
In this study, we design and develop an alternative full<br />
range vacuum gauge using new concept to integrate with the<br />
MENS technology. It is performed to use an elastic body with<br />
its dynamic response under different pressures to read accurate<br />
vacuum pressures of the environment with wide range. The<br />
system consists of a micro-scale metal film attached to the<br />
cantilever beam. It is located inside a vacuum chamber and<br />
used to read the vacuum pressure according to its dynamic<br />
damping response.<br />
II. PADDLE DESIGN & FABRICATION<br />
It is the fact that for a vibrated spring-mass system the decay<br />
rate of amplitude is proportional to the surrounding pressure of<br />
gaseous. Therefore, it can be extended to use as the concept for<br />
pressure sensing device. Here, a paddle liked sensing device<br />
was designed and developed with a structure including a proof<br />
mass, a tapered beam structure and a frame used to support the<br />
spring-mass system. The function of tapered beam structure is<br />
similar to the role of spring in spring-mass system. In addition,<br />
the cyclic motion can be provided on the proof mass because<br />
the compliance force on such structure is driven by<br />
electrostatic force due to the tapered beam bent down or up.<br />
The design of using the tapered beam was to avoid the plane<br />
stress concentrate on the root of the beam when the sensing<br />
device is bent. The sensing device was fabricated using 4<br />
160
inches silicon wafers which have thickness of 250μm. Special<br />
high conductivity silicon wafers were chosen to make these<br />
sensing devices because they can be directly used to couple the<br />
capacitance in the system. Before fabrication started, wafers<br />
need to go through the standard RCA cleaning procedures in<br />
order to remove the organism and particles on its surface.<br />
Silicon nitride layers were then growth on both sides using low<br />
pressure chemical vapor deposition (LPCVD). This nitride<br />
layer is used as an etching barrier due to its high etching<br />
selectivity compare with silicon. Next, silicon nitride was<br />
patterned by photoresist and reactive ionic etching (RIE).<br />
After the silicon nitride was patterned, the entire wafer will be<br />
immersed in KOH solution at 85℃. Non-protect regime of<br />
silicon will then be removed on both sides until the full<br />
thickness of silicon had been etched through. Finally, the<br />
silicon nitride layers were taken away using hydro fluoric acid.<br />
The complete fabrication sequence is shown in Fig. 1 and the<br />
photograph of the final sensing structure through complete<br />
fabrication flow is shown in Fig. 2.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
comparison. Photograph of vacuum system shows in Fig. 3.<br />
Fig. 3. A photograph of system<br />
The sensing system we proposed was then used to obtain the<br />
dynamic response from the capacitance measurement. There<br />
are two electrodes mounted around this sensing device, the<br />
sensing electrode placed on top of the sensing device and the<br />
driving electrode mounted underneath of the sensing device.<br />
The gap between the sensing electrode and PCB will couple a<br />
capacitance in a fixed value unless for the sensing device is<br />
moved or under vibration. The amplitude of the displacement<br />
current will change from the coupling capacitance if there are<br />
changes in the gap between the device and the sensing<br />
electrode. The amplitude of the displacement current will be<br />
proportional to the capacitance, and also be inverse<br />
proportional to the gap between electrode and device. The<br />
schematic of measurement shows in Fig.4.<br />
Fig. 1. Complete stepwise processing sequence for sample<br />
process<br />
Fig. 4. A schematic of measurement<br />
Fig. 2. A photograph of sample after process<br />
III. SYSTEM SET UP<br />
The system can be characterized in two major parts; the first<br />
is the high vacuum system which used to provide different<br />
vacuum pressure during measurement. Prepared vacuum<br />
pressure was controlled by passing nitrogen gas flow into the<br />
chamber operated using a high accuracy mass flow controller<br />
(MFC). Simultaneously, the pressure inside the chamber was<br />
measured using preexisted linearly wide range vacuum gauge<br />
(WRG) which consists with a low vacuum gauge and a high<br />
vacuum gauge thus to give the precise vacuum pressure for the<br />
As shown in Fig. 4 when putting a fixed DC voltage on the<br />
bottom electrode, it will provide an electrostatic force and pull<br />
the sensing device bending down. If we suddenly removed the<br />
drive voltage on the bottom electrode, this sensing device will<br />
bend up and then bend down due to its compliance force.<br />
Finally, the sensing device will damp out to its balance<br />
position and the vibration time constant of the device can be<br />
used to determine the environmental pressure surround it.<br />
Here, the bottom electrode is used to provide the<br />
electrostatic force and induce the sensing device bent or<br />
vibrate. The equivalent impedance of the capacitance between<br />
them is a reciprocal of vibration frequency. If the sensing<br />
device oscillate at the lower frequency such as several Hz to<br />
hundreds Hz, the capacitance between it and the sensing<br />
electrode would provide an equivalent impedance in millions<br />
ohms. On the other hand, the displacement current we<br />
161
measured may be reduced to several nano-ampere even<br />
pico-ampere. In general, it is difficult to measure and acquire<br />
such small amplitude of displacement current. Therefore, a<br />
100 kHz sine waveform signal has been designed to carry the<br />
lower frequency signal through the coupling capacitor.<br />
Although, the displacement current obtained from the<br />
capacitor has low frequency and high frequency components,<br />
the lock-in amplifier added here can extract the low frequency<br />
signal that are needed for our measurement from the mixed<br />
signal with a 100kHz reference signal. Capacitance of the<br />
coupling capacitor can be easily determined through the<br />
comparison of the other reference capacitor that has a known<br />
capacitance. The other capacitor also has a 100kHz bias signal,<br />
but has 180 degree phase off with another 100kHz sine wave.<br />
The schematic of the measurement tool is shown in Fig.5. Fig.<br />
5 (a) shows 3-D diagram and (b) shows the cross-section view<br />
of entire system with details in each part. This capacitance<br />
measurement was controlled using Labview software and a PC.<br />
The electrical circuit design of entire capacitance<br />
measurement is shown in Fig.6.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Fig. 6. Schematic of measurement signal circuit design<br />
During the experiment, a step voltage was provided on the<br />
bottom electrode. The sensing device started vibration after<br />
sudden turned off the deflection voltage. The dynamic response<br />
of the sensing device will be recorded continuously until the<br />
sensing beam gets back to its balance position. The amplitude<br />
of vibration will be reduced with time because of the air<br />
damping due to the gaseous atoms.<br />
IV. RESULTS AND DISCUSSION<br />
Fig. 7 shows the results of one particular experiment tested<br />
at the low vacuum (1.6 x 10 -2 torr). The figure shows the proof<br />
mass position was slight decreased from its maximum value at<br />
the beginning. The maximum value of each cycle can be fitted<br />
using an exponential decay function defined as (1). Decay rate<br />
of the entire damping behavior can be calculated using (2).<br />
)( ⋅= (1) , δ ( ) (2)<br />
0<br />
⋅− t<br />
eAta<br />
δ = ln / aa<br />
nn<br />
+ 1<br />
In Fig. 7, the free decay response of the beam vibrated at the<br />
low vacuum has a decay time constant in 3.05 second. Not<br />
only the vacuum pressure would affect the decay time constant<br />
in the dynamic behavior of the sensing beam but also the<br />
intrinsic property in its material. Silicon is one of the low loss<br />
materials which can be used to avoid the decay time constant<br />
influenced by the variation of intrinsic material properties.<br />
Fig. 5. Schematic view of measurement apparatus (a) 3D view<br />
(b) cross-section view<br />
Fig. 7. Sample free damping decay versus time at 10 -2 torr<br />
Here we carried out a series of experiments at the different<br />
vacuum pressures. Vacuum pressure can be adjusted form the<br />
lowest pressure to the nearly atmosphere. The decay time<br />
constant obtained here shows a significant difference between<br />
162
samples tested in different vacuum pressures. Test results<br />
plotted in Fig. 8, Fig. 9, and Fig. 10 show sample decay rate<br />
versus vacuum pressure on three different ranges of pressure<br />
such as 0 - 0.01torr, 0 - 0.1torr, 0 - 0.025torr, respectively. We<br />
observed the significant trend of decay rate versus different<br />
pressure in free damping experiments. Decay rate found here<br />
shows linearly proportional to the vacuum pressure. As a result,<br />
this sensing technique can be used to calibrate as a standard<br />
pressure sensor as long as the decay rate has been found using<br />
the free damping method.<br />
Fig. 8. Sample decay rate versus vacuum pressure (0 -<br />
0.01torr)<br />
Fig. 9. Sample decay rate versus vacuum pressure (0 - 0.1torr)<br />
Fig. 10. Sample decay rate versus vacuum pressure (0 -<br />
0.025torr)<br />
V. CONCLUSION<br />
A novel pressure sensing technologies has been proposed in<br />
this study. The sensing technology used to determine the<br />
vacuum pressure depending on the viscosity of gaseous. The<br />
results show that the air pressure has a significant effect on the<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
decay time constant of free vibration of the sensing device.<br />
Cantilever beam was designed in tapered shape which can<br />
effectively reduce the bending stress concentrate on the root.<br />
This design improved the volumetric problem in spin rotor<br />
gauge, at the same time it also gives us a wider range<br />
measurement in pressure. The results also show the linear<br />
correlation between the damping decay rates versus vacuum<br />
pressure. Thus it gives us a full range vacuum pressure sensing<br />
capacity ranging from 0.2 to 1 x 10 -8 torr linearly.<br />
REFERENCES<br />
[1] VACUUM TECHNOLOGY & APPLICATION<br />
National Science Council Precision Instrument Development Center<br />
[2] J. W. Beams, J. L. Young, and J. W. Moore, Journal of Applied Physics,<br />
17, 886 (1946)<br />
[3] C. J. Tong, Y. C. Cheng and M. T. Lin, Microsystem technologies,<br />
16(7), 1131 (2010).<br />
[4] C. J. Tong and M. T. Lin, Microsystem technologies, 15(8), 1207<br />
(2009).<br />
[5] B. C. S. Chou, Y. M. Chen, M. OuYang and J. S. Shie, Sensors and<br />
Actuators A, 53, 273 (1996).<br />
[6] J. S. Shie, B. C. S. Chou and Y. M. Chen, Journal of Vacuum Science<br />
and Technology A, 13 (1995).<br />
[7] Y. Wang and M. Esasshi, Sensors and Actuators A, 66, 213 (1998).<br />
[8] K. R. Williams and Richard S. Muller, International Electron Devices<br />
Meeting, 387 (1992).<br />
[9] K. R. Williams and R. S. Muller, Transducers, 97, 1249 (1997).<br />
[10] D. H. Baker and R. A. Outlaw, in 45 th AVS international Symposium<br />
(1998).<br />
[11] D. C. Calting, Sensors and Actuators A, 64, 157 (1998).<br />
[12] R. T. Bayard and D. Alpert, Review of Scientific Instruments, 21, 571<br />
(1950).<br />
[13] J. K. Fremerey, Journal of Vacuum Science & Technology A, 3(3),<br />
1715 (1985).<br />
[14] J. W. Beams, J. L. Young, and J. W. Moore, Journal of Applied Physics,<br />
17, 886 (1946)<br />
[15] J. K. Fremerey, Journal of Vacuum & Science Technology, 9, 108<br />
(1972).<br />
[16] J. K. Fremerey, Review of Scientific Instruments, 44, 1396 (1973).<br />
[17] J. K. Fremerey and K. Boden, Journal of Physics E, 11,106 (1978).<br />
[18] J. W. Beams, D. M. Spitzer and J. P. Wade, Review of Scientific<br />
Instruments, 33, 151 (1962).<br />
[19] J. Harbour and R. O. Lord, Journal of Scientific Instruments, 42,105<br />
(1965).<br />
[20] G. Comsa, J. K. Fremerey and B. Lindenau, in Proceedings of the 7 th<br />
International Vacuum Congress, Vienna, Vol. I, 157 (1977)<br />
[21] G. Comsa, J. K. Fremerey, B. Lindenau, O. Messer and P. Rohl, Journal<br />
of Vacuum & Science Technology, 17,642 (1980).<br />
[22] G. Messer, in Proceedings of the 8th International Vacuum Congress,<br />
Cannes, Vol. II, 191 (1980).<br />
[23] G. Comsa, J. K. Fremerey and B. Lindenau, in Proceedings of the 8 th<br />
International Vacuum Congress, Cannes, Vol. II, 218 (1980).<br />
[24] G. Messer and L. Rubet, in Proceedings of the 8th International<br />
Vacuum Congress, Cannes, Vol. II, 259 (1980).<br />
[25] Cong Shu-Ren, Wan Yon-liang and Lu Jia-huo, Vac. Sci. Technol.<br />
(China) 2, 64 (1982).<br />
[26] G. Reich, Journal of Vacuum & Science Technology, 20, 1148 (1982).<br />
[27] K. E. McCulloh, Journal of Vacuum & Science Technology A, 1,168<br />
(1983).<br />
[28] P. J. Van Ekeren, M. H. O. Jacobs, J. C. A. Offringa, and C. O. De Kruif,<br />
J. Chern. Thermodynamics IS, 409 (1983).<br />
[29] Fujio Tamura, United States Patent 5,033,306 (1991).<br />
[30] J. H. Martin and W. P. Kelley, United States Patent 5,528,939 (1996).<br />
[31] J. H. Martin, United States Patent 5,939,635 (1999).<br />
[32] R. C. Gutierrez, C. B. Stell, T. K. Tang, V. Vorperian, J. Wilcox, K.<br />
Shcheglov and W. J. Kaiser, United States Patent 6,085,594 (2000)<br />
[33] R. Correale, C. Maccarrone and M. Busso, United States Patent US<br />
7,059,192 B2 (2006)<br />
[34] R. Correale, United States Patent US 7,334,481 B2 (2008).<br />
163
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
Design and Development of Vibrational<br />
Mechanoelectrical MEMS Transducer for<br />
Micropower Generation<br />
Rolanas Dauksevicius 1 , Genadijus Kulvietis 1 , Vytautas Ostasevicius 2 , Ieva Milasauskaite 2<br />
1 Department of Information Technologies, Vilnius Gediminas Technical University<br />
Sauletekio al. 11, LT-10223 Vilnius, Lithuania<br />
2 Institute for High–Tech Development, Kaunas University of Technology<br />
Studentu str. 65, LT-51369 Kaunas, Lithuania<br />
Abstract- The paper is devoted to design, numerical<br />
modeling and analysis of vibration-driven mechanoelectrical<br />
MEMS transducer based on piezoelectric cantilever-type<br />
microstructure, which function is to act as a micropower<br />
generator in wireless sensor networks. This study also deals<br />
with fabrication and experimental investigation of<br />
piezoelectric PVDF thin films intended for energy harvesting<br />
applications. The first part of the paper presents finite element<br />
model of the transducer, which is a multiphysics one,<br />
combining mechanics, piezoelectricity and fluid-structure<br />
interaction in the form of squeeze-film damping governed by<br />
nonlinear compressible isothermal Reynolds equation.<br />
Subsequently the model is subjected to modal, harmonic and<br />
transient analyses in order to determine the effect of viscous<br />
air damping and geometrical parameters on device dynamical<br />
and electrical performance. The second part of the paper<br />
considers aspects of formation of PVDF thin films. The quality<br />
of the produced thin films and their material characteristics<br />
are evaluated by means of scanning electron and atomic force<br />
microscopy as well as using X-ray diffractometry and FT-IR<br />
spectrometry techniques. Performed experiments reveal that<br />
fabricated PVDF samples possess distinct crystalline phases,<br />
with alpha-phase being predominant.<br />
I. INTRODUCTION<br />
Constant progress in low-power electronics promotes<br />
rapid development of large variety of battery-operated<br />
portable, wearable, implantable and embedded devices<br />
including autonomous wireless sensors, which have huge<br />
potential in body area networks, condition monitoring and<br />
ambient intelligence applications. Despite the fact that<br />
energy density of batteries has increased by a factor of 3<br />
over the past 15 years [1], frequently their usage has a<br />
significant negative effect on device size and operational<br />
cost. For example, conducting their maintenance for a largescale<br />
sensor networks consisting of hundreds or thousands<br />
of sensor nodes may be extremely unpractical and<br />
uneconomical. In some cases batteries is not a feasible<br />
solution, e.g. in providing reliable long-term power to<br />
remote sensing systems that operate in harsh environments<br />
such as downholes in mining, oil/gas extraction as well as<br />
nuclear reactors, deep-sea or space applications. For this<br />
reason alternative approaches are the subjects of active<br />
research work worldwide. Several possibilities are<br />
considered including [1,2]: (a) energy storage systems with<br />
larger energy densities (e.g. miniaturized fuel cells),<br />
however, still significant work is required for the realization<br />
of commercial devices; (b) wireless powering solutions (as<br />
employed in RFID tags), however, their tailoring for more<br />
power intensive devices would require dedicated<br />
transmission infrastructures; (c) harvesting ambient energy<br />
by using vibration/motion or thermal energy, light or RF<br />
radiation, acoustic noise, etc. Energy harvesters can<br />
typically supply power in the range of 0.01 − 1 mW<br />
depending on the employed conversion principle.<br />
Meanwhile, the consumption of common wireless sensor<br />
nodes is between 1 and 20 μW, with values reaching up to<br />
100 μW for relatively complex nodes operating at high<br />
data-rates [1]. Kinetic energy harvesting is particularly<br />
attractive as structural vibrations are ubiquitous in the<br />
environment. For example, it is well suited for supplying<br />
energy to autonomous sensors in condition monitoring of<br />
industrial machines or civil structures. Piezoelectric and<br />
electromagnetic transduction mechanisms are considered to<br />
be the most promising for vibrational harvesters, while<br />
electrostatic devices are presently limited by their high<br />
impedance and output voltages, which reduce the amount of<br />
available current. Piezoelectric micropower generators<br />
(PMPGs) have the advantages of relatively simple geometry<br />
and fewer peripheral components. Moreover, it is not<br />
difficult to integrate microelectronic circuits on the same<br />
chip because the process for depositing both thin and thick<br />
piezoelectric films is a fairly mature technology [2,3].<br />
However, the majority of current micro-scaled PMPGs do<br />
not generate sufficient energy to directly power most<br />
electronics including MEMS-based devices. Significant<br />
research efforts are currently focused on two principal<br />
approaches for improving efficiency of these generators: (a)<br />
development of hybrid micropower supply units comprising<br />
both on-board storage and energy harvesting from<br />
environmental vibrations, optimization of energy generation<br />
164
l c<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
t c t p<br />
w c<br />
t m<br />
l<br />
w m<br />
(a)<br />
(b)<br />
h 0<br />
Air film<br />
Fig. 1. Schematic representation of the modeled PMPG, which consists of a<br />
double-layer cantilever structure (composed of supporting Si layer t c and<br />
piezoelectric layer t p atop) with proof mass at the free end. Air film of<br />
thickness h 0 is present between an imaginary fixed ground surface<br />
and the bottom boundary of the proof mass (drawn not to scale).<br />
TABLE I<br />
DESIGN PARAMETERS OF THE PMPG<br />
Description and symbol Value Unit<br />
Length of uniform cantilever l c 2500 μm<br />
Length of proof mass l m 1500 μm<br />
Width of cantilever w c 300 μm<br />
Width of proof mass w m 3000 μm<br />
Thickness of cantilever supporting layer t c 20 μm<br />
Thickness of piezoelectric layer t p 20 μm<br />
Thickness of proof mass t m 1000 μm<br />
Young’s modulus of supporting layer and proof mass (Si) E Si 200 GPa<br />
Density of supporting layer and proof mass (Si) ρ Si 2330 kg/m 3<br />
Poisson’s ratio of supporting layer and proof mass (Si) ν Si 0.33 -<br />
and accumulation subsystems; (b) improvement of charge<br />
density via material and structural enhancements including<br />
expansion of operating frequency range of the devices by<br />
means of tuning of their resonant frequency or widening<br />
their bandwidth [1-4].<br />
Research results reported in this paper deal with the<br />
design and material issues of the PMPGs. The second<br />
section presents finite element modeling and simulation<br />
results of the device with emphasis on dynamic analysis of<br />
viscous air damping (squeezed air-film) effects that may be<br />
encountered during operation of the micropower generator.<br />
The third section considers aspects of fabrication of PVDF<br />
thin films and provides results of their experimental<br />
characterization by using SEM, AFM, XRD and FT-IR<br />
analysis methods.<br />
II.<br />
FINITE ELEMENT MODELING AND SIMULATION<br />
A. Model Description<br />
The design of analyzed PMPG is based on bi-layer<br />
cantilever structure with proof mass at the free end (Fig. 1<br />
and Table I). The supporting cantilever layer and the proof<br />
mass are made from silicon, while PZT-5A is used for<br />
piezoelectric layer, which is positioned on the top of the<br />
supporting layer and is poled along the thickness direction<br />
resulting in transverse (“d 31 ”) operation mode. Such<br />
configuration is chosen since it enables the condition of low<br />
resonance frequency to be fulfilled. This is required in<br />
typical applications of vibrational PMPGs, i.e. powering<br />
of wireless sensors installed in industrial or civil structures.<br />
(c)<br />
(d)<br />
Fig. 2. The first four vibration modes of the analyzed PMPG: (a) the 1 st out-ofplane<br />
flexural mode (184 Hz), (b) the 1 st torsional mode (458 Hz), (c) the 2 nd<br />
torsional mode (1628 Hz), (d) the 2 nd out-of-plane flexural mode (1689 Hz).<br />
(a)<br />
(c)<br />
(d)<br />
Fig. 3. 3D contour plots illustrating distribution of air pressure forces in the<br />
gap for the corresponding structural mode shapes in Fig. 2.<br />
These environments commonly generate vibrations that are<br />
characterized by small acceleration (< 1g) and low<br />
frequencies (up to approximately 200 Hz).<br />
Finite element model of the PMPG was realized within<br />
COMSOL Multiphysics by employing the “Piezoelectric<br />
Application” mode. Piezoelectric layer has got electrodes on<br />
its bottom and top faces. Due to low thickness mechanical<br />
behavior of the electrodes may be neglected. Their electrical<br />
behavior is evaluated by imposing proper electrostatic<br />
boundary conditions: the bottom face is grounded, while the<br />
top one is set to “Floating potential” condition. For the rest<br />
of faces of the piezoelectric layer the condition of “Zero<br />
charge/Symmetry” is applied.<br />
When designing PMPGs with bulky proof mass at the end<br />
of the cantilever structure one should consider a possibility<br />
of manifestation of a specific case of viscous air damping<br />
phenomenon referred to as squeeze-film damping, which<br />
occurs when a structure of large lateral dimensions, that<br />
is in relatively close proximity to a fixed surface, vertically<br />
(b)<br />
165
Fig. 4. Amplitude-frequency characteristics of end point of the cantilever<br />
structure, obtained in the vicinity of the fundamental frequency<br />
of the PMPG in the presence of squeeze-film damping for a<br />
constant air-film thickness (h 0 = 50 μm) at different levels<br />
of ambient pressure p 0: 100 Pa, 1 kPa, 5 kPa, 10 kPa,<br />
25 kPa, 100 kPa (curves from top to bottom).<br />
Fig. 5. Amplitude-frequency characteristics of end point of the cantilever<br />
structure, obtained in the vicinity of the fundamental frequency of the<br />
PMPG in the presence of squeeze-film damping for a constant<br />
ambient pressure (p 0 = 100 kPa) at different air-film thickness h 0:<br />
50 μm, 75 μm, 100 μm, 150 μm (curves from bottom to top).<br />
The topmost curve (green) is a frequency<br />
response with damping excluded.<br />
moves towards this nearby rigid surface with a thin air-film<br />
in-between. Thus, if the bottom face (boundary) of the proof<br />
mass is located relatively close to some stationary ground<br />
surface, then during transverse motion of the mass its fairly<br />
small displacement in normal direction would compress (or<br />
pull back) a significant amount of air out of (or into) the<br />
narrow gap. However, the viscosity of the air will limit the<br />
flow rate along the gap, and thus the pressure will be<br />
increased inside the gap and act against the structure. The<br />
squeezed air-film between the mass and ground surface will<br />
likely to have a significant effect on PMPG dynamic<br />
behavior due to the induced counter-reactive pressure force<br />
that is exerted on the vibrating cantilever structure [5].<br />
Nonlinear compressible isothermal Reynolds equation is<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
usually used for modeling of squeeze-film damping<br />
occurring in micro-scale [6]:<br />
∂ ⎛ 3 ∂P<br />
⎞ ∂ ⎛ 3 ∂P<br />
⎞ ⎛ ∂P<br />
∂h<br />
⎞<br />
⎜ Ph ⎟ +<br />
⎜ Ph<br />
⎟ 12μ=<br />
eff ⎜h<br />
+ P ⎟ , (1)<br />
∂x<br />
⎝ ∂ ⎠ ∂yx<br />
⎝ ∂<br />
⎠ ⎝ ∂t<br />
∂t<br />
⎠<br />
μ<br />
eff =μ . (2)<br />
1.159<br />
⎛<br />
0PL<br />
⎞<br />
1+<br />
9.638<br />
⎜ a<br />
⎟<br />
⎝ hp 00 ⎠<br />
Here the total pressure in the gap P and the gap thickness h<br />
are functions of time and position (x, y). μ is the dynamic<br />
viscosity of the gas, μ eff is the effective viscosity coefficient,<br />
which is used to account for gas rarefaction effects (a model<br />
of T. Veijola [7] is used here; it is adopted by the COMSOL<br />
as one of the optional approaches), p 0 is the initial (ambient)<br />
pressure in the gap, L 0 is the mean free path of air particles<br />
at atmospheric pressure P a , and h 0 is the initial air-film<br />
thickness. For the P a = 101325 Pa, L 0 ≈ 65 nm. Total<br />
pressure in the gap is equal to P = p 0 + Δp, where Δp is an<br />
additional film pressure (variation) due to the squeezed airfilm<br />
effect.<br />
“Film Damping Application” mode, which uses (1) and<br />
(2), was added to the piezoelectrical model in order to<br />
simulate frequency and time responses of the PMPG taking<br />
into account the effect of squeeze-film damping (a<br />
linearized version of (1) is used for harmonic analysis).<br />
B. Numerical Analysis<br />
Numerical study of the developed PMPG finite element<br />
model commenced from the determination of the natural<br />
frequencies and the associated vibration mode shapes (Fig.<br />
2). Performed modal analysis indicates that the fundamental<br />
frequency of the PMPG is equal to 184 Hz. Fig. 2 illustrates<br />
the first four mode shapes: a) the 1 st out-of-plane flexural<br />
mode, b) the 1 st torsional mode, c) the 2 nd torsional mode, d)<br />
the 2 nd out-of-plane flexural mode. This analysis also<br />
provided results on distribution of air pressure forces in the<br />
gap when the structure is vibrating in its flexural and<br />
torsional resonant modes. Pressure mode shapes in Fig. 3<br />
reveal obvious coupling between structural displacements of<br />
the structure and pressure distribution in the gap. For<br />
example, in the 1 st torsional mode (Fig. 2(b)), the upward<br />
motion of left side of the proof mass corresponds to a<br />
concave profile in the respective region of pressure mode<br />
shape (Fig. 3(b)), which indicates the reduction of pressure<br />
in this part of the gap (i.e. decompression effect). And, in<br />
contrast, the downward motion of right side corresponds to<br />
a convex pressure profile – zone of increased pressure with<br />
respect to atmospheric (i.e. compression effect).<br />
The aim of the subsequent numerical experiments was to<br />
determine influence of viscous air damping on dynamical<br />
behavior of the PMPG and its generated voltage. The<br />
simulations were performed with zero structural damping.<br />
The model was subjected to sinusoidal kinematic excitation<br />
by applying vertical acceleration through body load that is<br />
equal to F z =aρ in each subdomain, where a=Ng (N=0.1,<br />
g=9.81 m/s 2 ) and ρ is density of the corresponding material<br />
(Si or PZT-5A).<br />
166
Vp (V)<br />
3,5<br />
3,0<br />
2,5<br />
2,0<br />
1,5<br />
1,0<br />
0,5<br />
0,0<br />
-0,5<br />
-1,0<br />
3,090<br />
0,464<br />
0,050<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
0,025<br />
0,022<br />
0,021<br />
0,020<br />
0,020<br />
0,01 0,1 1 10 100<br />
Ambient pressure P 0 (kPa)<br />
Fig. 6. Peak voltage vs. ambient pressure for a constant air-film thickness<br />
h 0 = 50 μm.<br />
Vp (V)<br />
9,0<br />
8,0<br />
7,0<br />
6,0<br />
5,0<br />
4,0<br />
3,0<br />
2,0<br />
1,0<br />
0,0<br />
0 100 200 300 400<br />
Air-film thickness h 0 (μm)<br />
Fig. 7. Peak voltage vs. air-film thickness for a constant ambient pressure<br />
p 0 = 100 kPa.<br />
Firstly, frequency response analysis was conducted with<br />
parametric solver that was applied in order to sweep over<br />
the frequency range of 0 − 250 Hz. Thereby a series of<br />
amplitude-frequency characteristics were computed with<br />
different values of initial ambient pressure p 0 (Fig. 4) and<br />
air-film thickness h 0 (Fig. 5) since they determine the<br />
magnitude of the induced squeeze-film damping. Results in<br />
Fig. 5 demonstrate that when the PMPG operates under<br />
conditions of atmospheric pressure, air-film thickness of<br />
less than 100 μm leads to significant reduction of vibration<br />
amplitude implying a high level of air damping. The<br />
resonance peak is suppressed for h 0 less than approximately<br />
25 μm. Fig. 4 reveals that in the case of h 0 = 50 μm,<br />
ambient pressures of more than 1 kPa severely dampen the<br />
motion. It should be noted that the exerted squeeze-film<br />
damping force is nearly the same for ambient pressures that<br />
are in the range of 25 − 100 kPa. Computations reveal that if<br />
the PMPG has to operate under h 0 of less than 50 μm, the<br />
working environment should be rarefied to pressures not<br />
exceeding 10 Pa in order to avoid excessive levels of air<br />
damping. This is also confirmed by results in Fig. 6, which<br />
exhibits variation of generated peak voltage as a function of<br />
ambient pressure when h 0 = 50 μm. The graph reveals a<br />
pronounced increase in voltage when p 0 is reduced from 100<br />
Pa to 10 Pa. Fig. 7 illustrates that influence of air damping<br />
when operating at atmospheric pressure is also appreciable<br />
for larger air gaps of several hundred micrometers.<br />
Fig. 8. Time responses of end point of the cantilever structure to a harmonic<br />
base excitation at 192 Hz for p 0 = 100 kPa and with different air-film<br />
thickness h 0: 10 μm, 25 μm, 50 μm, 100 μm (curves from bottom<br />
to top in positive y-axis direction). The topmost curve (green)<br />
is a time response with damping excluded.<br />
Squeeze-film damping influence on PMPG dynamics is<br />
also obvious from results of transient analysis (Fig. 8),<br />
which was carried out by applying sinusoidal kinematic<br />
excitation at frequency of 192 Hz, which is close to the<br />
fundamental frequency of the transducer. One can observe<br />
that under atmospheric pressure air gaps of less than 100<br />
μm significantly reduce vibration amplitude, which is in<br />
agreement with findings of harmonic analysis.<br />
III.<br />
FABRICATION AND CHARACTERIZATION OF<br />
PIEZOELECTRIC THIN FILMS<br />
A. Formation of PVDF Films<br />
A mixture of PVDF pellets and granules (Mr ~ 180,000;<br />
Mw ~ 534,000; produced by Aldrich) was chosen for the<br />
purpose of this research since they could be easily dissolved<br />
in common polar solvents, resulting in relatively simple<br />
PVDF thin film formation process. Dimethylformamide<br />
(DMF) was chosen as a solvent ([8-10] reports that PVDF<br />
dissolves better in polar solvents), and 10%wt PVDF<br />
solution was produced while stirring the mixture of PVDF<br />
and DMF at 100°C for 20 minutes. Parallel to, silicon<br />
substrate was cleaned by boiling it in acid cleaner, treating<br />
ultrasonically in acetone and etching with plasma. The<br />
solution was transferred onto the substrate by means of dip<br />
coating, later on drying the solvent at 110°C for 10 minutes<br />
and melting the produced film at 200°C for another 10<br />
minutes in electric furnace. As a result, PVDF film of 15<br />
μm thickness was obtained.<br />
B. Experimental Study<br />
The main purpose of the experimental investigation was<br />
to analyze the morphology and crystallinity of produced<br />
PVDF thin film since these are one of the most important<br />
parameters affecting piezoelectric PVDF properties. To<br />
date, at least four crystalline structures – β, γ, δ and ε – with<br />
permanent dipole moment are described for PVDF. In all<br />
these crystal forms the chains are packed in the unit cell in<br />
167
(a)<br />
Fig. 9. Morphology of PVDF thin film obtained by means of industrial<br />
microscope Nikon Eclipse LV150. Magnification: (a) 10×, (b) 100×.<br />
(a)<br />
(b)<br />
Fig. 10. SEM micrographs obtained at magnification of: (a) 325×, (b) 1900×.<br />
(b)<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
different magnification levels: a) 325×, b) 1000×. The<br />
results confirm the formation of spherulitic structure with<br />
spherulites of 3-5 μm that are characteristic to β phase, and<br />
are particularly clearly visible in Fig. 10 (b).<br />
Final morphology analysis was performed by means of<br />
atomic force microscope MTM NT-206, while “Surface<br />
View” software was used for data processing, visualization<br />
and analysis. The roughness of the surface was determined<br />
to be relatively high as indicated by high average roughness<br />
(R a ) and root means square (R q ) values. Data analysis and<br />
Fig. 11 also reveal that investigated surface is dominated by<br />
deep valleys (skewness coefficient R sk is negative).<br />
Moreover, Fig. 12 proves the existence of two different<br />
phases of formed PVDF film. In general, the overall<br />
microscopy results comply well with those, described in<br />
other scientific papers and reveal that α phase is<br />
predominant over β in case of the analyzed PVDF sample.<br />
Assumption that α phase is more easily obtained as<br />
always resulting from melt crystallization at any<br />
temperatures was also verified by IR analysis, which was<br />
performed by Nicolet 6700 FT-IR spectrometer. The<br />
measurements were taken in the range of 500 – 1500 cm -1 ,<br />
which covers the fingerprint region for crystalline phases of<br />
PVDF. Intensive absorption bands at 611, 766, 797, 855,<br />
975 and 1414 cm -1 correspond to α phase, whereas β phase<br />
is revealed only by peak at 1251 cm -1 in Fig. 13.<br />
The same applies to diffractogram of PVDF sample<br />
registered by DRON-3 X-ray powder diffractometer with<br />
Cu Kα radiation. XRD spectra α usually contain sharp peaks<br />
due to crystallities, which in Fig. 14 are observed at<br />
2θ=20.14° (referent to the sum of the diffractions in plane<br />
(110) and (200) characteristic to β phase) and 18.30°<br />
(referent to the diffraction in (020) plane, characteristic to α<br />
phase) as well as 26.52° (referent to the diffraction in (021)<br />
plane, characteristic to α phase). The respective d spacing<br />
was calculated for mentioned values of 2θ and is presented<br />
in Table II.<br />
TABLE II<br />
CALCULATED d SPACING VALUES FOR 2θ<br />
Phase α Phase β Phase<br />
2θ 18.30 20.14 26.52<br />
d (Å) 4.848 4.405 3.353<br />
such a way that the dipoles associated with individual<br />
molecules are parallel, leading to a nonzero dipole moment<br />
of the crystal, yet in the fifth crystal modification (α)<br />
molecular dipoles are antiparallel and there is no net crystal<br />
dipole. Microscopy, X-ray and infrared (IR) tests may<br />
reveal the crystalline structures, as for each case there are<br />
typical documented results.<br />
Firstly, the morphology of thin film was obtained by<br />
means of industrial microscope Nikon Eclipse LV150. Fig.<br />
9 displays PVDF samples magnified by a factor of 10 and<br />
100. The images reveal that the surface of the films is<br />
homogeneous, with distinct spherulitic structure, indicating<br />
presence of different crystalline phases.<br />
Further insight into surface morphology was gained by<br />
means of versatile scanning electron microscope (SEM)<br />
Raith eLiNE. Fig. 10 presents micrographs obtained at<br />
Fig. 11. AFM images on PVDF samples: 3D surface morphology.<br />
168
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
IV. CONCLUSIONS<br />
A multiphysics finite element model of a PMPG was built<br />
that couples mechanic, piezoelectric and fluidic domains.<br />
The latter is represented by air-film between bottom face of<br />
the proof mass and stationary ground surface. Nonlinear<br />
compressible isothermal Reynolds equation is used to<br />
evaluate counter-reactive pressure force that is generated by<br />
squeezed air-film during operation of the generator.<br />
Numerical dynamic analyses revealed that significant air<br />
damping may be induced under atmospheric pressure<br />
leading to substantial reduction of voltage output,<br />
particularly for air gaps below 100 μm. At gaps below 50<br />
μm (at atmospheric pressure), the generated open circuit<br />
voltage is negligible as the exerted air pressure force is so<br />
Fig. 12. AFM images on PVDF samples: 3D phase identification. high that it suppresses the resonance. In this case only<br />
reduction of ambient pressure below ca. 10 Pa may restore<br />
power generation capability of the PMPG. These results<br />
demonstrate that during design of the device its<br />
configuration has to be tailored so as to minimize<br />
detrimental influence of squeeze-film damping.<br />
Application of PVDF films is planned for fabrication of<br />
the PMPG in the future. This study reported initial results of<br />
experimental characterization of morphology and<br />
crystallinity of the produced polymer films with thickness<br />
of 10 − 20 μm. Application of different analysis techniques<br />
revealed that α phase dominates in all the samples.<br />
ACKNOWLEDGMENT<br />
This research was performed under postdoctoral<br />
fellowship, which is funded by EU Structural Funds project<br />
“Postdoctoral Fellowship Implementation in Lithuania”.<br />
Fig. 13. FT-IR absorption spectra of PVDF sample.<br />
Fig. 14. X-ray diffractogram.<br />
REFERENCES<br />
[1] R.J.M. Vullers, R. van Schaijk, I. Doms, C. Van Hoof, R. Mertens,<br />
"Micropower energy harvesting," Solid-State Electron., vol. 53,<br />
pp. 684-693, 2009.<br />
[2] K.A. Cook-Chennault, N. Thambi, A.M. Sastry, "Powering<br />
MEMS portable devices - a review of non-regenerative and<br />
regenerative power supply systems with special emphasis on<br />
piezoelectric energy harvesting systems," Smart Mater. Struct.,<br />
vol. 17, 2008.<br />
[3] A. Khaligh, P. Zeng, C. Zheng, "Kinetic Energy Harvesting Using<br />
Piezoelectric and Electromagnetic Technologies−State of the Art,"<br />
IEEE Trans. Ind. Electron., vol. 53(3), pp. 850-860, 2010.<br />
[4] D. Zhu, M.J. Tudor, S.P. Beeby, "Strategies for increasing the<br />
operating frequency range of vibration energy harvesters: a<br />
review," Meas. Sci. Technol., vol. 21, 2010.<br />
[5] R.M. Lin, W.J. Wang, “Structural Dynamics of Microsystems -<br />
Current State of Research and Future Directions,” Mech. Syst. Sig.<br />
Process., vol. 20, pp. 1015-1043, 2006.<br />
[6] K.S. Breuer, “Chapter 9. Lubrication in MEMS,” in The MEMS<br />
Handbook, M. Gad-el-Hak, ed. CRC Press, 2002.<br />
[7] T. Veijola, H. Kuisma, and J. Lahdenperä, “The influence of gassurface<br />
interaction on gas film damping in a silicon<br />
accelerometer,” Sens. Actuators, A, vol. A66, pp. 83-92, 1998.<br />
[8] J. Inderherbergh, "Polyvinylidene fluoride (PVDF) appearance,<br />
general properties and processing," Ferroelectrics, vol. 115, pp.<br />
295-302, 1991.<br />
[9] R. Gregorio, "Determination of the alpha, beta, and gamma<br />
crystalline phases of poly(vinylidene fluoride) films prepared at<br />
different conditions," J. Appl. Polym. Sci., vol. 100(4), pp. 3272-<br />
3279, 2006.<br />
[10] V. Sencadas, R.G. Filho, "Processing and characterization of a<br />
novel nonporous poly(vinilidene fluoride) films in the beta phase,"<br />
J. Non-Cryst. Solids, vol. 352, pp. 2226-2229, 2006.<br />
169
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Interfacial Configurations and Mixing Performances<br />
of Fluids in Staggered Curved-Channel Micromixers<br />
Jyh Jian Chen, Chun Huei Chen, and Shian Ruei Shie<br />
Department of Biomechatronics Engineering, National Pingtung University of Science and Technology<br />
1, Shuefu Road, Neipu, Pingtung 91201, Taiwan<br />
Abstract- A parallel laminar micromixer with staggered<br />
curved channels is designed and fabricated in our study. The<br />
split-and-recombination (SAR) structures of the flow channels<br />
result in the reduction of the diffusion distance of two fluids.<br />
Furthermore, the impinging effects increase the mixing<br />
strength whereas one stream is injected into the other. The<br />
particles trajectories are utilized to numerically examine the<br />
mixing and fluidic behaviors inside the staggered curved<br />
microchannel with tapered structures. The effects of various<br />
Reynolds numbers and channel configurations on mixing<br />
performances are investigated in terms of the experimental<br />
mixing indices and the computational interfacial patterns.<br />
I. INTRODUCTION<br />
Because of the vast application fields of micromixers, such<br />
as DNA hybridization [1], direct methanol fuel cell (DMFC)<br />
[2] and cell sorting [3], the mixing efficiency in these devices<br />
is very important for the overall process performance. With<br />
the progressing of microfabrication technology, micromixers<br />
gradually move from the sub-systems of micro total analysis<br />
systems into the crucial components of MEMS. Mixture of<br />
fluids in a microchannel is strongly restricted to molecular<br />
diffusion due to the low Reynolds number. In order to speed<br />
up the mixing process in microfluidic systems, passive<br />
micromixers with the advantages of low cost, easy<br />
fabrication and no additional power have been applied in the<br />
development to enhance mixing processes.<br />
Parallel laminated mixers with simple two-dimensional<br />
structures are fabricated without difficulty, and mixing in<br />
such laminar flows can be very easily enhanced. Two<br />
representative micromixers were discussed in detail before.<br />
One design splits the main stream into several narrow<br />
streams and rejoins them together. A circular vortex<br />
micromixer with several tangential inlets was presented by<br />
Bohm et al. [4]. The mixing could be performed in a shorter<br />
timescale. The other design is a device with multiple<br />
intersecting channels. Nguyen et al. [5] demonstrated a<br />
micromixer with a square obstacle on the square-wave flow<br />
channel. Results showed that mixing index increased rapidly<br />
with decreasing microchannel width.<br />
When liquid is directed through curved channels, the fluid<br />
at the center experiences a higher centrifugal force than the<br />
surrounding liquid. Therefore, a pair of counter-rotating<br />
vortices is generated and ejects fluid toward the outer wall;<br />
this will enhance the stretching and folding of the flow<br />
element. This mechanism has been employed by many<br />
researchers for heat transfer enhancement [6, 7]. These<br />
vortices (known as Dean Vortices) as a result of differential<br />
centrifugal forces acting on the fluid at the center and at the<br />
surrounding regions also provide enhanced mixing. Howell<br />
et al. [8] fabricated a micromixer with three quarters of a<br />
circular channel. The longitudinal variation of the radial<br />
distribution of the dye is evident. While increasing the aspect<br />
ratio increases the mixing. Yamaguchi et al. [9] expressed<br />
that the interface configuration was affected by secondary<br />
flows induced by centrifugal forces. Simulation results were<br />
validated by images through confocal fluorescence<br />
microscope. Jiang et al. [10] presented a channel comprising<br />
four circular arcs and two straight inlet and outlet sections.<br />
For Dean Numbers, K, larger than 143 (corresponding to<br />
Reynolds numbers, Re, of 313), the interface stretching got<br />
increased and it indicated that chaotic mixing occurred.<br />
Kockmann et al. [11] presented the concentration<br />
distribution in a channel with a 90° bend. The length of the<br />
interface was enlarged by the vortex flow, and the potential<br />
for an exchange of the liquids was increased. Sudarsan and<br />
Ugaz [12] demonstrated a planar split-and-recombine<br />
micromixer. Parallel liquid streams first traveled through a<br />
curved segment that induced simultaneous 90° rotations in<br />
the upper and lower halves of the channel, at which point the<br />
flow was spilt into multiple streams that continued along<br />
curved trajectories such that each individual split stream<br />
experienced a second pair of 90° rotations. It was capable of<br />
generating multiple alternating lamellae of individual fluid<br />
species. Mouza et al. [13] illustrated a micromixer that<br />
comprises a semicircular curved channel and a<br />
split-and-recombine unit consisting of two semicircular<br />
microchannels that form a circle. At relatively low flow<br />
rates, where the secondary Dean flows were weak, the<br />
addition of geometrical features considerably promoted fluid<br />
mixing.<br />
A two-dimensional curved rectangular channels is<br />
designed in our study. The flow system is composed of<br />
several staggered three quarters of ring-shaped channels. The<br />
secondary flow patterns of the curved channels with various<br />
configurations are numerically and experimentally analyzed.<br />
In order to quantify the mixing as a function of the distance<br />
along the curved channel and the interfacial line length,<br />
linear regression is utilized to predict the interfacial line<br />
length at different mixing index.<br />
170
II.<br />
MATHEMATICAL MODEL AND NUMERICAL<br />
METHODOLOGY<br />
To study Dean Vortex flows with regard to mixing<br />
applications, the geometry of the curved channels and the<br />
schematic diagram of the physical features is expressed in<br />
Fig. 1. The flow system is composed of several staggered<br />
three quarters of ring-shaped channels. The angle between<br />
the lines from the center to two intersections of two<br />
consecutive channels is 90°, and the angle between two lines<br />
of the centers of three consecutive channels is 0°.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
channel. For an accelerated convergence, the algebraic<br />
multigrid (AMG) iterative method is applied for pressure<br />
corrections, and the conjugates gradient squared (CGS) and<br />
preconditioning (Pre) solvers are utilized for velocity and<br />
species corrections. The solution is considered converged<br />
when the relative errors of all independent variables are less<br />
than 10 -4 between successive sweeps.<br />
Poor grid systems can enhance numerical diffusion effects.<br />
If liquid fluids flow diagonally through the simulated grid,<br />
then the numerical effect takes the form of an extra high<br />
diffusion rate. In the proposed grid systems, meshing is<br />
generally aligned in the flow direction in the computational<br />
domain (Fig. 2). Grid-sensitivity tests are done for the preset<br />
Re with several grids. The values of mixing index at the<br />
outlet section for the five mesh densities are also shown in<br />
Table 1. Finally, the mesh density with 8.663×10 5 has been<br />
chosen for further investigation since the mixing indices at<br />
the specific location are almost the same and the numerical<br />
results are grid-independent.<br />
Fig. 1. Schematic diagram of the physical features.<br />
The numerical results presented in this work are based on<br />
the solution of the incompressible Navier-Stokes equation<br />
and a convection-diffusion equation for a concentration field<br />
by means of the finite-volume method.<br />
U 0=⋅∇ (1)<br />
<br />
−∇=∇⋅<br />
μ<br />
2∇+<br />
UPUU<br />
<br />
2<br />
DU<br />
∇=∇⋅<br />
φφρ<br />
(3)<br />
ρ (2)<br />
where U is the fluid velocity vector, ρ is the fluid density, P<br />
is the pressure, μ is the fluid viscosity, φ is the mass<br />
concentration and D is the mass diffusivity. Eq. (3) must be<br />
solved together with Eqs. (1) and (2) in order to achieve<br />
computational coupling between the velocity field solution<br />
and the concentration distribution.<br />
The dimensionless groups characterizing the Dean Vortex<br />
flows are the Reynolds number, which expresses the relative<br />
magnitudes of inertial force to viscous force.<br />
Re = UD H<br />
ν<br />
(4)<br />
where U, D H , and υ denote the velocity, the hydraulic<br />
diameter, and the kinematic viscosity, respectively, and the<br />
Dean number, which expresses the relative magnitudes of<br />
inertial and centrifugal forces to viscous force<br />
= Re<br />
(5)<br />
( ) 5.0<br />
H<br />
RDK<br />
where R is the radius of curvature.<br />
Three-dimensional structured grids are employed, and the<br />
SIMPLEC algorithm is used. All spatial discretizations are<br />
then performed using a second-order upwind scheme with<br />
limiter. The simulation is carried out for a steady state using<br />
the commercial software CFD-ACE+ TM . A fixed-velocity<br />
condition is set at the inlet; the boundary condition at the<br />
outlet is a fixed pressure. At the inlet, the concentrations<br />
normalized to 1 and 0 are prescribed in the halves of the<br />
Number of nodes<br />
Fig. 2. The grid system in the computational domain.<br />
Table 1 The analysis of the grid size independence.<br />
Mixing index<br />
Relative difference<br />
in mixing index<br />
3.577×10 5 0.794 -<br />
4.992×10 5 0.754 5.301%<br />
6.582×10 5 0.725 4.001%<br />
8.663×10 5 0.703 3.129%<br />
9.984×10 5 0.695 1.151%<br />
The uniformity of mixing at sampled sections is assessed<br />
by determining the mixing index of the solute concentration.<br />
The standard deviation of the concentration on a cross<br />
section normal to the flow direction is calculated. And the<br />
standard deviation on the inlet cross section is also computed<br />
and introduced to normalize the one on the specific cross<br />
section. Thus the mixing index can be obtained. The mixing<br />
index φ of the solute concentration, which is defined as<br />
and<br />
σ<br />
D<br />
ϕ 1−= (5)<br />
σ<br />
D 0,<br />
1<br />
σ =<br />
II (6)<br />
D<br />
N<br />
2<br />
∑(<br />
i<br />
−<br />
ave<br />
)<br />
N i=<br />
1<br />
where σ D is the standard deviation of the concentration on a<br />
cross section normal to the flow direction, σ D,0 is the standard<br />
deviation on the inlet cross section, I ave is the averaged value<br />
of the concentration over the sampled section, and I i is the<br />
171
concentration value. The mixing index φ range is from 0 for<br />
no mixing to 1 for complete mixing.<br />
IV.<br />
FABRICATION PROCESS AND FLOW VISUALIZATION<br />
For experimental characterization of mixing performance<br />
of passive micromixers, the staggered microstructures with<br />
curved channels are fabricated. The mixer geometry consists<br />
of a structure comprising a three-quarter ring-shaped channel<br />
and two three-eighth ring-shaped channels per segment. The<br />
flow device is fabricated using a replica molding method.<br />
Initially, a thin film is fabricated by patterning a cleaned<br />
silicon wafer with epoxy-based negative photoresist (SU-8).<br />
The resist is then soft baked on a level hotplate. The channel<br />
pattern is fabricated by photolithography using a chrome<br />
photomask. After development, the master is washed and<br />
baked to fix the photoresist. Once the mold is complete, the<br />
wafer is rinsed in deionized (DI) water and dried with<br />
nitrogen. After pouring the polydimethylsiloxane (PDMS)<br />
prepolymer mixture onto the wafer, microstructures are<br />
fabricated using a PDMS replica molding process. The<br />
PDMS prepolymer mixture which is thoroughly mixing the<br />
base solution and curing agent using a 10:1 weight ratio is<br />
degassed with a mechanical vacuum pump to remove air<br />
bubbles. The PDMS is then cured in an oven and the replicas<br />
are peeled off from the mold. The inlet and outlet holes are<br />
then drilled. Methanol is used as a surfactant to prevent<br />
oxygen-plasma-treated PDMS replica and glass slide from<br />
being irreversibly bonded when aligned improperly. After<br />
bonding, the designed microchannels which consist of<br />
twenty two identical mixing elements are fabricated. Figure<br />
3 shows the image of the microchannel for one typical<br />
micromixer in this study. The radius of curvature of the<br />
channel is 550 μm. The entrance and outlet of the<br />
microchannel have 0.01 mm 2 square cross-section.<br />
Fig. 3. The image of the fabricated microchannel made of PDMS.<br />
For the mixing experiment in pressure-driven flows, two<br />
different fluids are injected into the microchannels using a<br />
programmable syringe pump (KDS-101, kdScientific Inc.,<br />
USA) at preset constant flow rates, shown in Fig. 4. The flow<br />
rates are ranging from 0.003 ml/min to 0.3 ml/min<br />
corresponding to the Re from 0.5 to 50. The experimental<br />
setup for testing the performance of the fabricated<br />
micromixers is described as follows. Two syringes are<br />
loaded with 0.31 mol/L phenolphthalein and 0.33 mol/L<br />
NaOH dissolved in 99% ethanol. The NaOH solution shows<br />
a pH value of about 13. Phenolphthalein solution, as a pH<br />
indicator, has a characteristic of changing color from<br />
transparent to purple at pH values greater than 8. As a result<br />
of the rapid reaction between phenolphthalein and NaOH,<br />
the interface between two streams turns purple within a<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
negligible time [14]. Once the steady state is attained, the<br />
color changes are observed by an optical microscope<br />
(Eclipse 50i, Nikon, Japan) and CCD camera (DC80, Sony,<br />
Japan). Images are captured at each segment along the<br />
downchannel direction. The captured images are converted<br />
into grayscale images that give the luminance of the image in<br />
256 levels, and analyzed using image processing software<br />
(MATLAB, The MathWorks, Inc., USA).<br />
Fig. 4. The setup of the measurement system.<br />
Confocal fluorescence microscopy system is used for<br />
observing of the interface of the water and fluorescent<br />
solution in a three-dimensional manner. To verify the<br />
simulation results, we use a confocal microscope (Leica TCS<br />
SP2, Leica Corp., Germany) to monitor the mixing behaviors<br />
at the cross-sections of the mixing channel. One syringe is<br />
filled with fluorescent solution (99 % DI water and 1 %<br />
Rhodamine B, Fluka, Germany) while the other is filled with<br />
DI water only. The images of the fluorescent solution are<br />
excited at 543 nm with a He-Ne green laser and the signal of<br />
fluorescent emission can be detected in red (585 to 615 nm).<br />
Only the portion containing rhodamine emits light when<br />
exposed to laser. The fluorescence is monitored with a<br />
confocal microscope equipped with an air objective (10× /0.4,<br />
∞/0.17/A). The XY cross-section is scanned with a<br />
resolution of 102×1024, and the YZ cross-section is scanned<br />
with a resolution of 1024×240 (total distance along the z-axis<br />
is 100 μm with an interval of 1 μm). The micromixers are<br />
designed for investigating the effects of various operational<br />
and geometric parameters on mixing.<br />
IV. RESULTS AND DISCUSSION<br />
A staggered curved-channel micromixer is designed and<br />
fabricated in our study. The inlet and outlet have square 100<br />
μm cross-section. As the width of the channel is constant, it<br />
equals 100 μm. When the width of the three-quarter ring is<br />
tapered, it is reduced from 100 μm to 50 μm, shown in Fig. 5.<br />
The depth of the channel is always kept at 100 μm. And then<br />
the hydraulic diameter, D H , is equal to 100μm. Results are<br />
performed for K ranging from 0.23 to 22.36, corresponding<br />
to Re between 0.5 and 50.<br />
Outlet:H 100μm×W 50μm<br />
Fig. 5. The layout of the micromixer with geometric scales.<br />
Transverse Dean Flows arise from centrifugal forces when<br />
172
fluids travel along a curved channel. This effect induces a<br />
secondary flow. Figure 6 shows concentration distributions<br />
and vector planes of various cross sections a-e. The<br />
simulation results and confocal images are illustrated at an<br />
inlet velocity of 0.5 m/s. Re is equal to 50, and K equal to<br />
22.36. The red- and blue-colored liquids are utilized in the<br />
computation. The fluid with red color stands for species A,<br />
and the fluid with blue color for species B. Fluid flows<br />
around a curved channel and the fluid near the center<br />
experiences a larger centrifugal force than that near the<br />
surrounding. The velocity along the central axis is the largest<br />
and is the most strongly affected by the centrifugal forces.<br />
Two counter-rotating vortices coinciding with its plane of<br />
curvature, above and below the symmetry plane of the<br />
channel, are created. As a result, fluid is transported in the<br />
outward direction and is transported back by recirculation<br />
along the channel walls. Thus, the vertical interface that<br />
crosses the central axis is distorted. As the fluid proceeds to<br />
the curved channel, main stream is separated as two streams.<br />
Fluids in the angled channels show blue fluids surrounded by<br />
red fluids and the lamellae of two species. It accompanies by<br />
a corresponding increase in interfacial area, shown in Figs.<br />
6(a) and (c). In vector planes, the length of the arrow means<br />
the magnitude of the velocity vector. Compared with the<br />
vector planes between two branch channels, a large amount<br />
of fluid tends to flow along the original curved channel (Fig.<br />
6(b)) and the rest of the fluid moves into the angled channel<br />
(Fig. 6(a)). The uneven split of two fluids inside the<br />
staggered channels can be observed. Then two streams<br />
merge and it produces a strong impact around the<br />
interconnection (Fig. 6(e)). And then the fluid is divided into<br />
two sub-streams again. The vertical interface observed in the<br />
inlet is heavily and permanently distorted by the Dean<br />
Vortex, SAR microstructures and the impinged effect. The<br />
mixing performance is increased. Confocal images are used<br />
to qualitatively compare with the computational results. The<br />
fluorescence images are classified into two distinct regions, a<br />
red region from rhodamine and a black region from DI water.<br />
The results of numerical results are compatible with the<br />
visual experiment.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
(c)<br />
(d)<br />
(e)<br />
Fig. 6. The mixing characteristics at nine cross-sectional areas along the<br />
downchannel.<br />
A particle trajectory is the path of a particle moving in a<br />
fluid. Being able to visualize this trajectory can be very<br />
helpful in understanding flow patterns and flow distribution.<br />
A streak line is defined as a line formed by the particles<br />
which pass through a given location in the flow field. At the<br />
steady state, the streak lines coincide with the particle<br />
trajectories. In this study, the streak lines are determined by<br />
integrating the vector equations for motion and obtained<br />
from CFD-ACE+ TM software. Top view of streak lines<br />
through the mixing channels is depicted in Fig. 7(a). The<br />
streak lines stretch from the inlet, then split into two streams,<br />
and merge into one main stream. It shows most of the fluid<br />
keeps flowing in the original channel and the rest of the fluid<br />
flows into the angled channel. After passing the second split<br />
portion, a similar trend can be observed. This uneven split of<br />
the streams increases the contact surface of the mixing fluids.<br />
Fig. 7(b) demonstrates the magnified images of the streak<br />
lines near the two split portions of the curved channel marked<br />
by two blue ovals in Fig. 7(a). Two streams merge and<br />
produce an impact around the interconnection. Due to the<br />
split-and-recombine and the impinging effects, the mixing<br />
performance can be improved.<br />
(a)<br />
(a)<br />
(b)<br />
(b)<br />
Fig. 7. (a) Top views of streak lines through the mixing channels. (b) Top views<br />
of streak lines through the mixing channels at the interconnection.<br />
173
The mixing length is a distance that a fluid will keep its<br />
original characteristics before dispersing them into the<br />
surrounding fluid. By means of the mixing length, the<br />
required channel length of the micromixer can be<br />
demonstrated. Fig. 8 shows a series of images captured at<br />
specific segments of the channels at specific flow velocity,<br />
and the fabricated microchannels which consist of twenty<br />
two identical mixing elements are studied. The flow rates are<br />
0.15 ml/min corresponding to the Re equal to 25 and K equal<br />
to 11.18. As shown in Fig. 8(a), the flow proceeds forward<br />
near the inlet, the mixing of two fluids is only through<br />
molecular diffusion across the interface of the two liquids. So<br />
two parallel streams meet at the exact center of the channel<br />
and, thus, the interface is clearly observed in the channel.<br />
Then more reacted solution stream passes through the inner<br />
half of the channel and the variation of the concentration<br />
distribution along the radical direction can be seen clearly.<br />
Furthermore, the reacted solution is spread across the<br />
channel and multiple streams become visible. For the<br />
staggered curved channels with constant-width structures<br />
shown in Fig. 8(b), the similar results can be observed. With<br />
smaller impact effect near the recombined portions in the<br />
channels than that of staggered curved channels with tapered<br />
structures, it shows the mixing is not very well. The mixing<br />
performance of the continuous curved channels is<br />
demonstrated in Fig. 8(c). Lack of SAR structures the mixing<br />
is poorer than that of staggered curved channels with tapered<br />
structures. However, the path length per segment is the<br />
longest compared with that of staggered curved channels.<br />
The Dean Vortices inside the continuous curved channels<br />
induce strong secondary flows. The mixing is comparable to<br />
that of staggered curved channels with constant-width<br />
structures. The mixing index at twenty two specific locations<br />
is measured and calculated at different micromixers. The<br />
dashed line represents a mixing index equal to 0.9, and the<br />
mixing length is the channel length required for achieving<br />
the mixing index of 0.9. Mixing index of 0.9 denotes that<br />
mixing fluids are in a well mixed status. The resulting mixing<br />
lengths are at the seventh, eighteenth and over twenty-second<br />
segments corresponding to the downstream distances of 24.5<br />
mm, 63 mm and 77 mm, respectively. Due to the large vortex<br />
flow combined with the SAR effect and the impact effect,<br />
two fluids are folded into each other. Notably, due to the<br />
increases of the interfaces of the two fluids, mixing is<br />
improved.<br />
Staggered Curved Channels with Tapered structures<br />
1 5<br />
10 15<br />
(a)<br />
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May 2011, Aix-en-Provence, France<br />
Staggered Curved Channels with Constant-width structures<br />
1 5<br />
10 15<br />
(b)<br />
Continuous Curved Channels<br />
1 5<br />
10 15<br />
(c)<br />
(d)<br />
Fig. 8. Top view images of the specific segments of the staggered channel<br />
with tapered structure at specific Reynolds number. Mixing index changes<br />
along the downchannel direction of the micromixer for different<br />
micromixers.<br />
The increased interface area of two fluids can promote a<br />
mass transfer based on diffusion. The configurations of<br />
interfacial lines between two different fluids play an<br />
important role in the microchannels. The numerical results of<br />
the interfacial line length at four specific locations are<br />
calculated. Initially, the fluid interface is described by a<br />
vertical straight line across the inlet. The shapes of the<br />
interfaces of the cross-sectional planes for staggered curved<br />
channels with tapered structures at different Re are shown in<br />
Fig. 9(a). The flow near the central axis is the most strongly<br />
affected by the inertia. In the case of Re of 10, the interfacial<br />
distortion is negligible. For Re equal to 50, the interface is<br />
much more distorted. The interface stretching factors are<br />
poltted in Fig. 9(b) as a function of the number of mixing<br />
segments. This factor is defined as the interface length at a<br />
certain position divided by the initial interface length. From<br />
the figure, it is obvious that at Re=10 nearly no stretching<br />
occurs, while for Re=50, the stretching can be seen obviously.<br />
The plots of the interfacial line length as a function of<br />
174
measured mixing indices are depicted in Fig. 9(c). In order to<br />
quantify the mixing as a function of the distance along the<br />
curved channel and the interfacial line length, linear<br />
regression is utilized to predict the interfacial line length at<br />
different mixing index. From this linear equation the<br />
R-Squared value is equal to 0.960549. It can be found that<br />
the value is a high correlation and the value of mixing index<br />
can be predicted by the value of the interfacial line length.<br />
(a)<br />
(b)<br />
R-Squared=0.960549<br />
(c)<br />
Fig. 9. Experimental mixing indices for various interfacial line lengths changes<br />
along the downchannel direction of the micromixer.<br />
IV. CONCLUSION<br />
A parallel laminar micromixer with two-dimensional<br />
curved rectangular channels is designed and fabricated in our<br />
study. The flow system is composed of several staggered<br />
three quarters of ring-shaped channels. The centrifugal<br />
forces in curved flow channels make fluids to produce<br />
secondary flows. Two counter-rotating vortices, above and<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
below the symmetry plane of the channel, coincide with its<br />
plane of curvature. The bifurcation structures of the flow<br />
channels result in the reduction of the diffusion distance of<br />
two fluids. Furthermore, the impinging effects increase the<br />
mixing strength whereas one fluid is injected into the other<br />
fluid. The confocal fluorescence images demonstrate the<br />
changes of the cross-sectional concentration distributions<br />
along the downchannel direction. Phenolphthalein solution,<br />
as a pH indicator, is used for examining the mixing<br />
characteristics of three different curved channels. It can be<br />
seen that the mixing performance of the staggered curved<br />
channels with tapered structures shows superior. The shapes<br />
of the interfaces for staggered curved channels with tapered<br />
structures at different Re are investigated. Results reveal that<br />
the interface configuration of two fluids is affected by the<br />
secondary flows, and the value of mixing index can be<br />
predicted by the value of the interfacial line length..<br />
ACKNOWLEDGMENT<br />
The authors would like to thank the National Science<br />
Council of the Republic of China, Taiwan, for financially<br />
supporting this research under Contract No.<br />
99-2313-B-020-009-. And we are grateful to the National<br />
Nano Device Laboratories for MEMS processes.<br />
REFERENCES<br />
[1] M. K. McQuain, K. Seale, J. Peek, T. S. Fisher, S. Levy, M. A.<br />
Stremler, and F. R. Haseltona, “Chaotic mixer improves microarray<br />
hybridization,” Anal. Biochem., vol. 325, pp. 215-226, 2004.<br />
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2004.<br />
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2008.<br />
[4] S. Bohm, K. Greiner, S. Schlautmann, S. de Vries, and A. van den<br />
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Analysis Systems, micro-TAS 2001, pp. 25-27, 2001.<br />
[5] T. N. T. Nguyen, M. C. Kim, J. S. Park, and N. E. Lee, “An effective<br />
passive microfluidic mixer utilizing chaotic advection,” Sensor. Actuat.<br />
B-Chem., vol. 132, pp. 172-181, 2008.<br />
[6] P. M. Ligrani, S. Choi, A. R. Schallert P., Skogerboe, “Effects of<br />
Dean vortex pairs on surface heat transfer in curved channel flow,” Int. J.<br />
Heat Mass Tran., vol. 39, pp. 27-37, 1997.<br />
[7] E. A. Sewall, D. K. Tafti, A. B. Graham, K. A. Thole,<br />
“Experimental validation of large eddy simulations of flow and heat transfer<br />
in a stationary ribbed duct,” Int. J. Heat Fluid Fl., vol. 27, pp. 243-258, 2006.<br />
[8] P. B. Howell, Jr., D. R. Mott, J. P. Golden, and F. S. Ligler, “Design<br />
and evaluation of a Dean vortex-based micromixer,” Lab Chip, vol. 4, pp.<br />
663-669, 2004.<br />
[9] Y. Yamaguchi, F. Takagi, K. Yamashita, H. Nakamura, H. Maeda,<br />
K. Sotowa, K. Kusakabe, Y. Yamasaki, and S. Morooka, “3-D simulation<br />
and visualization of laminar flow in a microchannel with hair-pin curves,”<br />
AIChE, vol. 50, pp. 1530-1535, 2004.<br />
[10] F. Jiand, K. S. Dress, S. Hardt, M. Kupper, and F. Schönfeld,<br />
“Helical flows and chaotic mixing in curved micro channels,” AIChE, vol.<br />
50, pp. 2297-2305, 2004.<br />
[11] N. Kockmann, T. Kiefer, M. Engler, and P. Woias, “Convective<br />
mixing and chemical reactions in microchannels with high flow rates,”<br />
Sensor. Actuat. B-Chem., vol. 117, pp. 495-508, 2006.<br />
[12] A. P. Sudarsan, and V. M. Ugaz, “Multivortex micromixing,”<br />
PNAS, vol. 103, pp. 7228-7233, 2006.<br />
[13] A. A. Mouza, C. M. Patsa, and F. Schönfeld, “Mixing performance<br />
of a chaotic micro-mixer,” Chem. Eng. Res. Des., vol. 86, pp. 1128-1134.<br />
[14] E. F. Caldin, Fast reactions in solution, Wiley, New York, 1964.<br />
175
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May 2011, Aix-en-Provence, France<br />
<br />
Micro probe array fabrication by using the microlens<br />
array mask through proximity printing<br />
Tsung-Hung Lin 1 , Hsiharng Yang 2 , Ching-Kong Chao 3<br />
1 Graduate Institute of Engineering, National Taiwan University of Science and Technology, Taipei 106, Taiwan<br />
2 Institute of Precision Engineering, National Chung Hsing University, Taichung, Taiwan 402<br />
3 Department of Mechanical Engineering, National Taiwan University of Science and Technology, Taipei, Taiwan 106<br />
Abstract- This study presents a novel and precision<br />
process to fabricate an array of micro metal probes. The<br />
process includes microlens array mask with the proximity<br />
printing in ultraviolet (UV) lithography and Ni electroforming<br />
technology. The tip formation of micro cone probe array<br />
utilizes the microlens array mask through geometrical optics.<br />
Due to the light pass through a microlens, a microlens has a<br />
focal point. The simulated results of various focal lengths using<br />
different diameter of microlenses, the different photoresist<br />
microcone probe array molds can be fabricated. The micro<br />
cone probe array will have great potential in the area of field<br />
emission display applications.<br />
I. INTRODUCTION<br />
Recently, the microprobes play important roles in<br />
many fields, such as probes used in scanning probe<br />
microscopy (SPM) [1] and atomic force microscope (AFM)<br />
[2], microprobes in field emission [3], probe cards [4] and<br />
data storage [5], and so on. The semiconductor and MEMS<br />
devices become smaller and testing process during their<br />
production should follow such a high density trend. The<br />
probe card provides the interface between test equipment<br />
and IC device. In the probe card, controlled collapse chip<br />
connection type probe is usually used and many groups have<br />
developed different types of probe card like cantilever type<br />
and vertical type. The cantilever type contact probe was<br />
used for the linearly arranged electrode pad. In case of the<br />
pads which are irregularly arranged on the entire area of a<br />
chip, the cantilever type contact probe cannot satisfy the<br />
requirement. The vertical type probe cards are required to<br />
measure the controlled collapse chip connection type<br />
devices which have an irregular arrangement. The primary<br />
approach fabrication processes for the vertical probe tip was<br />
bulk- micromachining using deep reactive ion etcher and<br />
electro-plating with the material of Ni. In this study, the<br />
fabrication of micro-cone vertical probe array is proposed<br />
here to provide a novel method. The micro-cone vertical<br />
probe tips with various angles were produced. The<br />
microprobes can be fabricated using several manufacturing<br />
processes, which create small mechanical structures of<br />
silicon, polymer and metal. However, the probe tip can have<br />
several shapes, such as, quadrilateral pyramid, and cone, etc.<br />
There are four main approaches for fabrication of the<br />
different material probe tips. The first is an etching<br />
technique that includes chemical etching and plasma etching<br />
[6, 7]. The silicon chemical etching uses silicon electrolytic<br />
anodization in aqueous hydrofluoric acid, in combination<br />
with light, to etch patterns onto the silicon. The chemical<br />
etching based techniques can produce very sharp tips using<br />
the under-cut control strategy. However, it is difficult to<br />
obtain high density arrayed micro probes with well<br />
controlled morphologies.<br />
Because the etching rate is hard to control and the<br />
working area is restricted by the wafer size, plasma etching<br />
has the drawback of requiring expensive plasma-based<br />
equipment. The second fabrication process to realize<br />
all-metal probes is to use a focused ion or electron beam to<br />
crack an organometallic gas. The resulting tips have a good<br />
aspect ratio and radius of curvature, but this serial process is<br />
slow and it is difficult control the shape of the tips [8]. The<br />
third approach for fabricating tips-shaped polymer<br />
microstructures with patterned metallic coatings has been<br />
developed. This process involves three techniques including<br />
micro-molding, patterned metal layer transfer, and<br />
electrochemical-base sacrificial layer [9].<br />
The fourth fabrication process for the probe tips is<br />
bulk-micromachining using a deep reactive ion etcher or<br />
multiple-exposure in ultraviolet (UV) lithographic and<br />
electro-plating with Ni or Ni–Co material [10-12]. This<br />
study presents a novel process for fabricating the micro<br />
metal probe array. The proposed process will have great<br />
potential in the area of field emission display applications.<br />
The fabrication method of the micro-cone vertical probe<br />
array is presented. The experimental results showed that<br />
micro metal vertical probes with 45° and 60° tip angles. The<br />
proposed method can precisely control the geometric profile<br />
of vertical probe array. This work also offers the new<br />
fabrication method for probe card fabrication.<br />
II. Lithography characteristics<br />
Microlens projection lithography is a kind of non-contact<br />
projection lithography [13]. A plano-convex micorlens<br />
comes in one of its surfaces plane (plano) and the other<br />
convex. Plano convex microlenses have a positive focal<br />
176
length, which enables them to focus the parallel light or<br />
make parallel light out of point. The proximity exposure<br />
operations using plano-convex microlens on a mask mode<br />
occur the effect of refractive focusing. Fig. 1(a) shows the<br />
simulation plots for focusing characteristics of the<br />
plano-convex microlens mask using the software TracePro.<br />
As a parallel beam passing through a glass layer (refractive<br />
index n glass =1.8) with thickness of 1mm and a plano-convex<br />
photoresist microlens (refractive index n Az4620 =1.55) with<br />
diameter of 60μm and height of 5um, the obtained focal<br />
length is about 112μm. Fig. 1(b) shows the intensity<br />
distribution in the photoresist and after development as the<br />
light passes through the microlens array mask. A smooth<br />
and convex cone light intensity profile in photoresist. The<br />
concave micro-cone profile can be fabricated through proper<br />
operational parameters using a positive photo-resist. The<br />
desired microstructures are formed after the development<br />
process.<br />
The exposure gap (d) was set to focal length of a<br />
plano-convex lens. Fig. 2 shows the simulated relationship<br />
between the exposure gap and diameter of lens. The final<br />
concave mold micro-structure geometry can be determined<br />
using the resulting intensity distribution after exposure and<br />
development.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Fig. 2. Relationship between the focal length and microlens diameter.<br />
III.<br />
EXPERIMENTAL SECTION<br />
A. Preparation of microlens arrays mask<br />
Microlens array in photoresist (AZ4620) was firstly<br />
fabricated. When photoresist patterns are formed by<br />
lithography process, the photoresist patterns can be heated<br />
above its glass temperature. It allows surface tension to<br />
produce plano-convex microlens array. The areas between<br />
neighboring microlenses can be covered with an opaque<br />
layer of metals to block the transmission of stray light. This<br />
layer of metals acts as an aperture stop that avoids the<br />
formation of features in the area of photoresist not covered<br />
by the lenses. Fig. 3 schematically shows the process for the<br />
fabrication of microlens array mask. Fig. 3 (a)(b)(c)(d) show<br />
the formation of aperture stops by lift-off of photoresist with<br />
metals and the production of aperture stops using<br />
sputter-deposited, respectively. The process requires the use<br />
of an aligner to fabricate microlens on top of the aperture<br />
stops in Fig. 3(e)(f).<br />
(a)<br />
(b)<br />
Fig. 1. Schematic diagram showing the UV light path from the<br />
source to the substrate, (a) UV light passing through a microlen<br />
mask to substrate, (b) focal length simulation for a microlens mask<br />
with 60μm in diameter and 5μm in height.<br />
Fig. 3. Fabrication of micrlens array mask with aperture stops using<br />
reflow of melted photoresist. (a) spin coating photoresist (AZ4620) on<br />
the glass substrate, (b) photoresist pattern defined by UV lithography<br />
process, (c) sputtering a metallic film over the substrate surface, (d)<br />
removing the photoresist with solvent, (e) fabricating microlens on the<br />
top of the aperture stops using the aligner, (f) thermal reflow of melted<br />
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May 2011, Aix-en-Provence, France<br />
photoresist for microlens array formation.<br />
<br />
B. Microlens Photolithography<br />
Figure 4 shows the proposed micro probe array fabrication<br />
process. Desired patterns are transferred from the designed<br />
microlens array mask in the proximity printing ultraviolet<br />
(UV) lithographic process. In this experiment, a microlens<br />
array mask was fabricated using the thermal reflow process<br />
onto a glass. The each pohotoresist microlens with a<br />
diameter of 60μm , 70μm and the pitch distance for two<br />
adjacent microlens was 90μm. The upper and lower rows<br />
were arranged in equidistance. The Silicon wave substrate<br />
was then spun with a layer of positive photoresist (AZ4620)<br />
18μm thick. The spin condition was 500rpm for 40 seconds.<br />
Prebaking in a convection oven at 90℃<br />
for 3 minutes is a<br />
required procedure. This removes the excess solvent from<br />
the photoresist and produces a slightly hardened photoresist<br />
surface. The mask was not stuck onto the substrate. The<br />
sample was exposed through the microlens array mask using<br />
a UV mask aligner (EVG620). This aligner had soft, hard<br />
contact or proximity exposure modes with NUV (near<br />
ultra-violet) wavelength 350-450nm and lamp power range<br />
from 200-500 W. A slice of glass was inserted between the<br />
photoresist and microlens array mask to create a gap shown<br />
in Fig. 4a. The gap was adjusted to 100μm. Exposure was<br />
then conducted for about 40 seconds. The threedimensional<br />
array was completed after exposure and dip<br />
into the developer for 2 minutes and cleaning with deionized<br />
water. The micro-cone probe tips mold was produced as<br />
shown in Fig. 4(b).<br />
Fig. 4. Flow chart for micro probe array fabrication, (a) proximity UV<br />
exposure, (b) photoresist molding, (c) Ni electroforming, (d) micro probe<br />
array peel-off.<br />
C. Fabrication of micro metal probe tip<br />
Electroforming was carried out in a 10 L electroforming<br />
tank. The electroplating process requires a conductive layer<br />
to be deposited if the substrate itself is non-conductive.<br />
Therefore, a seed layer of copper (250 nm) was deposited<br />
usinganE-beam evaporator. The substrate is connected to a<br />
cathode, with nickel pellets acting as the anode. An in-tank<br />
circular filtration system including a filter tube and carbon<br />
treatment was used. The filter used in this work has 5 lm<br />
pores, which is the finest commercially available density<br />
tube. High purity is required in the electroforming process to<br />
avoid impurity deposits onto the microstructures. The<br />
template was placed into a Ni electroplating bath to form the<br />
metallic micro probes shown in Fig. 4c. The detailed<br />
ingredients of the Ni electroplating bath are listed in Table I.<br />
The deposition of Ni was uniformly controlled using an air<br />
pump for agitation to mix the electrolyte bath. Because of<br />
the micrometer range feature size at the end, a very slow<br />
deposition rate at 1 ASD was applied for 2 h to maintain the<br />
completed step coverage and duplication quality. The<br />
sample microstructure was observed as shown in Fig. 4d.<br />
Optical microscopy (OM) and a 3D surface profiler were<br />
used to measure the characteristics of the resulting<br />
microcone probe array structures.<br />
TABLE I Ni electrolyte composition<br />
Ni (NH2SO3)2_4H2O 500 (g L -1 )<br />
Boric acid 45(gL -1 )<br />
Current density 1 ASD (A dm -2 )<br />
pH 4<br />
Temperature 50 °C<br />
Agitation<br />
Air pump<br />
Wetting agent 3 (mL L -1 )<br />
Ⅳ. RESULTS AND DISCUSSION<br />
The lithography using microlen array mask can produce<br />
the micro-cone probe tips mold. As shown in Fig. 5, under<br />
the conditions of 150 °C high temperature and 10 minutes,<br />
patterns of microlens were successfully formed in the<br />
aperture stops over the whole glass substrate by using OM<br />
observation. From measurement, the microlens array was<br />
found to have a diameter of 60μm and height of 5μm.<br />
According to the experimental results, the micro cone probe<br />
arrays mold were classified after development using<br />
different focal length and diameter of lens. Fig. 6 shows<br />
micro probe arrays mold by using OM observation. As the<br />
exposure gap is 100μm and pohotoresist microlens with a<br />
diameter of 60μm, micro probe arrays mold with concave<br />
cone were formed. Then, using the exposure gap is larger;<br />
the photoresist structure is flat. The UV light and photoresist<br />
of gap is less; a flat down micro cone mold was formed. A<br />
small exposure gap is not suitable for micro cone mold<br />
fabrication because the thick photoresist will not have<br />
enough thickness to produce concave structures. The<br />
concave micro cone structure surface is quite smooth. The<br />
micro metal probes after the electroforming process. The Ni<br />
micro probes was fabricated. The 3D surface of Ni micro<br />
cone probe profiler was measured using 3-D surface profiler<br />
in Fig. 7. The fabricated structure was fine and had clear<br />
surface. Using the proximity ultraviolet (UV) lithography<br />
methodtofabricate micro cone probe mold and furthermore<br />
replication of Ni micro cone probe array is practicable.<br />
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Fig. 5 OM photograph of microlen array mask in photoresist.<br />
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“Microfabricated microneedles: a novel approach to<br />
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2000.<br />
[8] T. Akiyama et al.,“Fabrication and testing of an integrated force<br />
sensor based on a MOS transistor for applications in scanning<br />
force microscopy,”Proceedings of the IEEE Micro Electro<br />
Mechanical Systems (MEMS), p 141-146, 1997<br />
[9] H. Zhou et al.,“A new process for fabricating tip-shaped polymer<br />
microstructure array with patterned metallic coatings,”Sensors<br />
and Actuators A: physical, pp. 296-301, 2009.<br />
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Fig. 6.OM photograph of microcone probe array mold in photoresist.<br />
mechanical characterization of micro electro mechanical system<br />
based vertical probe tips for micro pad measurements,”The<br />
Japan Society of applied physics, vol. 45, pp.9238-9243 ,2006.<br />
[11] B.H.Kim,H.C.Kim,S.D.Choi,K.Chun,J.B.KimandJ.H.<br />
Kim,“A robust MEMS probe card with vertical guide for a fine<br />
pitch test,”J. Micrromech. Microeng., vol.17, pp. 1350-1359,<br />
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Polym. Adv. Technol., vol. 18 , pp. 841-844, 2007.<br />
Fig. 7 Experimental results of the Ni micro cone probe array using<br />
three-dimensional profile measurement.<br />
V. CONCLUSION<br />
The fabrication method of the micro-cone probe array is<br />
presented. The fabrication process provides an effective<br />
way to manufacture a micro probe mold as the master<br />
mold for replication in mass production. The proposed<br />
method can precisely control the geometric profile of<br />
probe array. The application of this probe array hasagreat<br />
potential in the area of field emission display.<br />
ACKNOWLEDGMENT<br />
This work was supported by the National Science<br />
Council (series no. NSC98-2221-E-005-058-MY3) of<br />
Taiwan.<br />
REFERENCES<br />
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<br />
Crescent Shaped Alignment Marks Applicable to<br />
Self-alignment of Micro-parts with and without<br />
Positive and Negative Poles<br />
Shouhei Shiga 1 , Dong F. Wang 1 , Takao Ishida 2 , and Ryutaro Maeda 2<br />
1<br />
Micro Engineering & Micro Systems Laboratory, Ibaraki University (College of Eng.), Hitachi, Ibaraki 316-8511 Japan<br />
(Tel: +81-294-38-5024; Fax: +81-294-38-5047; E-mail: dfwang@mx.ibaraki.ac.jp)<br />
2 Ubiquitous MEMS and Micro Engineering Research Center (UMEMSME), AIST, Tsukuba, Ibaraki 305-8564, Japan<br />
Abstract<br />
A “crescent-shaped” binding alignment mark, more<br />
applicable to the self-alignment than reported<br />
“tear-drop/elliptical hole” pattern, has been designed and<br />
comparatively studied with other possible alignment marks. In<br />
order to further apply this novel design to micro-parts with<br />
positive and negative poles on the binding sites, a modified<br />
“crescent-shaped” pattern with an insulated space area, defined<br />
as “crescent-shaped/interval” for self-alignment of micro-parts<br />
with two poles has been therefore proposed and discussed. The<br />
fabrication process using micromachining has been studied and<br />
both the substrates and micro-parts with alignment marks have<br />
been fabricated for next self-alignment verification.<br />
Keywords- Self-alignment; Alignment mark; Crescent-shaped<br />
pattern; Surface energy, Overlap ratio; Crescent-shaped/interval<br />
pattern; Positive and negative poles<br />
I. INTRODUCTION<br />
The integration of micro-parts in alignment with an<br />
integrated circuit is a highly important task in assembly process.<br />
In any case, a uni-directional control is required since dies,<br />
packaging or optical elements, i.e. LED etc., must be positioned<br />
to the corresponding sites of the substrate with the correct<br />
angular orientation.<br />
In stead of complicated robotic manipulation or principled<br />
restriction in traditional lithography, fluidic self-assembly<br />
(FSA) is becoming an emerging technology for its high<br />
efficient in-parallel registration or three-dimensional automatic<br />
alignment. Current self-assembly techniques for micro-scale<br />
parts are based on two major mechanisms. One is<br />
capillary-driven self-assembly [1-2] and the other is<br />
shape-directed self-assembly.<br />
The size effect of square micro-parts on the capillary-driven<br />
interaction between the square micro-parts and the square<br />
binding sites was previously studied [3], and the interaction can<br />
be confirmed until 0.3 mm parts × 0.3 mm binding sites.<br />
A two-dimensional alignment mark of a tear-drop/elliptical<br />
hole with a tip angle of 60 o [4] was developed to increase the<br />
recovery angle and reduce the energy barrier to uni-directional<br />
micro-part alignment. The work reported that the standard<br />
deviation of the aligned angular orientation was 0.9 o and the<br />
lateral accuracy was 15 μm; the re-aligned assembly yield was<br />
100 %.<br />
In this study however, a two-dimensional asymmetric<br />
“crescent-shaped” alignment mark has been newly designed<br />
and a “crescent-shaped/interval” pattern has been further<br />
proposed to be applicable to micro-parts with two poles.<br />
Ⅱ. SELF-ALIGNMENT PRINCIPLE<br />
For the FSA process herein, self-alignment is caused by<br />
capillary force, which occurs between lubricant and SAM.<br />
When micro-parts are introduced onto adhesive droplets on<br />
binding sites (receptor sites), their hydrophobic faces (usually<br />
gold faces) can be attracted to the adhesive droplets. The<br />
Au-patterned side of each micro-part that is shown in Fig. 1 is<br />
hydrophobic and the other side is hydrophilic. In this case<br />
(a)<br />
(b)<br />
Si SAM Lubricant<br />
Fig. 1. Schematic illustrating self-alignment, (a): Parts move and rotate with an<br />
angle of θ for coincidence; (b):Using uni-directional Au pattern, part is<br />
self-aligned to the determined direction.<br />
180
however, a rotation for coincidence by self-alignment will<br />
occur as shown in Fig. 1(a), and the rotation has the potential to<br />
realize a uni-directional self-alignment if a special binding<br />
pattern is employed as shown in Fig. 1(b). This implies that<br />
two-dimensional asymmetric patterns are useful to aligning<br />
micro-parts to a determined direction on substrate.<br />
11-13 <br />
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<br />
surface-energy model for the self-alignment of flat silicon parts<br />
[4] can thus be applied and the total interfacial energy E 1 before<br />
self-alignment can be approximated by Equation (1):<br />
E<br />
= γ S + P<br />
(1)<br />
1 Lub , Water<br />
γ<br />
SAM , Water<br />
Ⅲ. ANALYZING SELF-ALIGNMENT USING OVERLAP RATIO<br />
To accomplish a high alignment yield, a suitable binding<br />
pattern must be a global energy minimum (with maximum<br />
overlap), while other (local) energy minima (low energy<br />
barrier) corresponding to an unsuitable pattern design must be<br />
avoided.<br />
However, the overlap ratio, as defined in Fig. 2, which is the<br />
ratio of the overlap area to the total area of the binding site, is<br />
maximized in the fully aligned orientation, and gradually<br />
declines toward the energy barrier. The overlap ratio is<br />
therefore very useful to find out suitable patterns for<br />
self-alignment. In this case, two important parameters, the<br />
energy barrier and the angular span of the designed shape, must<br />
be kept as small as possible to ensure a highly efficient<br />
uni-directional alignment. Since direction-specific and high<br />
precision self-alignment depends on adhesion force, overlap<br />
ratio and correct pattern design predicted by the surface energy<br />
model, the aforementioned parameters should be carefully<br />
analyzed by simulation with the surface energy model.<br />
where γ Lub,Water is the interfacial energy between lubricant<br />
adhesive and water, γ SAM,Water is the interfacial energy between<br />
SAMs and water, S is the binding area on substrate, and P is<br />
the binding area on micro-part. Supposing S = P in present<br />
case, the total interfacial energy E 2 after self-alignment can then<br />
be described by the following Equation (2):<br />
E2<br />
= γ<br />
γ<br />
E<br />
Lub,<br />
Water<br />
Lub,<br />
Overlap<br />
SAMLub<br />
1<br />
( AS ) γ ( − )<br />
Overlap<br />
+−<br />
SAM , Water<br />
AP<br />
Overlap<br />
×+<br />
A<br />
(2)<br />
−−+=<br />
γγγ<br />
(<br />
, SAM Lub , Water SAM , Water<br />
) A Overlap<br />
where γ Lub,SAM is the interfacial energy between lubricant<br />
adhesive and SAMs, and A overlap is the overlap area as defined in<br />
Fig. 2. If using Δ E to express the interfacial energy change as<br />
the following Equation (3),<br />
Δ<br />
( γ Lub , SAM<br />
− γ Lub , Water<br />
− γ SAM Water<br />
) A Overlap<br />
E =<br />
,<br />
(3)<br />
the total interfacial energy E 2 after self-alignment can be written<br />
as<br />
= 12<br />
+ ΔEEE (4)<br />
When an adhesive droplet sits on a substrate surface with a<br />
contact angle α, Young’s equation gives the following<br />
relationship:<br />
Fig. 2. Schematic illustration of the overlap ratio, which is defined as the ratio<br />
of the overlap area to the total area of the binding site.<br />
A. Surface energy model<br />
The interfacial energy minimization gives rise to the<br />
attraction of the micro-parts to the binding sites. As shown in<br />
Fig. 3, if the thickness of the adhesive droplet is thin enough<br />
that two dimensional approximation is valid and the sidewall<br />
interfacial energy of the adhesive is negligible, the<br />
surface-energy model for the self-alignment of flat silicon<br />
γ Lub,Water<br />
α<br />
γ SAM,Water<br />
Lub<br />
γ Lub,SAM<br />
SAM<br />
Au<br />
Fig. 3. Schematic figure of the contact angle between the lubricant and the<br />
surface modified by SAMs in water.<br />
γ<br />
,<br />
γ cosα<br />
+ γ<br />
= (5)<br />
SAM Water Lub,<br />
Water<br />
Lub,<br />
SAM<br />
The above Equation (5) can thus be rewritten as<br />
γ<br />
[ cos( α ) 1]<br />
Lub , SAM<br />
− γ<br />
Lub,<br />
Water<br />
− γ<br />
SAM , Water<br />
−= γ<br />
Lub,<br />
Water<br />
+<br />
where ∀α<br />
∈ ( 0 , 180 )°° . Equation (6) means that ΔE is less<br />
than zero for any adhesive, micro-part surface and environment<br />
medium.<br />
B. Overlap ratio<br />
From the above Equations (3) and (6), it can be noted that<br />
the interfacial energy has a minimum value when the<br />
micro-part is exactly aligned with the binding site (receptor<br />
site). Since E 1 is a constant, theΔE thus appears to be linearly<br />
proportional to the overlap area A overlap as:<br />
[ cos( α ) ] A Overlap<br />
Lub , Water<br />
+<br />
(6)<br />
ΔE ∝ −γ 1<br />
(7)<br />
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The above Equations indicate that E<br />
<br />
2 decreases in value<br />
1<br />
when A overlap increases in value. The decreasing of E 2 will make<br />
self-alignment advance towards a higher precision. The<br />
0.8<br />
self-alignment completes at the time that A overlap reaches the<br />
maximum and E 2 drops to the minimum.<br />
IV.<br />
NEWLY PROPOSED “CRESCENT-SHAPED” ALIGNMENT<br />
MARK FOR SELF-ALIGNMENT<br />
In order to accomplish a high alignment yield, a novel<br />
asymmetric alignment mark, so called “crescent-shaped ”<br />
pattern, as shown in Fig. 4, is originally designed based on the<br />
above surface-energy model, to increase the recovery angle and<br />
reduce the energy barrier to uni-directional micro-part<br />
alignment.<br />
The basic concept of the “crescent-shaped” pattern can be<br />
described using the following three terms:<br />
1. There is only one line to form a linear symmetry;<br />
2. The pattern consists of a curved shape;<br />
3. The shape is wide in the middle and pointed at each end.<br />
Overlap ratio<br />
0.6<br />
0.4<br />
0.2<br />
0<br />
Crescent-shaped<br />
Tear-drop<br />
Square<br />
0 30 60 90 120 150 180 210 240 270 300 330 360<br />
Offset angles(a)<br />
Fig. 5. Overlap ratio between a moving part and a binding site for comparison<br />
of three kinds of alignment patterns.<br />
1<br />
Wafer preparation<br />
2<br />
Cr/Au sputtering<br />
3<br />
Cr/Au etching<br />
Fig. 4. Schematic figure for detailed description of proposed crescent-shaped<br />
alignment mark, where the square means the surface of micro-part, and the<br />
crescent-shaped pattern is the binding mark for self-alignment.<br />
The overlap ratio of three kinds of alignment marks,<br />
including crescent-shaped, tear-drop/elliptical holes, and<br />
square, as a function of the offset angles has been simulated and<br />
comparatively shown in Fig. 5. It can be noted that both the<br />
proposed “ crescent-shaped ” pattern and the reported<br />
“tear-drop”pattern with elliptical holes have much lower span<br />
angles and energy barriers than the “square” pattern, and are<br />
therefore expected to have only one stable orientation<br />
compared to the “square” pattern of four maximum overlap<br />
ratios at four offset angles (0 o , 90 o , 180 o , and 270 o ). Since the<br />
energy barrier of the proposed “crescent-shaped” pattern is<br />
lower than that of the reported “tear-drop” pattern with<br />
elliptical holes, the “crescent-shaped” pattern is thus expected<br />
to be the most suitable one for self-alignment in this work.<br />
For fabricating both the substrates and micro-parts for<br />
self-alignment verification, a 300-μm-thick silicon wafer was<br />
used as a starting material.<br />
The one-mask etching process is typically shown in Fig. 6.<br />
Simply, the sputtered Cr/Au was patterned by lithography and<br />
was then wet etched to transfer the alignment mark patterns<br />
from the mask.<br />
Si Au Resist<br />
Fig. 6. Typical one-mask process chart for fabricating both substrates and<br />
micro-parts with alignment marks.<br />
Fig. 7 shows the three patterns designed for the present<br />
work (Figs. 7(a’), 7(b’), 7(c’)), and fabricated patterns after wet<br />
etching process (Figs. 7(a), 7(b), 7(c)) for comparison,<br />
respectively.<br />
(a)<br />
(a’)<br />
4<br />
Resist removing<br />
Fig.7. Typical binding site design for verification: (a’) crescent-shaped, (b’)<br />
tear-drop, (c’) square; Fabricated patterns after wet etching: (a) crescent-shaped,<br />
(b) tear-drop, (c) square.<br />
(b)<br />
(b’)<br />
(c)<br />
(c’)<br />
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V.<br />
<br />
MODIFIED CRESCENT PATTERNS FOR DEVICES WITH<br />
POSITIVE AND NEGATIVE POLES<br />
A modified “crescent-shaped” alignment mark, called<br />
“crescent-shaped/interval” as shown in Fig. 8, has been further<br />
proposed. The basic idea is to separate the “crescent-shaped”<br />
pattern into two parts with an insulated space area between<br />
them, so as to correspond to positive and negative poles,<br />
respectively.<br />
In order to study the effect of interval a, as defined in Fig. 8,<br />
on the surface energy during self-alignment, the overlap ratio as<br />
a function of the offset angles has been calculated and<br />
compared for different interval a. It can be noted that the<br />
self-alignment of micro-parts with positive and negative poles<br />
could be achieved with a maximum misalignment angle of 30 o<br />
degree, when the interval is set as 1:26, as shown in Fig.9.<br />
(a)<br />
25 a<br />
(b)<br />
Fig. 10. Target substrate design (Fig. 10(a)) for integration of prototype<br />
micro-parts (Fig. 10(b)) with positive and negative poles.<br />
Fig. 8. A modified crescent-shaped alignment mark, called<br />
crescent-shaped/interval pattern, proposed for micro-parts with positive and<br />
negative poles.<br />
VI. CONCLUSIONS<br />
A newly designed two dimensional asymmetric alignment<br />
mark, so called as “crescent-shaped” pattern, has been derived<br />
as the most suitable one for self-alignment of the micro-parts to<br />
the binding sites of the substrate, using overlap ratio based on<br />
the surface energy model of a capillary effect. The energy<br />
barrier of the proposed “crescent-shaped” pattern is<br />
theoretically lower than that of the reported “tear-drop” pattern<br />
with elliptical holes. A modified “crescent-shaped” alignment<br />
mark, so called as “crescent-shaped/interval”, has been further<br />
proposed and designed to be applicable to micro-parts or<br />
devices with positive and negative poles on the binding sites.<br />
ACKNOWLEDGEMENT<br />
Part of this work was supported by MEMS Inter<br />
University Network and performed in the Ubiquitous MEMS &<br />
Micro Engineering Research Center (UMEMSME) of National<br />
Institute of Advanced Industrial Science & Technology (AIST).<br />
Fig. 9. Overlap ratio between a moving part and a binding site with a<br />
relation to different interval a in crescent-shaped/interval pattern.<br />
A target substrate with special binding sites with<br />
“crescent-shaped/interval” alignment mark for integration of<br />
micro-parts with poles, where a uni-directional self-alignment<br />
is necessary, has been designed for further practicable study, as<br />
shown in Fig. 10. Firstly, both the Au surface on substrate and<br />
the backside of micro-parts are patterned with proposed marks.<br />
Secondly, electrodes are connected to Au wire on the substrate<br />
after self-alignment. Finally, the integration is demonstrated to<br />
examine the feasibility of uni-directional self-alignment.<br />
REFERENCES<br />
[1] A. Terfort, N. Bowden, and G.M. Whitessides, Three-dimensional<br />
self-assembly of milllimeter-scale components, Nature, 386 (1997)<br />
162-164.<br />
[2] U. Srinivasan, D. Liepmann, and R.T. Howe, Microstructure to<br />
substrate self-assembly using capillary forces, J. Microelectromech.<br />
Syst., 10 (2001) 17-24.<br />
[3] D. F. Wang, and S. Shiga, Fabrication ofmicro and nanostructures<br />
using self-assembly, Proceedings of Ibaraki District Conference<br />
2009, pp.127-128, JSME, Tsukuba, Japan.<br />
[4] C. Lin, F. Tseng, and C. Chieng, Orientation-specific fluidic<br />
self-assembly process based on a capillary effect, J. Micromech.<br />
Microeng., 19 (2009) 1-12.<br />
[5] K.F. Bohringer, U. Srinivasan, and R.T. Howe, Modeling of<br />
capillary forces and binding sites for fluidic self-assembly, Proc.<br />
IEEE workshop on Micro electro Mechanical<br />
183
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Simulation of 3D SOI-Structures for MEMS elements<br />
Igor KOGUT, Victor HOLOTA, Victor DOVHIJ (Precarpatian U., Ivano-Frankivsk, Ukraine)<br />
Anatoliy DRUZHININ (National U. "Lvivska Politechnika", Lviv, Ukraine)<br />
Text unavailable at the time of printing.<br />
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<br />
Studies of optical and crystal properties of ALD<br />
grown ZnO<br />
David Elam 1 , Anastasiia Nemashkalo 2 , Yuri Strzhemechny 2 , Chonglin Chen 1 , Arturo Ayon 1 , Andrey Chabanov 1<br />
1 Department of Physics and Astronomy, University of Texas San Antonio, One UTSA Circle, San Antonio, TX 78249<br />
2 Department of Physics and Astronomy, Texas Christian University, TCU Box 298840 Fort Worth, TX 76129<br />
Abstract- In this paper, we report structural and<br />
photoluminescence properties of ZnO thin films grown<br />
on single crystal ZnO substrates by means of Atomic<br />
Layer Deposition (ALD).<br />
I. INTRODUCTION<br />
In the last decade, zinc oxide has become a popular<br />
subject of research exhibiting a number of useful properties.<br />
For example, as a direct band gap semiconductor with a<br />
band gap in the UV at ~3.3eV, zinc oxide has applications<br />
in UV laser diodes, although large defect densities have<br />
been reported as an obstacle [1]. Doping zinc oxide with<br />
transition metals causes zinc oxide thin films to exhibit<br />
weak ferromagnetic behavior, a property which may have<br />
interesting applications in electronics [2]. Zinc oxide is also<br />
a piezoelectric material with potential applications in<br />
MEMS devices, sensing, and power generation [3].<br />
As a semiconductor, zinc oxide has seen only limited use<br />
due to its tendency to dope itself n-type. Resistivities < 1<br />
Ohm-cm have been observed in undoped films [4]. This<br />
has resulted in difficulties in reliably doping zinc oxide<br />
films p-type. Such an attempt may simply result in an<br />
insulating material. This situation has been improved<br />
recently with more mature doping techniques [5].<br />
Although, the underlying cause of this natural doping has<br />
been the subject of debate, it is thought to originate on<br />
oxygen vacancies and zinc interstitials. These defects have<br />
been demonstrated to act as electron donors. A competing<br />
candidate, however, is hydrogen impurities [6-7]. The exact<br />
cause of this doping behavior may very well be a<br />
combination of the two, and may depend on the deposition<br />
method. Hydrogen impurities may be particularly relevant<br />
in metal organic deposition techniques such as MOCVD<br />
and ALD. Understanding the defect structure of these films<br />
may help to understand the mechanical and electrical<br />
properties of the thin film.<br />
In this paper, we report structural and optical properties<br />
of ZnO thin films grown on single crystal ZnO substrates by<br />
means of Atomic Layer Deposition (ALD). Our x-ray<br />
diffraction data and photoluminescence measurements<br />
indicate a good crystallinity and low concentration of lattice<br />
defects of the ZnO films, which depend on the deposition<br />
temperature.<br />
II. RESULTS AND DISCUSSION<br />
ZnO films were grown at 120 C and 200 C using<br />
diethylzinc and water as precursors. The films were grown<br />
on oxygen-terminated single crystal ZnO substrates to a<br />
thickness of 100nm. The X-ray diffraction data were<br />
obtained using a Rigaku Ultima IV XRD. The XRD<br />
measurements show that the films are polycrystalline with a<br />
strong preferred c-axis orientation.<br />
Fig. 1. 2θ/ω scan of the (002) diffraction peak of the ZnO films, as<br />
compared to the substrate.<br />
TABLE I<br />
XRD data of the ZnO films on ZnO substrate.<br />
Sample 2-theta c lattice constant FWHM<br />
(002)<br />
Substrate 34.4532 5.20194 0.0505<br />
120C 34.4461 5.20298 0.0480<br />
200C 34.4527 5.20202 0.0496<br />
(001)<br />
120C 17.0333 5.2012 0.0372<br />
200C 17.0477 5.1968 0.0479<br />
The (002) peak, shown in Fig.1, is essentially due to the<br />
substrate, since the signal from the film and substrate<br />
overlap at the (002) peak. The second peak in the (002)<br />
diffraction pattern is from the Cu Kα 2 wavelength. On the<br />
other hand, the (001) peak, shown in Fig 2, is due to the<br />
film, as it cannot be seen in hexagonal zinc oxide. The (001)<br />
peak is due to point defects, such as oxygen vacancies,<br />
which produce deformations of the ZnO lattice. The (001)<br />
peaks in the films grown at 120 and 200 C have different<br />
shapes and are slightly shifted. The c lattice constant of the<br />
185
ZnO films were determined from the (001) peak positions<br />
(Table 1). The shift in the peak positions of the films grown<br />
at 120 and 200 C corresponds to 0.2% difference in the c<br />
lattice constant.<br />
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<br />
of the second lower-energy luminescent component usually<br />
associated with the Zn interstitial defects.<br />
To further elucidate the nature of these differences we<br />
intend to further analyze the temperature dependences of the<br />
relative intensities, spectral positions and widths of the<br />
excitonic luminescent features in the PL spectra of the<br />
studied specimens.<br />
Fig. 2. 2θ/ω scan of the (001) diffraction peak of the ZnO films grown at 120<br />
and 200 C.<br />
Our photoluminescence (PL) and surface photovoltage<br />
(SPV) spectroscopy results also indicated that the obtained<br />
thin films are of high quality. In particular we observed that,<br />
on the one hand, “deep” defect signatures in the PL and<br />
SPV spectra have low relative intensity. On the other hand<br />
the excitonic peaks in the near-bang gap range of the lowtemperature<br />
PL spectra are intense and narrow whereas the<br />
super-band gap SPV transitions are sharp and well-defined.<br />
In comparison, for ZnO thin films vapor phase-deposited on<br />
single-crystalline ZnO substrates [8] the reported excitonic<br />
lines were order of magnitude broader in the lowtemperature<br />
spectra. Also, the XRD results for those films<br />
showed a strong mosaicity which limited the structural<br />
quality of their film.<br />
Nonetheless, the thin films grown at different<br />
temperatures exhibited discrepancies in their optical and<br />
structural characteristics. For example, the relative<br />
intensity of the ~ 3.33 eV luminescence peak in the 10 K PL<br />
spectra, commonly attributed to extended structural defects<br />
[9], was significantly greater for the sample grown at 120°C<br />
(cf. Fig. 3A). Moreover, the deep level (visible)<br />
luminescence in the room temperature PL spectra is also<br />
different. For the same sample grown at 120°C the broad<br />
emission exhibits a substantial red shift (cf. Fig. 3B). The ~<br />
2.4 eV emission is most often attributed to oxygen<br />
deficiency whereas the observed shift may indicate presence<br />
Fig. 3. Photoluminescence at 10K (A) and 293K (B) of the ZnO films grown<br />
at 120 and 200 C.<br />
ACKNOWLEGDEMENT<br />
This research was supported by the US Army Research<br />
Grant No. 54484-RT-ISP and National Science Foundation<br />
Grant No. DMR-0934218.<br />
REFERENCES<br />
[1] T. P. Rao et al. Jallcom 485 (2009)<br />
[2] F. Pan et al. Materials Science and Engineering R 62 (2008) 1-<br />
35<br />
[3] Sheng Xu, et al. Nature Nanotechnology 5 (2010) 366<br />
[4] C. Jagadish and S. Pearton (Ed.) “Zinc Oxide Bulk, Thin Films<br />
and Nanostructures”, Elsevier Limited (2006)<br />
[5] Eun-Cheol Lee, K. J. Chang Physica B 376-377 (2006) 707-<br />
710<br />
[6] F. Sun et al. Applied Surface Science 256 (2010) 3390-3393<br />
[7] Y. J. Lin et. al J. Appl. Phys. 99 (2006)<br />
[8] A. Zeuner, H. Alves, D. M. Hofmann, and B. K. Meyer, M.<br />
Heuken, Bläsing, and A. Krost, Appl. Phys. Lett., 80, 2078<br />
(2002).<br />
[9] B. K. Meyer, H. Alves, D. M. Hofmann, W. Kriegseis, D.<br />
Forster, F. Bertram, J. Christen, A. Hoffmann, M. Straßburg, M.<br />
Dworzak, U. Haboeck, and A. V. Rodina, Phys. Stat. Sol. (b)<br />
241, 231 (2004).<br />
186
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<br />
A Methodology for the Pull-in Voltage of Clamped<br />
Diaphragms<br />
Joseph Lardiès, Marc Berthillier<br />
Institute FEMTO-ST; DMA; UMR CNRS 6174<br />
Rue de l’Epitaphe<br />
25000 Besançon, FRANCE<br />
Abstract- Due to the electrostatic excitation principle,<br />
MEMS based capacitive microphones exhibit a nonlinear<br />
behaviour and this communication investigates these<br />
nonlinearities with the aim to optimize the input voltage signal.<br />
The nonlinear electrostatic force due to the bias voltage is<br />
combined with the 2D load-deflection model of a square or<br />
circular diaphragm to evaluate the pull-in voltage. An<br />
analytical solution is derived to calculate the electrostatic<br />
pressure, the pull-in voltage and the deflection profile of the<br />
diaphragm. Numerical results with clamped square and<br />
circular diaphragms are presented showing the effectiveness of<br />
the method.<br />
I. INTRODUCTION<br />
MEMS-based capacitive microphones offer advantages<br />
due to their small size, high sensibility, batch fabrication<br />
capability and low power consumption. A MEMS<br />
capacitive-type microphone is basically an electrostatic<br />
transducer converting electrical energy into mechanical<br />
energy and vice-versa. A good design requires a large<br />
displacement from the bias voltage for efficient energy<br />
coupling between the movable microplate or diaphragm and<br />
the air. The microplate can also be deflected by ambient<br />
pressure if the cavity beneath the microplate is vacuum<br />
sealed, which is necessary for immersion applications.<br />
However, optimum energy coupling is achieved when the<br />
plate is near the structural instability known as pull-in,<br />
where the largest stable plate deflection occurs. Beyond this<br />
point, the movable plate snaps onto the fixed plate (or<br />
substrate). Many resonance applications demand better<br />
understanding of MEMS behaviors, especially near the pullin<br />
instability. Finite element method (FEM) simulations and<br />
analytical plate or membrane models have been used to<br />
analyze resonating microstructures effects. However, most<br />
FEM simulations are computationally inefficient or<br />
breakdown near pull-in of electrostatically actuated<br />
structures, and membrane model ignore plate bending,<br />
which is needed for bending dominate microstructures.<br />
The central component in micro-electro-mechanical<br />
systems is the mechanical resonator which constitutes a<br />
capacitive transducer and is formed with two plates: a fixed<br />
plate and a movable plate. Due to the electrostatically force,<br />
when the gap between the two plates becomes two thirds of<br />
the initial gap, the movable plate is not stable, we have a<br />
’’push-down’’ phenomenon and the MEMS fails. From the<br />
dynamics point of view, the system loses its stability and<br />
the gap being equal to two-thirds of the initial gap is termed<br />
the minimum gap in MEMS [1].<br />
The electrostatic force associated with the voltage is non<br />
linear due to its inverse square relationship with the airgap<br />
thickness between the capacitor electrodes. This gives rise<br />
to the pull-in phenomenon that causes the movable structure<br />
(membrane) to collapse if the bias voltage exceeds the pullin<br />
limit and limits the effective range of deformation of the<br />
structure. Accurate determination of the pull-in voltage, or<br />
the collapse voltage, is critical in the design process to<br />
determine the sensitivity, harmonic distortion and the<br />
dynamic range of a MEMS-based capacitive transducer. In<br />
[2] a method is provided to approximate pull-in voltage for<br />
cantilevers, fixed-fixed beams and circular diaphragms<br />
under electrostatic actuation in which the pull-in voltage<br />
depends on the undeformed gap and on the linear elastic<br />
response to an applied uniform load. In this communication<br />
an analytical solution is described to calculate the pull-in<br />
voltage and diaphragm deflection under electrostatic<br />
actuation, for square and circular diaphragms. The method<br />
incorporates both the nonlinearities of the electrostatic force<br />
and the large deflection model for a clamped square or<br />
circular diaphragm. The developed analytical method<br />
allows for a fast, more accurate determination of the<br />
developed electrostatic pressure, maximum diaphragm<br />
deflection for different bias voltage and the pull-in voltage.<br />
The method can easily be extended to the cases of<br />
cantilevers and fixed-fixed beams.<br />
II.<br />
MODEL DEVELOPMENT<br />
A. Parallel-Plate Approximation<br />
A parallel plate approximation is first considered to<br />
highlight the major aspects of the analysis. A schematic cross<br />
section of a MEMS capacitive-type microphone is shown in<br />
Fig. 1. An external voltage V is applied between the upper and<br />
lower conductors, which causes the upper conductor to<br />
electrostatically deflect downwards. Deflection increases with<br />
voltage until pull-in is reached. A static displacement of the<br />
diaphragm of a capacitive cell due to the bias voltage is shown<br />
187
this figure.<br />
Fig. 1. Cross-section of a MEMS-based capacitive sensor.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
2<br />
xd<br />
M + Kx = Fe (x) =<br />
2<br />
td<br />
ε<br />
0<br />
0<br />
2<br />
VA<br />
(1)<br />
2<br />
− )xd( 2<br />
The mechanical elastic force is F m (x) = Kx and ε 0<br />
is the<br />
permittivity of the free space. At the static equilibrium the<br />
mechanical elastic force equals the electrostatic attraction<br />
force and the relationship between the voltage V and<br />
displacement of the movable plate is :<br />
We introduce a simplified one-dimensional (1-D) pull-in<br />
model in which the pull-in voltage depends on the<br />
undeformed gap and on the linear elastic response to an<br />
applied uniform load. While not numerically accurate, this<br />
model has the virtue of providing a functional form, which,<br />
for many structures, can be approximated analytically by<br />
solving a suitable linear equation. Fig.2 shows a lumped 1-<br />
D pull-in model, which provides some guidance in how the<br />
functional form is developed. The problem is approximated<br />
by a rigid body suspended by a lumped linear spring with<br />
spring constant K.<br />
V=(d 0 – x) ε(/xK A) 2 (2)<br />
The maximum of the voltage is obtained for dV/dx=0 and<br />
from this equation we obtain the distance where the pull-in<br />
occurs: x pi = d 0 /3 and the pull-in gap is d pi = 2d 0 /3. The pull-in<br />
voltage for this ideal parallel plate structure is then :<br />
3<br />
0<br />
0<br />
V pi = ε2 A) 7(/dK8<br />
(3)<br />
and the spring constant of the movable plate is given by:<br />
0<br />
3<br />
K = 27 ε 0<br />
AV pi / (8 d 3 0 ) (4)<br />
Fig. 2. A simplified mechanical model for a parallel-plate capacitor.<br />
As shown in Fig. 2 the fixed plate of the capacitor with<br />
area A is connected with a constant supply voltage V. The<br />
other plate of the capacitor with mass M and area A is movable<br />
and rigid. The support of the moving plate is modeled through<br />
an equivalent spring with stiffness K. Without any electrostatic<br />
force, the gap between two plates of the capacitor in the<br />
MEMS is d 0 : it is the initial air gap. The coordinate x is shown<br />
in Fig. 2 and the origin is at equilibrium without any voltage.<br />
By applying a voltage across the plates, an electrostatic<br />
attractive force F e (x) is induced which leads to a decrease of<br />
the gap spacing, thereby stretching the spring. This results in<br />
an increase of the mechanical elastic force (or spring force)<br />
F m (x) which counteracts the electrostatic force. Pull-in<br />
instability occurs as a result of the fact that the electrostatic<br />
force increases non-linearly with decreasing gap spacing,<br />
whereas the mechanical elastic force is a linear function of the<br />
change in the gap spacing. In simple terms, the pull-in voltage<br />
can be defined as the voltage at which the restoring spring<br />
force can no longer balance the attractive electrostatic force.<br />
Our purpose is to determine this pull-in voltage.<br />
B. First Order Analysis<br />
Neglecting any damping within the system, the equation of<br />
motion of the movable plate due to an electrostatic attraction<br />
force F e (x) caused by a constant supply voltage V is:<br />
If the applied voltage is increased beyond the pull-in<br />
voltage, the resulting electrostatic force will overcome the<br />
elastic restoring force and will cause the movable plate to<br />
collapse on the fixed plate and the capacitor will be short<br />
circuited.<br />
To obtain stiffness due to the electrostatic force we expand<br />
(1) using a Taylor series approximation about the nominal<br />
distance x 0 :<br />
2<br />
2<br />
xd ⎛<br />
M + ε<br />
2<br />
td<br />
⎟ ⎞<br />
0<br />
VA<br />
⎜ K − x =<br />
3<br />
⎝ −<br />
00<br />
)x( ⎠<br />
d<br />
1<br />
2<br />
ε<br />
0<br />
2<br />
⎡<br />
− 1n<br />
VA x2 N<br />
⎤<br />
0<br />
−<br />
0<br />
)xx(n<br />
⎢1<br />
− ∑<br />
−<br />
⎥ (5)<br />
2<br />
⎢ − 3n<br />
⎣<br />
−<br />
1n<br />
00<br />
)xd(<br />
00<br />
)xd(<br />
00<br />
)xd(<br />
⎥⎦<br />
− =<br />
The electrostatic attraction force effectively modifies the<br />
spring constant K of the movable plate and the effective spring<br />
constant at a specified voltage V is :<br />
2<br />
⎛ ε ⎞<br />
K effective =<br />
0<br />
VA<br />
⎜ K − ⎟<br />
(6)<br />
3<br />
⎝ −<br />
00<br />
)x( ⎠<br />
d<br />
The amount of modification is termed as spring softening<br />
and the resonant frequency of the structure is shifted from<br />
ω res<br />
= M/K to ωres = M . /K<br />
effective<br />
The simple parallel-plate approximation method assumes<br />
that the beam has a linear spring constant, considers a piston<br />
like motion, and predicts that the pull-in when the highest<br />
deformation exceeds one-third of the gap. This analysis<br />
neglects the effects of the fringing field capacitances and<br />
188
excludes the nature of the fixed boundary conditions, nonuniformity<br />
of the electrostatic pressure, effects of the residual<br />
stress, and the developed nonlinear distribution due to the<br />
stretching of the beam. For wide beams with small airgaps,<br />
errors up to 20% have been reported in literature due to such<br />
approximations [2]. In the next section we analyze the case of<br />
a rigidly clamped square diaphragm separated from a rigid<br />
backplate by a small airgap.<br />
III.<br />
SQUARE DIAPHRAGM ANALYSIS<br />
A. Load-deflection characteristics of a square diaphragm<br />
The load-deflection analysis has been developed for the<br />
measurement of the mechanical properties of thin films [3-5],<br />
and the deflection is measured as a function of applied<br />
pressure as shown in Fig. 3. In [3] and [4] the biaxial modulus<br />
and the residual stress of the film are extracted from the data<br />
using various mathematical models.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
e σ<br />
Ee 3<br />
P(w 0 ) = C 1 w + C<br />
2<br />
a<br />
0 2 ( ν ) w (7)<br />
( − ν ) 1a<br />
024<br />
where P is the applied uniform pressure, w 0 the deflection of<br />
the diaphragm midpoint, e the diaphragm thickness, a half of<br />
the diaphragm side length, E the Young’s modulus, ν the<br />
Poisson’s ratio and σ the residual or internal stress. The<br />
dimensionless constants C 1 and C 2 are numerical parameters<br />
which are obtained from the results of Maier-Schneider et al.<br />
[5] :<br />
C 1 = 3,45 and C 2 ( ν ) = 1,994(1-0,271 ν )/(1- ν ) (8)<br />
If the midpoint deflection w 0 is known, the deflection of<br />
the diaphragm from mid-side to mid-side can be calculated<br />
using [6-8] :<br />
w(y,0)=w 0 (1+0,401(y/a)) 2 cos(πy/2a) (9)<br />
Fig. 3. Deflection of a square diaphragm in response to an applied pressure.<br />
Due to the presence of residual stress and a significantly<br />
large deflection of the diaphragm compared to its thickness,<br />
the developed strain energy in the middle of the diaphragm<br />
causes a stretch of the diaphragm middle surface. The<br />
deflection of the diaphragm middle surface corresponds to a<br />
nonlinear behavior of a rigidly clamped diaphragm and is<br />
known as spring hardening. Thus, the analytical solution for<br />
diaphragm deflection from electrostatic forces must account<br />
for this spring hardening effect in addition to the nonlinear and<br />
non uniform electrostatic forces. Tabata et al. [3] developed an<br />
analytical solution for the load-deflection of membranes. They<br />
found a relationship between the external pressure load and<br />
the membrane deflection to determine the residual stress and<br />
Young’s modulus of thin films. Pan et al. [4] compared the<br />
analytical solution with FEM analysis. They found that the<br />
functional form of the analytical results is correct, but some<br />
constants have to be corrected. Pan et al. also found that the<br />
analytical forms of the membrane’s bending lines do not<br />
describe the real behavior very accurately. Maier-Schneider et<br />
al. [5] found an analytical solution for the load-deflection<br />
behavior of a membrane by minimization of the total potential<br />
energy. A new functional form of the membrane’s bending<br />
shape was found which agrees well with experimental<br />
measurements and with FEM analysis.<br />
Following the large deflection model, for a rigidly<br />
clamped square diaphragm with built-in residual stress, the<br />
load-deflection relationship of the midpoint of the diaphragm<br />
under a uniform pressure P can be expressed as [4-7] :<br />
The 2-D distributed problem is approximated by a rigid<br />
body suspended by a lumped spring. The spring constant has<br />
units of N/m and is defined as F elastic /w Max where w Max is the<br />
maximum displacement of the diaphragm with no<br />
electrostatic load, but with a uniform distributed pressure<br />
load P. In our case we have: w Max =w 0 , P=P(w 0 ) and<br />
F elastic =P(w 0 )A, so the nonlinear spring constant of the square<br />
diaphragm is :<br />
P(w0<br />
) A e σ<br />
Ee 2<br />
K nl = = (C 1 + C 2 ( ν ) w ) A (10)<br />
2<br />
w0<br />
a<br />
( − ν ) 1a<br />
The deflection-dependent nonlinearity due to spring<br />
hardening appears in equation (11) where the square of the<br />
midpoint deflection variable w 0 has been obtained. For a test<br />
device we consider the parameters given in table 1.<br />
TABLE I<br />
MODEL PARAMETERS<br />
parameter e a d 0 E ν σ<br />
024<br />
value 0,8 1,2 3,5 169 0,28 20<br />
unity μm μm μm GPa - MPa<br />
Spring hardening resulting from the deflection of a<br />
clamped square diaphragm midpoint due to an applied uniform<br />
pressure is plotted in Fig.4. From this figure we can obtain the<br />
value of the non-linear spring.<br />
Spring constant (N/m)<br />
260<br />
255<br />
250<br />
245<br />
240<br />
235<br />
230<br />
225<br />
220<br />
0 0.5 1 1.5 2 2.5 3<br />
Diaphragm deflexion (m)<br />
x 10 -6<br />
Fig. 4. Spring constant and diaphragm deflection<br />
189
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May 2011, Aix-en-Provence, France<br />
B. Pull-in voltage evaluation of a square diaphragm<br />
<br />
The analysis carried out in section 2 for a parallel plate<br />
capacitor structure can be extended to the case of a fully<br />
clamped square diaphragm separated from a rigid backplate by<br />
a small airgap. The deflection of the diaphragm is due to the<br />
resultant effect of electrostatic and restoring elastic forces (air<br />
damping force is neglected). For a parallel plate configuration<br />
(Fig. 2) the non linear electrostatic force is always uniform.<br />
However, for a rigidly clamped diaphragm, the electrostatic<br />
force becomes non-uniform due to a hemispherical<br />
deformation profile of the diaphragm. Thus, to evaluate the<br />
deflection of a rigidly clamped diaphragm under an<br />
electrostatic force, it is necessary to obtain a uniform linear<br />
model of the electrostatic force that can be applied in loaddeflection<br />
equation (7). A uniform linearized model of the<br />
electrostatic force can be obtained from (5) by linearizing the<br />
electrostatic force about the zero deflection point x 0 = 0. The<br />
linearized electrostatic force is :<br />
⎛<br />
2 1 x<br />
F elect. = ε<br />
0<br />
V<br />
⎜<br />
A (11)<br />
2<br />
2d d<br />
⎝<br />
0<br />
+ 3<br />
0<br />
and the effective linearized uniform electrostatic pressure on<br />
the diaphragm is :<br />
P elect. =<br />
F<br />
elect.<br />
= A<br />
⎛<br />
⎜<br />
⎝<br />
⎟ ⎞<br />
⎠<br />
+<br />
2 3<br />
0<br />
d0<br />
⎟ ⎞<br />
⎠<br />
2 1 x<br />
ε<br />
0<br />
V<br />
(12)<br />
2d<br />
This equation is general and can be applied to any sort of<br />
diaphragms. We deduce the pull-in electrostatic pressure by<br />
replacing x by the pull-in deflection d 0 /3:<br />
P PI-elect. =<br />
5 ε V<br />
2<br />
P I0<br />
2<br />
0<br />
d6<br />
(13)<br />
where V PI represents the desired pull-in voltage. The applied<br />
uniform transverse pressure load of the rigidly clamped square<br />
diaphragm built in residual stress, obtained from elastic<br />
considerations (equation 7) equals the effective linearized<br />
uniform electrostatic pressure (equation 13) and at the distance<br />
where the pull-in occurs we obtain :<br />
e σ d<br />
C<br />
0 1<br />
2<br />
a 3<br />
+ C 2 ( ν )<br />
⎛<br />
0<br />
24⎟ ⎟ ⎞<br />
⎠<br />
Ee d 5 ε V<br />
2<br />
⎜ =<br />
3<br />
2<br />
( − ν ) 1a<br />
d6<br />
⎝<br />
3<br />
0<br />
P I0<br />
(14)<br />
The above equation is solved and we obtain the expression<br />
for the pull-in voltage for a clamped square diaphragm under<br />
an electrostatic pressure:<br />
V PI =<br />
d<br />
0<br />
a<br />
6<br />
5 ε0<br />
⎡<br />
⎢<br />
⎢<br />
⎢⎣<br />
1<br />
3<br />
d<br />
⎤<br />
C ⎛ d ⎞<br />
20<br />
ν Ee)(<br />
⎜<br />
σeC +<br />
⎟ ⎥<br />
(15)<br />
3<br />
3<br />
⎥<br />
( − ν ) 1a ⎝ ⎠ ⎥⎦<br />
22<br />
The pull-in voltage can be calculated using equation (15).<br />
With the parameters given in table 1 we obtain a value of<br />
17.45 volts. If we use equation (3), which corresponds to the<br />
ideal case of two parallel plates, we obtain V pi = 15.02 volts, a<br />
value which is 2.43 volts smaller than the value obtained with<br />
the method proposed in this communication. The mid-side to<br />
mid-side deflection profiles of the clamped diaphragm for<br />
different voltages is shown in Fig.5.<br />
Deflection (meter)<br />
10-6 -0.2<br />
-0.4<br />
-0.6<br />
-0.8<br />
-1<br />
10 Volts<br />
15 Volts<br />
17,45 Volts<br />
-1.2<br />
-6 -4 -2 0 2 4 6<br />
0 x Diaphragm side length (meter)<br />
x 10 -4<br />
Fig. 5. Diaphragm deflection for different bias voltage<br />
IV.<br />
CIRCULAR DIAPHRAGM ANALYSIS<br />
A. Load-deflection characteristics of a circular diaphragm<br />
Consider a circular plate of radius a and constant<br />
thickness e under a uniform transverse load p z = p 0 and an<br />
initial tension load N r = N 0 , as shown in Fig.6.<br />
Fig.6. Schematic of a clamped circular plate under an<br />
initial in-plane stress N 0<br />
Based on von Karman plate theory [6] for large circular<br />
plate deflections, the equilibrium equations for the<br />
symmetrical bending of this plate are :<br />
( r N ,r ) ,r - N θ =0 (16)<br />
( r Q r ) ,r + (r N r w ,r ) ,r + r p 0 = 0 (17)<br />
r Q r = (r M r ) ,r -<br />
M θ<br />
(18)<br />
where ( ), r indicates the differentiation with respect to the<br />
radial coordinate r, w is the normal displacement or the<br />
deflection of the plate in the z-direction, N r , N<br />
θ are the lateral<br />
loads, Q r is the shear force, M r and M θ are the bending<br />
moments. The shear force can be obtained by integrating (17)<br />
Q r + N r w ,r +<br />
1<br />
p0 r = 0 (19)<br />
2<br />
190
Using the radial and tangential midplane strains and<br />
curvatures [6] the second version of the shear force is :<br />
Q r = - D ( w ,rrr +<br />
r<br />
1<br />
w,rr -<br />
2<br />
r<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
1<br />
w, r ) (20)<br />
where D=Ee 3 /12(1- ν 2 ) is the flexural rigidity of the plate.<br />
Placing (20) into (19) produces:<br />
membrane behavior is revealed due to the nonlinearity of<br />
W(0)/P. In this Fig.7, there appears to be a transition from pure<br />
plate behavior to pure membrane behavior in the region from<br />
k ≈ 1 to k ≈ 20. For k20 a membrane behavior dominates the majority of the<br />
clamped circular plate.<br />
plate behavior<br />
w ,rrr +<br />
r<br />
1<br />
w,rr -<br />
1<br />
w, r -<br />
2<br />
r<br />
N r w,r =<br />
D<br />
p r<br />
0<br />
(21)<br />
D2<br />
W(0)/P<br />
10 0 k<br />
10 -1<br />
10 -2<br />
membrane<br />
behavior<br />
The analytical solution of (21) is given by [9] :<br />
10 -3<br />
6Pe(1 - ν<br />
w(r) =<br />
2<br />
k<br />
2<br />
)<br />
[<br />
( I ( kr/a ) - I (k) )<br />
0<br />
0 −<br />
+ ] (22)<br />
Ik (k)<br />
a2<br />
where I 0 ( ) and I 1 ( ) are the modified Bessel functions of the<br />
first kind [6], P is the nondimensional loading parameter and<br />
k is the nondimensional tension parameter. These two<br />
parameters are defined as:<br />
1<br />
10 -4<br />
10 -1 10 0 10 1 10 2<br />
22<br />
ra Fig.7 Center deflection normalized by the loading parameter as a function of<br />
2<br />
the tension parameter<br />
In order to have a thorough insight to the effects of initial<br />
tension upon the related geometrical responses, normalized<br />
deflection shapes are plotted in Fig.8. As the tension<br />
parameter k increases a sharp change in the curvature near the<br />
edge appears, in order to accommodate the zero slope<br />
boundary condition.<br />
P =<br />
0<br />
4<br />
ap<br />
; k = a<br />
4<br />
eE<br />
N 0<br />
=<br />
D<br />
a<br />
e<br />
2<br />
N012 (1 − ν<br />
e E<br />
)<br />
(23)<br />
Two limiting cases are of interest here: the pure plate case<br />
where the tension parameter has the value k=0 and the pure<br />
membrane case where k ∞→ . For the pure plate case, taking<br />
the small argument limit of the modified Bessel functions, the<br />
corresponding deflection is:<br />
W/W(o)<br />
1<br />
0.9<br />
0.8<br />
0.7<br />
0.6<br />
0.5<br />
0.4<br />
0.3<br />
0.2<br />
pure membrane<br />
0.1<br />
k=10<br />
pure plate<br />
0<br />
-1 -0.8 -0.6 -0.4 -0.2 0 0.2 0.4 0.6 0.8 1<br />
r/a<br />
w(r) =<br />
16<br />
3 e P (1- ν 2 ) (1-<br />
r 2 )()<br />
2<br />
a<br />
(24)<br />
For the pure membrane case, taking the large argument<br />
limit of the modified Bessel functions, the pure membrane<br />
deflection is:<br />
e3 P (1- ν 2 r 2<br />
) ( 1- )() (25)<br />
w(r )=<br />
2<br />
k<br />
In Fig.7 the center deflection of the clamped circular plate,<br />
normalized by the transverse load, is plotted against the initial<br />
nondimensional tension parameter k. The effect of initial inplane<br />
tension (or the lateral loads effect) is shown in this<br />
figure where two asymptotes are present. The horizontal part<br />
of the curve means the center deflection is in a linear<br />
proportion to the applied transverse load (W(0) ∝ P , see also<br />
(24)) and is not a function of the initial tension, thus it reflects<br />
a plate behavior. The inclined straight line indicates a<br />
nonlinear variation of the center deflection with the tension<br />
parameter k (W(0)/P ∝ 1/k 2 , see also (36)), hence, a<br />
a<br />
Fig. 8. Normalized deflection as a function of normalized radial distance<br />
B. Pull-in voltage evaluation of a circular diaphragm<br />
From equation (22) we deduce the uniform transverse<br />
pressure which is applied at the center of the clamped circular<br />
plate:<br />
p 0 =<br />
D2 2 1<br />
k w(0) ⎡1<br />
I (k) ⎤<br />
−<br />
0<br />
⎢ − ⎥<br />
2 k I (k)<br />
4<br />
a<br />
⎢⎣<br />
1<br />
⎥⎦<br />
(26)<br />
As shown in paragraph II, we assume that the pull-in effect<br />
occurs when the deflection of the movable plate is one-third of<br />
the original air gap d 0 and the uniform pressure load given by<br />
(26) is equal to the pull-in electrostatic pression given by (13).<br />
We obtain:<br />
⎛<br />
⎜<br />
⎝<br />
2<br />
kD2 d 0<br />
4<br />
a 3<br />
⎞ ⎡<br />
⎟<br />
⎢<br />
⎠ ⎢⎣<br />
1<br />
I (k)<br />
0<br />
−<br />
2 k I (k)<br />
1<br />
⎤<br />
⎥<br />
⎥⎦<br />
− 1<br />
2<br />
5 ε0<br />
V<br />
PI<br />
=<br />
2<br />
d6 0<br />
(27)<br />
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The above equation is solved and we obtain the<br />
<br />
24<br />
expression for the pull-in voltage for a rigidly clamped<br />
22<br />
circular plate:<br />
20<br />
pure membrane<br />
plate<br />
V PI =<br />
dk2 dD I (k)<br />
0 0 1 0<br />
2 ⎢ −<br />
a 5 ε 2 k I (k)<br />
0 ⎢ 1<br />
⎡<br />
⎣<br />
⎤<br />
⎥<br />
⎥⎦<br />
− 1<br />
(28)<br />
In the case of a pure circular plate the uniform transverse<br />
pressure is obtained from (24):<br />
p 0 =<br />
4<br />
a<br />
D64 w(0)<br />
(29)<br />
and at one-third of the original air gap we obtain the pull-in<br />
voltage for a pure clamped circular plate:<br />
2<br />
D d 6 4 5 ε0<br />
V<br />
0<br />
PI<br />
=<br />
4<br />
a 3<br />
2<br />
⇒ V PI =<br />
d6 0<br />
⎛<br />
⎜<br />
⎝<br />
⎞<br />
⎟<br />
⎠<br />
d8 0<br />
2<br />
a<br />
dD2<br />
0<br />
5 ε<br />
0<br />
(30)<br />
In the case of a pure circular membrane the uniform<br />
transverse pressure is obtained from (25):<br />
2<br />
kD4 w(0)<br />
p 0 =<br />
4<br />
a<br />
(31)<br />
and at one-third of the original air gap we obtain the pull-in<br />
voltage for a pure clamped circular membrane:<br />
a<br />
4<br />
2<br />
⎛<br />
⎜<br />
⎝<br />
2<br />
k d D 5 ε<br />
0<br />
4 0 V<br />
PI<br />
dk dD8<br />
0<br />
0<br />
=<br />
3<br />
2<br />
⇒ V PI = 2<br />
(32)<br />
d6 0<br />
a 5 ε<br />
0<br />
⎞<br />
⎟<br />
⎠<br />
To illustrate the above model of pull-in evaluation, a<br />
clamped circular diaphragm of Young’s modulus E = 169 GPa<br />
and Poisson’s ratio ν = 0,28 is considered. The thickness of the<br />
plate is e = 0,8 μ m and the airgap thickness is d 0 = 3,5 μ m.<br />
The permittivity in free space is ε 0 = 8,5x10 -12 F.m -1 and the<br />
residual in-plane stress is σ 0 =N 0 /e = 20 MPa. Fig. 7 shows<br />
the pull-in voltage for a plate having the previously device<br />
parameters and for a pure membrane as a function of radius. It<br />
is evident that there is negligible difference between the pullin<br />
voltage evaluated using a pure membrane model and given<br />
by (32) and the pull-in voltage our circular clamped plate.<br />
Pull-in voltage (V)<br />
18<br />
16<br />
14<br />
12<br />
10<br />
8<br />
6<br />
4<br />
0.5 1 1.5 2 2.5<br />
Diaphragm radius(m)<br />
x 10 -3<br />
Fig. 9. Pull-in voltage for a pure circular membrane and for a circular plate<br />
V. CONCLUSION<br />
The deflections of clamped square and circular plates<br />
under a uniform transverse load and in-plane tension loads<br />
have been studied in the communication. In particular the<br />
transition from plate behavior to membrane behavior has been<br />
described for a clamped circular plate. A new relatively simple<br />
closed-form model to evaluate the pull-in voltage associated<br />
with rigidly clamped diaphragms subject to an electrostatic<br />
force has been presented and numerical results have been<br />
obtained showing the effectiveness of the method in pull-in<br />
voltage evaluation.<br />
REFERENCES<br />
[1] S. D. Senturia, Microsystem Design. Springer, 2000.<br />
[2] P.O. Osterberg and S.D. Senturia, ’’M-Test: a test chip for MEMS<br />
material property measurements using electrostatically actuated<br />
test structures’’, J. of Microelectromechanical Systems, Vol. 6,<br />
pp.107-117, 1997.<br />
[3] O. Tabata, K. Kawahata, S. Suguyama and I. Igarashi,<br />
’’Mechanical property measurements of thin films using loaddeflection<br />
of composite rectangular membranes’’, Sensors and<br />
Actuators, Vol. 20, pp. 135-141, 1989.<br />
[4] J.Y. Pan, P. Lin, F. Maseeh, S.D. Senturia, ’’Verification of FEM<br />
analysis of load-deflection methods for measuring mechanical<br />
properties of thin films’’, IEEE Solid-State Sensors and Actuators<br />
Worshop, Hilton Head Island, pp. 70-73, 1990.<br />
[5] D. Maier-Schneider, J. Maibach and E. Obermeier, ’’A new<br />
analytical solution for the load-deflection of square membranes’’,<br />
J. of Microelectromechanical Systems, Vol. 4, pp. 238-241, 1995.<br />
[6] S. Timoshenko, Theory of Plates and Shells. Mc Graw Hill, 1959.<br />
[7] S. Chowdhury, M. Ahmadi and W.C. Miller, ’’Nonlinear effects in<br />
MEMS capacitive microphone design’’, International Conference<br />
on MEMS, NANO and Smart Systems 03, Banff, Alberta, Canada,<br />
2003.<br />
[8] J. Lardiès, O. Arbey and M. Berthillier, ’’Analysis of the pull-in<br />
voltage in capacitive mechanical sensors’’, Third International<br />
Conference on Multidisciplinary Design Optimization and<br />
Applications, 21-23 June 2010, Paris.<br />
[9] M Sheplak and J. Dugundji, ’’Large deflections of clamped circular<br />
plates under initial tension’’, J. of Applied Mechanics, Vol. 65,<br />
pp.107-115, 1998.<br />
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<br />
Optimisation and realisation of a portable NMR<br />
apparatus and Micro Antenna for NMR<br />
Patrick Poulichet 1 , Latifa Fakri-Bouchet 2 , Christophe Delabie 1 , Lionel Rousseau 1 , Abdenasser Fakri 1 and Anne Exertier 1<br />
ESIEE, 2 BD Blaise Pascal, 93162 Noisy Le Grand, France 1<br />
Laboratoire CREATIS - LRMN UMR CNRS 5220, INSERM U630, INSA de Lyon 3, rue Victor-Grignard,<br />
69616 Villeurbanne, France 2<br />
p.poulichet@esiee.fr (+33) 1.45.92.67.18<br />
Latifa.Fakri-Bouchet@creatis.univ-lyon1.fr (+33) 4.72.44.82.08<br />
Abstract- This paper is focused on two designs and<br />
realizations. The first one concerns a prototype of a portable<br />
NMR (nuclear magnetic resonance) apparatus. The second one<br />
concerns NMR micro antenna realization.<br />
For the first part, our goal is the NMR magnetic field<br />
homogeneity and the signal-to-noise ratio (SNR) improvement.<br />
Since de the volume of the sample to analyse is around 1 cm 3 ,<br />
the design is optimized to obtain a good SNR. Particularly, the<br />
magnet is chosen to obtain a high magnetic field with limited<br />
inhomogeneities. The receiver antenna is designed and<br />
optimized to have high feeling factor and then more sensitivity.<br />
A mixer and a low-pass filter are used in order to limit the<br />
bandwidth and reduce the thermal noise. The FID is digitized<br />
and addressed to a FPGA which averages successive<br />
acquisitions in order to increase the SNR. The final acquisition<br />
is processed for determining the FID spectrum.<br />
In the second part, a new concept of micro coil is presented<br />
in order to measure the small volumes and small<br />
concentrations samples by NMR spectroscopy at 4.7 T (200<br />
MHz proton frequency resonance).<br />
This micro sensor would offer the possibility of new<br />
investigation techniques based on micro coils’ implantation<br />
used for in vivo study of local cerebral metabolites of animals<br />
models.<br />
Keyword: NMR, single side, portable, magnet, micro<br />
coil.<br />
I. INTRODUCTION<br />
The NMR signal measurement requires a static magnetic<br />
field B0 and a pulsed frequency varying B1 magnetic field.<br />
The studied sample is submitted to these magnetic fields.<br />
Then, the B1 magnetic field is applied at a Larmor<br />
frequency (1T correspond to 42.57 MHz), is applied, the<br />
signal FID (Free Induction Decay) generated by<br />
magnetization motion is acquired using a detection coil.<br />
In mobile NMR, B0 is generated with an arrangement of<br />
magnets in order to deliver a homogeneous magnetic field.<br />
For measurement of the relaxation (T2 or T1) or a spectrum,<br />
inhomogeneities must be as low as possible particularly if<br />
the objective of the measurement is spectroscopy (ΔB0/B0<br />
≅ 10 ppm).<br />
Reference [4] is a review of the different ways of the array<br />
magnets setup in order to obtain the best B0 homogeneity.<br />
Reference [1] reports the construction of a mobile<br />
tomography device by exploiting the concept of movable<br />
permanent magnets in the shim unit of a Halbach array. The<br />
cross section of a banana placed inside the magnets is<br />
presented. Reference [2] proposes a complete procedure for<br />
permanent magnet design, fabrication, characterization and<br />
shimming. 1H NMR spectrum of a 3 mm 3 sample of water<br />
doped with CuSO4 in the shimmed magnet is presented.<br />
The half-height full width (HHFW) is about 12 ppm.<br />
Reference [3] presents a single-sided mobile NMR<br />
apparatus with a small Halbach magnet. It is lightweight,<br />
compact and exhibits good sensitivity.<br />
Reference [5] the spectrometer design that uses an FPGA.<br />
The system is composed of an FPGA chip and several<br />
peripheral boards for USB communication, direct-digital<br />
synthesis (DDS), RF transmission, signal acquisition, etc. In<br />
the FPGA have been implemented a number of digital<br />
modules including three pulse programmers, the digital part<br />
of DDS, a digital quadrature demodulator, dual digital lowpass<br />
filters, and a PC interface have been implemented.<br />
The feasibility to use a new generation of microcoils was<br />
proposed in a recent study [6]. It demonstrated potential<br />
opportunities in terms of increased signal-to-noise ratio<br />
(SNR), spatial resolution, and limits of detection (LOD) [7]<br />
compared to the surface-coil [8]. Their use for localized<br />
spectroscopic studies of NMR observable cerebral<br />
metabolites into 2mm 3 region of interest (ROI), aims to<br />
push limits of in vivo detection.<br />
II. MOBILE NMR<br />
In our design, we used the mains parts describe in [5] in<br />
order to realize portable NMR apparatus dedicated to<br />
portable NMR. The aim is to optimize the SNR. The magnet<br />
(alnico) uses two large piece of steel in order to reduce<br />
inhomogeneities of the magnetic field.<br />
Fig. 1 shows the functional schematic of the spectrometer<br />
connected to the PA (Power Amplifier) and connected to<br />
193
the LNA (Low Noise Amplifier). The duplexer use a circuit<br />
with two diodes for ensuring a good isolation when a signal<br />
is applied to the generation coil.<br />
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<br />
using USB protocol. There are some daughter cards<br />
connected to the FPGA as shown in Fig. 2. The FPGA also<br />
ensures that the delivery signal parameters are correct.<br />
Fig. 1. Functional schematic of the connexions between the<br />
spectrometer and the coils<br />
A. Optimisation of the SNR<br />
Like the volume of the sample to analyze is around 4,6 cm 3 ,<br />
we choose the dimension of the air gap, the wire diameter,<br />
the noise factor of the preamplifier in order to deliver a<br />
signal above the noise. The measured induction B 0 is 0,116<br />
T. With the constant gyromagnetique γ, we determine the<br />
Larmor frequency :<br />
f 1<br />
0<br />
= γ .<br />
0<br />
4.92<br />
2π<br />
B = MHz<br />
With the value of the Plank constant h and the Bolzman<br />
constant k, we obtain the magnetic moment M0 :<br />
B<br />
M0 = N. γ .3. = 9,9.10<br />
4. kT .<br />
2 2 0<br />
−6<br />
The FID value is : ξ = K. ω0. BM<br />
1. 0.<br />
V when K is an<br />
homogeneity factor, B 1 the alternative magnetic field and V<br />
the volume of the sample. In order to increase the B 1<br />
magnetic value, a separate inductance from the detection<br />
coil was chosen. With a current of nearly 5 A in the<br />
generation inductance, ξ = 14 µV. By taking into account<br />
the bandwidth of the detection coil Δf = 500 kHz, the<br />
spectral density of noise across this inductance is calculated:<br />
e n = 1,6.10 -10 V. With the noise factor of the preamplifier,<br />
the SNR is calculated: SNR ≅ 100. In this calculus, the<br />
noise is underestimated because only the thermal noise from<br />
the detection coil and the noise from the LNA are take into<br />
account.<br />
B. Spectrometer<br />
Fig. 2 shows the functional schematic and the realization<br />
of the spectrometer used to control FID generation,<br />
acquisition and processing. A Cyclone III FPGA is used to<br />
control the FID record, acquisition time, sequence repetition<br />
time used for NMR, FID storage and transfer to the PC<br />
Fig. 2. Functional schematic and realization of the<br />
spectrometer<br />
To digitize the FID, an ADC 14 bits - 65 MIPS is used.<br />
Then data are fed into the FPGA memory. After M samples<br />
of N points, the digital word is sent to the PC via USB port.<br />
When the acquisition parameters are defined, the<br />
generated signal is sent by the DDS and an amplitude<br />
modulation is applied if necessary via the DAC. The<br />
amplitude and frequency modulation can be used to<br />
generate complex shapes of chirps in order to compensate<br />
the inhomogeneities of B0.<br />
Fig. 3. Magnet and antenna connected to the duplexer<br />
The magnet and the duplexer of the Fig. 3 were connected<br />
to the LNA and the PA.<br />
Using two large pieces of steel, low inhomogeneities are<br />
expected. The sample of liquid of the Fig. 3 is placed inside<br />
the air gap of the magnet.<br />
Fig. 4 shows an acquired signal with a superposed noise<br />
and the same signal with noise removed. This last signal is<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
194
11-13 <br />
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generated by the treatment operated in the FPGA.<br />
<br />
good aspect and good uniformity of the deposited layer.<br />
Fig. 6. Micro coil on the right and the PCB used to tune and<br />
match the signal<br />
Fig. 4. Signal generated with and without the noise<br />
The process to deposite the thick resist is difficult to<br />
optimise because the thickness is around 50 µm. The resist<br />
have to be etched for 50 µm on a width of 20 µm. Thus, at the<br />
bottom of the hole, if there are some rests of the resist, the<br />
adherence of the cooper will be poor. It’s this problem that we<br />
can see on the Fig. 7; there a lack of cooper inside the cercle.<br />
Fig. 5 shows the shape of the signal generated from the<br />
spectrometer. We choose to deliver a sequence CPMG of three<br />
pulses [π/2 π π] separated by [t 2 t 4 t 6 ]. Each times are<br />
programmed in the VHDL code of the FPGA.<br />
Fig. 7 : Photography of the microcoil NMR<br />
Fig. 5 : Shape of the generated signal from the spectrometer<br />
The pattern represents in the Fig. 8 is used to verify that<br />
the resistivity of the layer copper deposited is correct. The<br />
thickness is 30 µm, the width is 42 µm and the total length<br />
of the pattern is about 10.9 mm. Thus resistivity calculation<br />
−8<br />
gives ρ = 2,88.10 Ω .m . After bake at 150 °C during 30<br />
−8<br />
minutes, the resistivity is ρ = 2.10 Ω .m . It is not far from<br />
−8<br />
the theoretical copper resistivity ρ = 1,72.10 Ω .m .<br />
II.<br />
MICROCOIL NMR<br />
NMR micro coil represented in Fig. 6 detects the FID and<br />
is connected to the PCB for tuning and matching circuit. It<br />
is connected to the PCB for tuning and matching. Then, the<br />
signal is fed to an external amplifier. To improve the SNR,<br />
the resistance of the micro coil has to be very low.<br />
Therefore, the thickness of the copper used to realize the<br />
coil has to be thick so usually it uses electroplating. We<br />
optimize the electroplating process: flow of the electrolyte,<br />
current density, additive in the electrolyte in order to obtain<br />
Fig. 8. Shape used to determine the resistivity of the<br />
electroplate copper<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
195
III.<br />
CONCLUSION<br />
Two important parts necessary for NMR measurement are<br />
described in this paper. In the first part, a portable NMR<br />
apparatus is described. In the future, the array of magnets<br />
will be modified to reduce the inhomogeneities of B0. The<br />
running works concern also the achievement of PC<br />
supervisor linked to the Portable NMR divice.<br />
In the second part, the realization of a micro coil is<br />
described.<br />
The two parts are designed to constitute an original NMR<br />
portable system for the analysis of low volumes samples.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
A special thanks to the folowing students that worked on<br />
one part of this projet: LIMA MIGLIORINI Fabricio,<br />
ZHANG Shu and ZHANG Hao.<br />
REFERENCES<br />
[1] E. Danieli, J. Mauler, J. Perlo, B. Blümich, F. Casanova, “Mobile<br />
sensor for high resolution NMR spectroscopy and imaging”<br />
Journal of Magnetic Resonance 198, 2009, pp 80-87.<br />
[2] C. Hugon, F. D’Amico, G. Aubert, D. Sakellariou, “Design of<br />
arbitrarily homogeneous permanent magnet systems for NMR<br />
and MRI: Theory and experimental developments of a simple<br />
portable magnet“, Journal of Magnetic Resonance 205, 2010,<br />
pp75-85.<br />
[3] W-H Changa, J-H Chena, L-P Hwang “Single-sided mobile NMR<br />
with a Halbach magnet” in Magnetic Resonance Imaging 24, 2006,<br />
pp 1095–1102.<br />
[4] V. DEMAS, P. J. PRADO, “Compact Magnets for Magnetic<br />
Resonance”, Concepts in Magnetic Resonance PartA, Vol. 34A(1),<br />
2009, pp 48–59.<br />
[5] K. Takeda, “OPENCORE NMR: Open-source core modules for<br />
implementing an integrated FPGA-based NMR spectrometer”,<br />
Journal of Magnetic Resonance 192, 2008, pp 218–229.<br />
[6] Baxan et al, C.R.Chim.2007.<br />
[7] Lacey et al, Chem.Rev.1999.<br />
[8] Kadjo , et al, ESMRMB, Valencia, 2008.<br />
BIOGRAPHIE<br />
Patrick Poulichet received his degree of electrical engineer at<br />
CNAM-Paris in 1998. He received the Ph.D. degree in Electrical<br />
Engineering from the École Normale Supérieure de Cachan (SATIE<br />
CNRS UMR 8029) in 2001. Since 1995, he has been with the<br />
Department of Electrical Engineering, at ESIEE (Ecole Supérieure<br />
d’Ingénieurs en Electronique et Electrotechnique) in France, as an<br />
Associate Professor. His research concerns portable NMR, integrated<br />
electronic and MEMS, and EMC.<br />
Latifa Fakri-Bouchet received her Ph. D degree in the field of<br />
biomedical engineering from Lyon 1 university in 1996. Lastly she<br />
obtained the “habilitation” for Research Heading (HDR, 2008). Since<br />
1995, she was assistant Professor (ATER), and then Associate<br />
Professor at Institut Universitaire de Technologie (IUT) of Lyon, and<br />
Laboratory CREATIS-LRMN UMR CNRS 5220, U630 INSERM,<br />
Université Lyon1 Claude Bernard, INSA de Lyon. Her main interest<br />
is in development of electronic circuits, and instrumentation<br />
dedicated to NMR biomedical applications, more particularly coils<br />
and micro coils design, and NMR instrumentation potentially leading<br />
to developments for clinical purposes and industrial applications.<br />
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<br />
Convex Corner Compensation for a Compact Seismic<br />
Mass with High Aspect Ratio Using Anisotropic Wet<br />
Etching of (100) Silicon<br />
Jyh-Cheng YU<br />
National Kaohsiung First University of Science and Technology<br />
2 Jhuoyue Rd.,Nanzih , Kaohsiung City 811,Taiwan, R.O.C.<br />
Abstract - This paper reports a novel convex corner<br />
compensation design for a compact seismic mass of high aspect<br />
ratio using KOH etching of (100) silicon. Anisotropic wet<br />
etching is often applied to fabricate a seismic mass due to cost<br />
advantage. Dimension of the convex corner compensation<br />
pattern is in proportional to etching depth, which restrain the<br />
miniature of seismic mass and supporting beams. If the width<br />
of the seismic is too small, overlap of compensation pattern<br />
occurs to cause a compensation failure. This study presents a<br />
corner compensation for a mesa with high aspect ratio using<br />
oriented compensating bands augmented with a<br />
mandatory separation and a -oriented beam to the<br />
truncated bands due to the overlap of adjacent<br />
compensation patterns. An empirical equation is presented from<br />
the simulation of anisotropic etching using Intellisuite. The<br />
design can produce a satisfactory wet etching mesa with the<br />
aspect ratio of 0.6, while a typical oriented<br />
compensating bands can only applied to a mesa with a largest<br />
aspect ratio of 0.35. A mesa is etched using 30%, 80°C KOH to<br />
verify the design feasibility.<br />
I. INTRODUCTION<br />
Mesa with thin supported beams is a typical configuration<br />
for inertia sensors [2][3][4] as shown in Figure 1. Wet bulk<br />
micromachining processes of (100) silicon such as KOH and<br />
TMAH are widely used to the fabrication of the seismic mesa<br />
due to cost advantages in comparison with dry etching [1].<br />
However, corner compensation pattern is required in the mask<br />
design for protrusion corners to prevent undercut in wet<br />
chemical etching as shown in Figure 2.<br />
Various compensation patterns were proposed to such as<br />
simple squares, triangles, oriented bands, <br />
oriented band with narrow beams, etc. [5][9]. The<br />
compensation of simple oriented bands is widely<br />
applied because perfect shaped convex corners can be produced.<br />
The width of the band is twice the etching depth and the length<br />
of the band depends on the etch rate of the {411} plane relative<br />
to the {100}. Typical length for KOH etching is about 3.2 times<br />
of etching depth as shown in Figure 3 [6].<br />
Figure 1 Typical configuration of a inertial sensor<br />
{111}<br />
{110}<br />
{001}<br />
151.9°<br />
{411} {411}<br />
{111}<br />
{110}<br />
Figure 2 Appearance of fast etching planes at the convex corner during<br />
anisotropic etching<br />
Figure 3 Dimensions and successive etched shape of the oriented<br />
compensating beam [6]<br />
Since the size of the compensation pattern is in proportional<br />
to etching depth, which refrains from further miniature of the<br />
devices with convex corners. If a typical band<br />
compensation is used, the constraints among the size of the<br />
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11-13 May 2011, Aix-en-Provence, France<br />
<br />
seismic mass and the length of the supported beams in Figure<br />
3 and Figure 4(a) are listed in (3) to (4).<br />
W = 2H (1)<br />
L =3. 2H (2)<br />
m<br />
> 23 HW<br />
(3)<br />
W3 > 21.1<br />
H<br />
(4)<br />
Some other comer compensations suggested a oriented<br />
band and fan-like [110]-oriented side beams [6] and a<br />
modified band design [7] to reduce the required groove<br />
width, W 3 for the compensation pattern. However, if the edge<br />
width for the corner compensation is limited, such as a small<br />
seismic mass with large etching depth, compensation designs<br />
fail due to the overlap of neighboring patterns as shown in<br />
Figure 4(b). If overlap of compensation occurs, the etched<br />
mesa dimensional will be incorrect and convex corners will<br />
undercut.<br />
Wn = 2*H - 0.0007C 2 - 0.02C (5)<br />
Ln= 3.2*H + 0.0101C 2 - 3.66C (6)<br />
B = 1.91C (7)<br />
where C is the overlap of adjacent conventional -oriented<br />
bands as shown in Figure 4(b).<br />
(a)<br />
Figure 4 Overlap of adjacent -oriented bands due to limited width of<br />
square<br />
This study presents a corner compensation for a mesa with<br />
high aspect ratio using oriented compensating bands<br />
augmented with a mandatory separation and a -oriented<br />
beam to the truncated bands due to the overlap of<br />
adjacent compensation patterns. An empirical equation is<br />
presented from the simulation of anisotropic etching using<br />
Intellisuite.<br />
II. METHODOLOGY<br />
In the wet etching of seismic mass, conventional oriented<br />
bands of corner compensation is good if the ratio<br />
between the etching depth and the width of the mesa top is<br />
smaller than 0.35. If the aspect ratio (H /W m ) is larger than 0.35,<br />
the adjacent compensation patterns overlap. Here a modified<br />
design is presented as shown in Figure 5. The overlap bands are<br />
truncated with a separation of 20 (μm). A [110]-oriented beam<br />
is introduced to the truncated pattern. The width and the length<br />
of the band and beam are related to the etching depth (H). An<br />
empirical formula for KOH etching is presented from various<br />
simulations using Intellisuite and least squared curve fitting as<br />
shown in (5)~(7).<br />
(b)<br />
Figure 5 A modified -oriented bands for limited width of square.<br />
If the aspect ratio of mesa is smaller than 0.35, a<br />
conventional oriented band compensation will perfect<br />
convex corners. If a small mesa is fabricated using deep bulk<br />
wet etching, the proposed design provides a satisfactory convex<br />
corner for the seismic mass with aspect ratio between 0.35 and<br />
0.625. Several compensation design based on the proposed<br />
method are shown in Figure 6 to Figure 8. With the increase of<br />
aspect ratio, minor corner undercut presents. If the aspect ratio<br />
is larger than 0.625, corner undercut will still aggravate.<br />
The verification result for a mesa with width of 400 (μm)<br />
and etching depth 250 (μm) is shown in Figure 9. The etchant<br />
is 30% 80°C KOH. The corresponding aspect ratio of the mesa<br />
is 0.675. Only minor undercut can be observed in convex<br />
corners using the proposed compensation design, while typical<br />
compensation using simple -oriented bands fail<br />
completely for this case.<br />
198
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May 2011, Aix-en-Provence, France<br />
<br />
Figure 6 Simulation result for the proposed compensation design of a seismic<br />
mass of 400 (μm) width and aspect ratio of 0.38<br />
III. CONCLUSION<br />
This paper reports a novel convex corner compensation<br />
design for a compact seismic mass of high aspect ratio using<br />
KOH etching of (100) silicon. An empirical equation is<br />
presented from the simulation of anisotropic etching using<br />
Intellisuite. The design can produce a satisfactory wet etching<br />
mesa with the aspect ratio up to 0.625 compared to a typical<br />
oriented compensating bands can only applied to a mesa<br />
with a largest aspect ratio of 0.35. The verification experiment<br />
for a mesa is etched using 30% 80°C KOH verifies the design<br />
feasibility.<br />
Keywords: KOH wet etching, fabrication of seismic mass,<br />
silicon etching, convex corner compensation, Intellisuite<br />
Figure 7 Simulation result for the proposed compensation design of a seismic<br />
mass of 400 (μm) width and aspect ratio of 0.57<br />
Figure 8 Simulation result for the proposed compensation design of a seismic<br />
mass of 400 (μm) width and aspect ratio of 0.625<br />
REFERENCE<br />
[1]. Marc, J. Madou., Fundamentals of microfabrication: the<br />
science of miniaturization, 2 nd Ed.,(2002)<br />
[2]. Chen, H., Shen, S., and Bao, M., “Over-range capacity of<br />
a piezoresistive microaccelerometer”, Sensor and<br />
Actuator A., Vol.58, No. 3, (1997), pp. 197-201<br />
[3]. Yu, J. and Lai., F.H., “Design And Fabrication Of<br />
Microaccelerometers Using Piezoelectric Thin Films”,<br />
Ferroelectrics., Vol.263, (2001), pp.101-106.<br />
[4]. Yu J., Lee C., Chang C., Kuo W., Chang C.: Modeling<br />
Analysis of a Tri-Axial Microaccelerometer with<br />
Piezoelectric Thin-Film Sensing Using Energy Method,<br />
paper accepted, to appear in Journal of Microsystem<br />
Technologies.<br />
[5]. Lang, Walter., “Silicon Microstructuring Technology”,<br />
Materials Science and Engineering, R17, 1996, pp. 1-55.<br />
[6]. Mayer, G K, Offereins, H L, Sandmaier, H and Kuhl, K,<br />
“Fabrication of non-underetched convex corners in<br />
anisotropic etching of (100) silicon in aqueous KOH with<br />
respect to novel micromechanic elements,” J.<br />
Electrochem. Soc., (1990) 137, 3947–3951.<br />
[7]. Zhang, Qingxin., Liu, Litian., Li, Zhijian., “A new<br />
approach to convex corner compensation for anisotropic<br />
etching of (100) Si in KOH”, Sensors and Actuators., A<br />
56, (1996), pp.251-254.<br />
[8]. Mayer, G K, Offereins, H L, Sandmaier, H and Kuhl, K,<br />
“Fabrication of non-underetched convex corners in<br />
anisotropic etching of (100) silicon in aqueous KOH with<br />
respect to novel micromechanic elements,” J.<br />
Electrochem. Soc., (1990) 137, 3947–51.<br />
[9]. Pal, P., Sato, K. and Chandra, S., “Fabrication techniques<br />
of convex corners in a (1 0 0)-silicon wafer using bulk<br />
micromachining: a review”, J. Micromech. Microeng. 17<br />
(2007) R111–R133.<br />
Figure 9 SEM of the etching result for the compensation in Figure 8<br />
199
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<br />
A Programmable Neural Measurement System for<br />
Spikes and Local Field Potentials<br />
Jonas Pistor, Janpeter Hoeffmann, Dagmar Peters-Drolshagen and Steffen Paul<br />
Institute of Electrodynamics and Microelectronics (ITEM.me)<br />
University of Bremen<br />
Abstract- This paper presents a configurable system<br />
measuring and pre-processing neurological data to be<br />
transmitted over a wireless RF datalink (the datalink<br />
itself is not part of this work). The system is capable of<br />
measuring spikes and/or local field potentials of up to<br />
128 electrodes with a variable resolution up to 16 Bit at<br />
a variable sample rate up to 10 kHz consuming roughly<br />
70mW. It represents the first step of the development of<br />
an implantable neurological measurement unit. In this<br />
paper the system-architecture, the digital system and the<br />
FPGA implementation of the developed neural<br />
measurement system are presented.<br />
I. INTRODUCTION<br />
Neural engineering represents a challenging part in<br />
medical engineering. One major goal in neural engineering<br />
is the development of an electrical interface to the human<br />
brain. The motivation for this endeavor lies in the possible<br />
cure of diseases like epilepsy, Parkinson and other<br />
metabolic disorders in the brain. In addition the ongoing<br />
research in neural prostheses relies on a proper<br />
understanding of the brain functions.<br />
To meet all the requirements of research and medical<br />
facilities it is crucial to measure neurological data in longterm<br />
and due to different applications with a highly flexible<br />
amount of measured neurological data in terms of number<br />
of channels and in terms of transmitted data rate.<br />
A. Related Work<br />
Groups at Brown University [1] and at Stanford<br />
University [2] have both presented their systems to measure<br />
neurological data. From the system point of view these two<br />
systems differ in integration degree and in focus of<br />
application.<br />
The Wireless, Ultra Low Power, Broadband Neural<br />
Recording Microsystem (NRM) developed at Brown<br />
University is an implantable device designed for<br />
transmitting cortical signals from 16 channels<br />
percutaneously over a wireless data link. Since the system is<br />
fully implantable it is necessary to incorporate a power and<br />
data telemetry interacting with the external receiving<br />
equipment. The NRM incorporates a two-panel system<br />
approach. The power demanding parts of the NRM are<br />
located in the cranial unit, whereas the sensing elements are<br />
placed in the cortical unit. The two panels are connected<br />
with a percranial cable.<br />
The HermesD-System developed at Stanford University<br />
is a measurement unit placed in skull-mounted aluminum<br />
housing with the focus on a high data rate transmitted via<br />
FSK over a relatively large distance.<br />
Since the measurement unit should not be implanted, the<br />
energy for the system can be provided by batteries. The<br />
HermesD-System provides a high degree of flexibility in the<br />
system architecture since the electrical components of the<br />
measurement unit could easily be adapted or exchanged<br />
during operation.<br />
At the University of Utah [3], the INI-Chip was<br />
developed. It is a highly integrated neural measurement<br />
system, designed for a chronically cortical implant. This<br />
single-chip device integrates all modules necessary for a<br />
fully implantable neural measurement system.<br />
B. This Work<br />
The digital system presented in this paper aims to fit into<br />
an intelligent neural measurement system, combining the<br />
advantages of a fully implantable medical device with the<br />
high flexibility of a skull-mounted external measurement<br />
unit.<br />
Since the hardware components of a fully implantable<br />
device cannot be adapted or exchanged, the desired<br />
flexibility has to come from a programmable electrical<br />
device. The neuro frontend presented in this paper is<br />
capable of measuring a dedicated subset or all of its up to<br />
128 electrodes enabled by a user defined channel mask.<br />
Besides the masking of channels, each channel is widely<br />
adjustable in terms of resolution, sample rate and input filter<br />
characteristics. All these adjustments can be done during<br />
operation, leading only to small gaps in data acquisition.<br />
The system provides a timestamp counter related to each<br />
data packet, which serves as a quality indicator for the<br />
signal integrity at all times.<br />
Besides the high degree of configurability in terms of<br />
collecting and handling highly parallel measurement data it<br />
is crucial to meet the requirements of a possible (wireless)<br />
RF Transceiver concerning the format of incoming<br />
(neurological) data. The digital system described in this<br />
paper is designed for a low power RF Transceiver (ZL70102<br />
from Zarlink Semiconductor) expecting a serial SPI-data<br />
stream which is divided into packets consisting of blocks<br />
with a fixed number of bits to be sent out.<br />
200
Now the challenge in data handling with this wide range<br />
of parallel recording bitrate (tens of Mb/s for 128 channels,<br />
with 16 bit resolution at 10kS/s down to single kb/s) due to<br />
the high flexibility is to ensure that the sent blocks are<br />
always fully filled with neurological data in terms of<br />
channel transitions between two active channels with a<br />
specific resolution and a given sample rate. To ensure this<br />
effective data transmission, additional circuitry is needed to<br />
sort the valid data and to buffer the data until one whole<br />
packet is complete with filtered neurological data ready for<br />
transmission.<br />
Besides the intended RF transceiver, any other transceiver<br />
taking a serial bit stream via SPI might be used in the<br />
presented system. The measured data rate of neural signals<br />
can easily be adapted to the data rate of the transceiver to<br />
achieve a maximum throughput in data and to meet the<br />
particular needs of the intended application.<br />
II. SYSTEM-ARCHITECTURE<br />
The system architecture, where the discussed neuro<br />
frontend is a key component, is shown in Fig. 1. The neural<br />
signals are recorded by an array of passive electrodes or<br />
needles. The signals are fed into the RHA2116 from Intan<br />
Technologies, a single-chip Fully Integrated 16-Channel<br />
Biopotential Amplifier Array [4]. This array incorporates a<br />
true DC decoupling of incoming neural signals, an LNA<br />
stage with adjustable bandwidths and an analog multiplexer<br />
routing a selected channel off the chip. The true DC<br />
rejection is necessary, since there are always built-in<br />
potentials at the interface between brain tissue and<br />
electrode.<br />
The 16 channels of preamplified and prefiltered neural<br />
signals are multiplexed in a cyclic order and digitized by the<br />
16 bit on-chip low power ADC incorporated in the<br />
RHA2116 from Intan Technologies.<br />
In order to measure up to 128 channels, one has to<br />
incorporate eight of these single-chip amplifier arrays,<br />
resulting in eight parallel channels, each carrying serial data<br />
of 16 channels running in our digital system. The digital<br />
system will be described in much more detail in chapter III-<br />
“Digital system”.<br />
As depicted in Fig. 1, the whole measurement system also<br />
requires a certain transceiver, in this case the ZL70102. The<br />
power supply of the system will be realized via an inductive<br />
link, since the system is designed for implantation and<br />
hence a wired power supply is not appropriate.<br />
The inductive link could also serve as a clock source<br />
where an external clock is modulated on the power carrier.<br />
III. DIGITAL SYSTEM<br />
The digital system allows the user to interact with the<br />
implanted system via an instruction set. The user is able to<br />
reconfigure the system in its flexibility and adjusting the<br />
system to meet the requirements. To do so, there exists a set<br />
of predefined instructions described in detail in the<br />
following sections.<br />
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<br />
HUMAN BRAIN<br />
ch1<br />
ch128<br />
neuro frontend<br />
RHA<br />
2116<br />
RHA<br />
2116<br />
RHA<br />
2116<br />
RHA<br />
2116<br />
RHA<br />
2116<br />
RHA<br />
2116<br />
RHA<br />
2116<br />
RHA<br />
2116<br />
8<br />
Inductive Powering<br />
& clock recovery<br />
digital system<br />
clk<br />
filter<br />
power gating<br />
of each RHA<br />
: INPUT<br />
: OUTPUT<br />
RF-<br />
Transceiver<br />
A. Commands<br />
On/Off: These commands enable or disable the digital<br />
system. Due to this commands the system is prompt to start<br />
from or settle to a known and well defined operational<br />
status.<br />
Channel Mask: To mask each of the 128 channels there is<br />
a bit string consisting of 128 entries, where every channel is<br />
represented by a bit in the string. In a measurement<br />
scenario, where only a subset of channels is of interest, all<br />
other channels can be disabled. The number and order of<br />
active channels obtained from the user defined channel<br />
mask, is fed into the Protocol Builder in order to generate a<br />
serial bit stream to the transceiver, consisting only the<br />
active channels in a close-packed manner.<br />
Resolution: The RHA2216 produces a serial data stream<br />
as the output of the on chip ADC. Each of the 16 bit<br />
samples represents a single measured value from one<br />
electrode at a time. Besides the digitized raw data, each<br />
ADC sample is led by three fixed bits, five bits for the<br />
current channel number and a parity bit for error checking if<br />
needed. So each sample value is represented by a 25-bit<br />
data frame. The adjustable resolution only affects the 16 bit<br />
of digitized raw data. During signal propagation through<br />
our digital system the data frame is reduced to its data-bits.<br />
After cutting off the non-data bits, the resolution is<br />
adjustable from 16 down to 1 bit, cutting off the LSBs<br />
successively. Especially for spike detection where the<br />
appearance of a spike is of interest, not the exact shape, the<br />
possible reduction in resolution displays a way for saving a<br />
lot of data. The data must not be processed as well as<br />
transmitted over the RF-Link, hence reducing the power<br />
consumption of the overall system.<br />
data from 8 RHAs<br />
V DD<br />
coupling<br />
1<br />
Base<br />
Station<br />
(ZL70102)<br />
Fig. 1.System architecture with main parts of the neural measurement<br />
system including the neuro frontend and the digital system.<br />
201
Fig. 3. Architecture of the Digital Circuit.<br />
Sample Rate: The on chip ADC has a native samplerate<br />
of 10kS/s/channel. If there is no need for this high sample<br />
rate e.g. during measurement of low frequency Local Field<br />
Potentials (LFPs) or due to a very limited data rate in the<br />
RF-Link, one can scale the sample rate down to the desired<br />
value. The sample rate can be reduced to 40S/s/channel.<br />
The reduction of the sample rate also leads to reduced<br />
power consumption similar to a reduced resolution.<br />
Filter Settings: In order to reconfigure the filter<br />
characteristics of the input filters, one can select between<br />
different settings. It is obvious that measuring Local Field<br />
Potential or Spikes (Action Potentials) resulting in a<br />
different need for filtering the incoming data.<br />
The command frame is shown in Fig. 2. It consists of an<br />
ASCI-“i”, Device_ID, Command, Parameter [opt.] and a<br />
linefeed. The command frame is led by the packet type<br />
‘0011’ and since the command frame structure is byte<br />
orientated, there is an additional nibble filled with zeros.<br />
The Device_ID is important, if there are more than one<br />
measurement systems running collateral. The command<br />
byte contains the bit string for the corresponding command.<br />
If the command carries any additional parameters like the<br />
128-bit channel mask or the resolution, the parameter block<br />
in the command frame is filled with data. The linefeed<br />
indicates the end of the command frame.<br />
B. Architecture of the Digital Circuit<br />
The digital system is divided into four major blocks and a<br />
set of registers, as shown in Fig. 3. The whole data path (the<br />
way from the measurement interface to the transceiver) is<br />
kept serial, which means a lot of effort for data<br />
manipulation, but reduces hardware and bandwidth<br />
overhead for setting any resolution of the measurement data<br />
between 1 and 16 Bit. The data path doesn’t cross the<br />
Controller, but is strongly influenced by the settings of the<br />
Fig. 2. Command Frame.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
register bank defined by the Controller.<br />
Transceiver Interface (TI): Modern integrated<br />
transceivers often contain a large number of settings and an<br />
integrated microcontroller. The TI initially configures the<br />
transceiver after power on. Setting the channel, the frame<br />
size or the modulation type are some examples for the<br />
initialization process. Once the initialization is done, the TI<br />
translates the serial measurement data coming from the<br />
Protocol Builder to meet the protocol of the transceiver.<br />
The TI contains an SPI Interface and writes the data into<br />
the TX Buffer of the transceiver. Besides the serial interface<br />
for the measurement data, the TI has an 8-bit parallel<br />
interface to the Controller for receiving commands and<br />
sending status messages. Fig. 5 shows the SPI output<br />
spi_sdo of the Transceiver Interface, where the RX buffer is<br />
polled and where the measurement data is written into the<br />
TX buffer of the transceiver.<br />
Measurement Interface (MI): This interface accesses the<br />
eight RHA2116 ICs. Each RHA samples 16 analog inputs<br />
with a resolution of 16 Bit at 10kS/s/channel, regardless of<br />
the quality and the characteristics of the input signals. As<br />
mentioned in the description of the resolution command<br />
one task of the MI is to gate the incoming data in order to<br />
extract the 16 bits of measurement data out of the 25-bit<br />
data frame.<br />
If the complete 25-bit data frame would be transmitted<br />
instead, one would need to have an RF-Transceiver capable<br />
of handling up to 11.52 MB/s additional transmit-data. In<br />
order to cut off the protocol overhead from the 16-bit raw<br />
data, there is an internal counter triggered by the<br />
synchronous pulses generated by the RHA. This internal<br />
counter is also used to cut off a subset of LSBs as specified<br />
in the resolution command.<br />
Therefore the MI-logic enables only valid data, in terms<br />
of user desired data out of the incoming serial bit stream to<br />
propagate in an ongoing serial manner to the Protocol<br />
Builder.<br />
To ensure that there is only valid data running through the<br />
MI, the Measurement Interface needs to evaluate the user<br />
defined parameters stored in the Register Bank and apply<br />
them on the serial data stream.<br />
To evaluate the channel mask, the MI has to compare the<br />
current active channel, represented by the channel number,<br />
with the corresponding entry in the channel mask. Even if<br />
the channel number is included in the 25-bit data frame, it is<br />
not uses for data handling, since the fact that the channel<br />
number is not encoded in an optimal way in terms of bit<br />
representation. The channel number encoded in the<br />
RHA2116 data stream is designed to maintain compatibility<br />
with a certain ADC from Analog Devices. To reduce the<br />
protocol overhead in data-handling the authors generated a<br />
new channel number encoded straight binary (Fig. 5 –<br />
mea_mux_ch[3:0]).<br />
Besides the channel mask the sample rate also has a direct<br />
influence on the outgoing serial data stream of each of the<br />
eight data lines. If there is an application specific need to<br />
reduce the sample rate e.g. by half, the MI logic disables<br />
every other sample value, containing one measurement<br />
202
value of every active channel. Therefore the MI has to know<br />
the beginning of a new measurement cycle. To do so, the<br />
MI evaluates a synchronization impulse, generated by the<br />
RHA every time one measurement cycle is finished. In<br />
addition it is important to execute adaptations to the<br />
commands only at a defined state of the MI to avoid<br />
unwanted intermediate states.<br />
Also the synchronization to external events like eye<br />
movement of the patient might be of interest in certain<br />
applications. In terms of signal integrity one also would like<br />
to ensure that every data packet is collected properly. If<br />
there is a loss in measurement data, it could be of interest<br />
which packets are missing. With a timestamp generated by<br />
the MI is it possible to detect any packet loss at the base<br />
station and to correct the inaccuracy of the integrated<br />
onboard quartz.<br />
The timestamp is an internal 16 bit counter incremented<br />
by each recorded measurement sample (10kHz) regardless<br />
of settings in the samplerate or channelmask. This<br />
timestamp is attached to every transmitted data packet (Fig.<br />
6). With this counter one is able to verify if the transmitted<br />
timestamp matches the expected one, or if there is a loss in<br />
synchronicity leading to degradation in signal validity.<br />
If there is a demand for operating only a dedicated subset<br />
of RHAs, the MI has the functionality to disable the<br />
outgoing data stream of the unwanted RHAs. Due to this<br />
switch, the further instances connected to the disabled<br />
RHAs stop operation as well. This leads to a reduced power<br />
consumption of the overall system.<br />
The MI also evaluates the parity-bit included in the 25-<br />
bit-data-frame of each of the eight RHAs. If there is a parity<br />
mismatch in one of the user defined active channels, there is<br />
an error-flag indicating a bit failure.<br />
Protocol Builder (PB): The Protocol Builder assembles<br />
the measurement data into a compact data packet including<br />
a header with the parameters from the Register Bank. The<br />
packet has a variable length, depending on the number of<br />
channels of the current measurement and the selected<br />
resolution. Several constraints made the Protocol Builder<br />
growing to the largest block of the design: The<br />
measurement data comes on eight lines in bursts of 4 MHz<br />
while the transceiver expects data to be clocked in with 4<br />
MHz. This means for the worst case an input rate of 32<br />
Mb/s. The data has to be stored until the next sample burst<br />
and written into the TX buffer. To keep the memory size<br />
and the latency low, we have implemented a dedicated<br />
memory management. To keep the memory at a minimum,<br />
the memory works bitwise, so if the resolution is chosen to<br />
have a width of 5 Bit, we only need 5 Bit of memory, the<br />
data of the following channel is written to the adjacent<br />
memory address. Furthermore the memory size and the<br />
latency is reduced by taking the order of the incoming<br />
channels into account. All eight RHAs are running<br />
synchronous and each delivers a sequence of 16 channels in<br />
a serial order. Each RHA has a dedicated memory<br />
(implemented as a circular buffer with read and write<br />
pointers) to handle the eight serial data streams<br />
simultaneously (shown in Fig. 4). The required memory size<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
depends on the bitrate ratio (measurement rate vs.<br />
transmission rate) and on the delay between the start of a<br />
measurement burst and the start of the transmission. Instead<br />
of transmitting the data from one RHA after another, the<br />
128 Channels are interleaved in a way that the channels are<br />
transmitted in the same order the measurement data comes<br />
in. One major challenge of this interleaving is the fact that<br />
the order is not fixed, but depends on the channels the user<br />
actually has selected for the running measurement. For this<br />
reason the Protocol Builder has an extra program memory<br />
called “channel stack”, which is written by the Controller<br />
every time the user defines a new channel mask.<br />
Fig. 5 shows two serial input lines of the Protocol Builder<br />
(coming from the Measurement Interface). The signal<br />
sdatabus_out[0] contains four blocks with nine bit of data,<br />
which is the measurement data according to the activation<br />
bits of the channel mask with the index 0,1,14 and 15. The<br />
signal sdatabus_out[7] contains 16 blocks, according to the<br />
channel mask bits 112-127. Bufferlevel0 and Bufferlevel1<br />
show, how the memory of the data lines is filled. The point<br />
where pstate gets the value three, the read sequence of the<br />
buffers is started and the measurement data is handed over<br />
as a compressed block to the transceiver interface. The<br />
signal rha_select (Fig. 5) shows (while pstate equals “3”),<br />
how the different memories (shown in Fig. 4) are read out in<br />
the following order: RHA0, RHA7, RHA0, RHA7 for a<br />
longer time and again RHA0, RHA7, RHA0, RHA7. The<br />
example shows how the distributed measurement data in the<br />
incoming data streams (mea_mux_ch 0-15) is packed into a<br />
relatively short sequence in the s_data signal (in the<br />
duration where pstate equals 3).<br />
The packed data can be decoded to the correct channels<br />
by interpreting the channel mask, which is part of the<br />
header.<br />
Controller: After power on, the system waits in a standby<br />
state for commands. The Transceiver Interface uses an 8-Bit<br />
parallel port for delivering commands and parameters to the<br />
Controller. The Controller writes the parameters of the<br />
Fig. 4. Protocol Builder.<br />
203
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May 2011, Aix-en-Provence, France<br />
<br />
FPGA<br />
(XC3S50AN)<br />
Connector for RF-<br />
Transceiver<br />
LED-Array<br />
Fig. 5. Data path from the analog frontend to the transceiver for one<br />
sample interval. In this example, the resolution was set to 9 Bit and the<br />
channel mask is set to FFFF000000000000000000000000C003, which<br />
means, that four channels of the first, and 16 channels of the last RHA<br />
are transmitted while the remaining RHAs 1-6 are ignored [5].<br />
commands into the according registers. In case of a new<br />
channel mask, the measurement gets interrupted, and the<br />
channel stack of the Protocol Builder is reprogrammed.<br />
After the “measurement on” command, the Controller<br />
checks if the transceiver is initialized and enables the<br />
measurement sequence.<br />
Register Bank: The registers store the values of the<br />
command parameters which are also included in the<br />
measurement data frame (Fig. 6): Channel_Mask[127:0],<br />
RHA_Filter[3:0], Sample_Rate[7:0] and Resolution[3:0].<br />
IV. RESULTS<br />
A. FPGA Prototype<br />
The FPGA prototype will serve as a development<br />
platform on the way to higher integration. It allows us to<br />
develop the software and test the compatibility with the<br />
latest version of our digital system. Fig. 7 shows the FPGA<br />
prototype without the transceiver board, which is normally<br />
placed on top of the stack. The prototype provides the<br />
smallest device of the Spartan3AN series, some simple<br />
debug capability (pinheader and LEDs), a power supply (for<br />
battery operation), sockets for the RHA modules, the<br />
external filter components and connectors for attaching up<br />
to 128 electrodes plus reference electrodes.<br />
The prototype has a size of 5x5 cm² and a height of 3 cm.<br />
The power consumption for the prototype with one<br />
connected RHA is about 120 mW. The dynamic loss of the<br />
digital core is below 200µW, which is the difference in<br />
overall power consumption between a running measurement<br />
and the digital core held in the reset state. The design<br />
utilizes 363 Flip-flops and 693 LUTs.<br />
Power supply<br />
& crystal unit<br />
Slots for<br />
additional<br />
RHAs<br />
Fig. 7. FPGA prototype for functional verification.<br />
RHA2216<br />
assembled on<br />
socket<br />
B. Test System<br />
For testing the prototype we use a National Instruments<br />
PXI System with a FlexRIO card. This allows us to reuse<br />
modules from the Simulation Testbench, which can be<br />
implemented on the FPGAs of the FlexRIO card. The User<br />
Interface (Fig. 8) was programmed using LabView and<br />
allows to control all available functions of the system. Fig. 9<br />
shows some qualitative measurement results from a sine<br />
stimulus.<br />
V. CONCLUSION<br />
The programmable neuro frontend presented in this paper<br />
covers nearly every conceivable application in neural<br />
engineering or neural prostheses. Due to the high degree in<br />
flexibility, one can easily shrink or expand the system<br />
performance in order to fit the functional constraints defined<br />
by the desired application.<br />
Even if the constraints are partly unknown or unspecified,<br />
one can use our system to evaluate and to determine the<br />
needed performance in order to meet the objective target.<br />
Fig. 6. Measurement Data Frame. Fig. 8. LabView Software for Controlling the Prototype [6].<br />
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<br />
especially fully implantable systems have a rather rigid<br />
design in terms of system architecture and therefore<br />
functionality. Actually the presented digital system tries to<br />
combine key advantages of the measurement systems listed<br />
in the first chapter: our digital system incorporates the<br />
flexibility of the Hermes-system, the aim of a fully<br />
implantation similar to the NRM - but without the cranial<br />
unit and therefore without the percranial cable - and the<br />
high degree of integration similar to the INI-chip, at least<br />
Fig. 9. Waveform view of recorded sine waves from the prototype [6]. for the digital system planned to included in an ASIC.<br />
After the target is met, the system still keeps its flexibility<br />
for future applications without any modifications necessary.<br />
If for some reasons a rigid system in terms of performance<br />
is needed, it is also possible to adapt the system on<br />
hardware base with a relative low effort, since all<br />
performance parameters are already known from the<br />
evaluation process.<br />
Due to the serial data handling and the fact that undesired<br />
measurement data and a lot of protocol overhead is removed<br />
in an early stage regarding signal propagation, the system<br />
has an increased performance with a low hardware<br />
complexity and therefore a reduced power demand.<br />
The flexibility of the system allows the user to fill the<br />
limited transceiver bandwidth with the best fitting product<br />
of resolution, number of channels and sample frequency,<br />
with respect to the particular application.<br />
A. Future Work<br />
Future work will concentrate on the ongoing increase in<br />
integration of all electrical components in order to achieve<br />
the goal of a fully implantable neural measurement micro<br />
system.<br />
The final goal is a single chip solution incorporating the<br />
whole signal path starting from the passive electrode/needle<br />
ending at the RF-Transceiver interface. To satisfy the<br />
demand in higher numbers of electrodes and thereby an<br />
increase in neural data, one has to think about sophisticated<br />
ways of data reduction without losing any neural<br />
information. A reduction in data rate through data<br />
compression also reduces the power consumption of the<br />
measurement system.<br />
Besides the flexibility in performance, it is also desirable<br />
to have a certain degree of redundancy if some parts of the<br />
system are malfunctioning. This redundancy has a direct<br />
influence on the system reliability, which is crucial for a<br />
non-removable fully implantable medical device. So the<br />
“perfect” system consists of several measurement units,<br />
each totally autonomous in terms of power supply and data<br />
link, carrying all the flexibility described in this paper.<br />
In this manner one gets a measurement system where<br />
each electrode is connected to at least two-subsystems, so<br />
there is a fairly high chance that each electrode is at least<br />
represented once in the overall system, able to propagate its<br />
neural data through the neural measurement system.<br />
B. Compared to Other Work<br />
Compared to other work, the work presented in this<br />
paper has a significant degree of flexibility. Other systems,<br />
ACKNOWLEDGMENT<br />
The authors would like to thank the German Federal<br />
Ministry of Education and Research (BMBF) for<br />
subsidizing this work within the KALOMED-project. Also<br />
the authors would like to thank Mr. Opel for his valuable<br />
support in technical implementation of the FPGA-based<br />
prototype.<br />
REFERENCES<br />
[1] Y.-K. Song, D. A. Borton, S. Park, W. R. Patterson, C. W. Bull, F.<br />
Laiwalla et al, “Active Microelectronic Neurosensor Arrays for<br />
Implantable Brain Communication Interfaces,” in IEEE Trans. on<br />
Neural Systems and Rehabilitation Engineering, vol. 17, no. 4,<br />
August 2009, pp. 339-345.<br />
[2] Henrique Miranda, Vikash Gilja, Cindy A. Chestek, Krishna V.<br />
Shenoy and Teresa H. Meng, “HermesD: A High-Rate Long-<br />
Range Wireless Transmission System for Simultaneous<br />
Multichannel Neural Recording Applications,” in IEEE Trans. on<br />
Biomedical Circuits and Systems, vol. 4, no. 3, June 2010, pp. 181-<br />
191.<br />
[3] Reid R. Harrison, Ryan J. Kier, Cynthia A. Chestek, Vikash Gilja,<br />
Stephen Ryu, Bradley Greger et al, “Wireless Neural Recording<br />
with Single Low-Power Integrated Circuit,” in IEEE Trans. on<br />
Neural Systems and Rehabilitation Engineering, vol. 17, no. 4,<br />
August 2009, pp. 322-329.<br />
[4] “RHA2116 – Fully Intergated 16-Channel Biopotential Amplifier<br />
Array”, intan Technologies, LLC, Datasheet, 19 May 2010.<br />
[5] Generated with the assistance of cadence-SimVision ©.<br />
[6] Generated with the assistance of LabView 2010 from National<br />
Instruments©.<br />
Contact: Jonas Pistor, Institute of Electrodynamics and<br />
Microelectronics (ITEM.me) University of Bremen, Otto-<br />
Hahn-Allee, NW1, 28359, Bremen, Germany, +49 421 218-<br />
62539 Email: pistor@me.uni-bremen.de<br />
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<br />
PANEL DISCUSSION<br />
HERMETICITY TESTS IN MEMS<br />
Marc DESMULLIEZ, Heriot-Watt University<br />
Suzanne COSTELLO, MCS Ltd.<br />
Wolgang REINERT, Fraunhofer Institute for Silicon Technology<br />
Steven MARTELL, Sonoscan Inc.<br />
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<br />
Hermeticity Test Methods for MEMS:<br />
Where are we?<br />
Marc Desmulliez<br />
MIcroSystems Engineering Centre, School of Engineering and Physical Sciences, Heriot-Watt University,<br />
Edinburgh, EH14 4AS, Scotland, UK.<br />
E: m.desmulliez@hw.ac.uk T: +44 (0) 131 451 3340<br />
ABSTRACT:<br />
This short presentation will set the scene of the agenda of this panel and indicates current commercial methods and<br />
R&D efforts in the field of hermeticity tests for low cavity volume packages such as those encountered in MEMS.<br />
Limitations of the military standards MIL-STD-883H and MIL-STD-750E will be briefly explained and guidelines to<br />
use existing the He fine leak test method will be provided.<br />
BIOGRAPHY:<br />
Marc Desmulliez is currently the founder-director of the Microsystems Engineering Centre (MISEC) at Heriot-Watt<br />
University, which is the 4 th largest academic MEMS research group in the UK. He is a physicist/electrical engineer<br />
by educational and professional background and has authored over 280 papers in the fields of MEMS,<br />
optoelectronics and advanced manufacturing techniques. He span out the Company MicroStencil Ltd in 2003 which<br />
is now operating as a partnership with DEK in Singapore.<br />
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11-13 May, Aix-en-Provence, France<br />
<br />
Standards for Hermeticity Test Methods<br />
Suzanne Costello<br />
MIcroSystems Engineering Centre, School of Engineering and Physical Sciences,<br />
Heriot-Watt University, Edinburgh, EH14 4AS, Scotland, UK.<br />
E: S.Costello@hw.ac.uk T: +44 (0) 131 451 3774<br />
ABSTRACT:<br />
Traditional hermeticity test methods and the standards used to ensure correct usage of these test methods have been<br />
shown to have limitations when applied to low cavity volume MEMS packages. Typical MEMS cavity volumes are<br />
well below the minimum volumes stated in the military standards leading to inaccuracies when the traditional<br />
hermeticity tests are carried out on such packages. Ultra low vacuum packaging is required for many MEMS<br />
applications, reducing the maximum acceptable leak rate of these packages below the measurable range of most test<br />
methods given in the military standards. New packaging materials used in MEMS manufacturing have limited the use<br />
of standard tests since new leak sources are apparent which the traditional test methods were not designed to measure<br />
and so are not considered in the military standards.<br />
SEMI MS8-0309 - GUIDE TO EVALUATING HERMETICITY OF MEMS PACKAGES was written to inform and<br />
guide users on the best way to quantify leaks which may adversely affect the performance of MEMS devices. Two<br />
further test standards are currently being written to provide further information and guidelines to quantifying<br />
permeation leaks and outgassing. The first of these standards, “SEMI Standards: Fluid Permeation through MEMS<br />
Packaging Materials”, will be discussed.<br />
BIOGRAPHY:<br />
Suzanne Costello graduated from Heriot-Watt University in Edinburgh in 2004 after receiving a masters degree in<br />
physics. She is currently working towards an engineering doctorate in microsystems engineering. Her sponsoring<br />
company is materials and failure analysis specialists, MCS Limited. Her research has been based on hermeticity<br />
testing of MEMS and low cavity volume microelectronic packages. She has published several papers in this field<br />
which have shown the theoretical limitations of the traditional hermeticity test methods when applied to today’s<br />
packages and the advantages of in situ test structures for assessing leak rates of low volume packages. Suzanne has<br />
also been involved in the task force working towards creating a new standard to assess fluid permeation through<br />
MEMS packaging materials.<br />
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11-13 May, Aix-en-Provence, France<br />
<br />
Q-factor monitoring as a 100% leak screen in<br />
industrial applications<br />
Wolfgang Reinert<br />
Fraunhofer Institute for Silicon Technology,<br />
Fraunhofer Str. 1<br />
D-25524 Itzehoe, Germany<br />
E: wolfgang.reinert@isit.fraunhofer.de T: +49 (0) 4821 17 4617<br />
ABSTRACT:<br />
Neon ultra fine leak testing enables a fast screen of the initial leak rate of hermetically sealed MEMS microresonating<br />
devices. The method is based on monitoring the Q-factor of resonating structures before and after a defined Neon gas<br />
bombing.<br />
The presentation will highlight the main characteristics of this method and the different modifications of the<br />
procedure that may be used in high volume production to save process time. Limitations and instabilities of the<br />
electronic Q factor measurement will be discussed to provide a better understanding and avoid false data<br />
interpretation. Last the role of H 2 O as a possible test gas for nano-leaks will be shortly explained.<br />
BIOGRAPHY:<br />
Dr.-Ing. Wolfgang Reinert studied Physics at the University of Bonn (Germany) . After a post-doctoral scholarship at<br />
the Technical University of Trondheim (Norway), he spent six years, as a scientist at the Centre for Microjoining<br />
Technology (CEM) in Germany. For the last 11 years, he is the group manager of „Advanced Packaging“ at the<br />
Fraunhofer ISIT.<br />
His research interests include: the construction of cap wafers for IR and interial sensors, the hermetic packaging of<br />
MEMS based on metallic sealing, the development of avanced getter and the development of pilot lines for MEMS<br />
module assembly.<br />
Dr. Reinert is the inventor of 46 patents or patent applications and has contributed to 3 technical books. He has<br />
authored numerous technical papers in the field of packaging.<br />
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11-13 May, Aix-en-Provence, France<br />
<br />
Evaluating the Seal Integrity of MEMS<br />
Hermetic Packages<br />
ABSTRACT:<br />
Steven R. Martell<br />
Technical Support Services Manager<br />
Sonoscan, Inc. , 2149 E. Pratt Blvd.<br />
Elk Grove Village, IL 60007<br />
T: 847-437-6400<br />
For many years Acoustic Microscopy (AM) techniques have been utilized to evaluate the quality of the bond between<br />
materials, especially for microelectronic devices. AM has been established as one of the few techniques within SEMI<br />
MS8-0309- GUIDE TO EVALUATING HERMETICITY OF MEMS PACKAGES that can provide reliability and<br />
quality control data, but little has been done to determine the minimum seal width required to ensure long term<br />
hermeticity.<br />
AM methods of non-destructive analysis incorporate techniques that provide data on how well a lid is bonded to the<br />
cavity package, the actual width and thickness of the seal material, plus any voids embedded within it. What is not<br />
known at this time is the minimum acceptable seal width/path required based on the permeability of the various seal<br />
materials.<br />
BIOGRAPHY:<br />
Steven R. Martell is the manager of technical support services at Sonoscan, allowing him to work with companies and<br />
standards organizations on an international basis. He has a B.S. in Mechanical/Ocean Engineering from University<br />
of Rhode Island.<br />
He is the current chairman of IPC's B-10a - Plastic Chip Carrier Cracking Task Group. He has received<br />
"Outstanding Performance", "Distinguished Committee Service" and "Special Recognition" awards for his leadership<br />
of the joint IPC/JEDEC working group that developed J-STD-020, J-STD-033, J-STD-075, etc. He is now working in<br />
cooperation with IPC and JEDEC, to coordinate these standards with other international standards organizations,<br />
such as EIAJ and IEC. He is also a contributing member to the SEMI MEMS and Solar standards committees, plus<br />
the chairman of the MEMS terms task group within SEMI. In addition, he was the main author of the revised Method<br />
2030 for Die Attach Evaluation within MIL-STD-883 through GEIA G12 standards committee. He is the author of<br />
over thirty papers and technical publications.<br />
Sonoscan, Inc., Corporate Headquarters: 2149 East Pratt Blvd ● Elk Grove Village, IL 60007 ● T: 847.437.6400 ●<br />
F: 847.437.1550 ● www.sonoscan.com Silicon Valley, CA ● Phoenix, AZ ● Boston, MA ● England ● Philippines ●<br />
Singapore ● Shanghai<br />
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<br />
PANEL DISCUSSION<br />
HIGH ADDED VALUE MEMS<br />
Jérémie BOUCHAUD, Director - Principal Analyst MEMS & Sensors, IHS iSuppli, Munich, Germany<br />
Jean Michel KARAM, Chairman & CEO, MEMSCAP, Bernin, France<br />
Sean NEYLON, CEO, Colibrys, Neuchâtel, Switzerland<br />
Thierry BRISARD, CEO, NEOSENS, Toulouse, France<br />
ABSTRACT:<br />
High-value MEMS as sensors and actuators for applications that are outside the high-volume consumer<br />
electronics and automotive volume markets, addressing the industrial, medical, energy, optical telecom and<br />
aerospace-defense market segments.<br />
According to iSuppli's five-year forecast, high-value MEMS revenue will hit $2.6 billion in 2014. This rapid<br />
growth is being driven by the value proposition brought by the MEMS devices to many systems and<br />
applications.<br />
The panel will start by an overview of the high added value MEMS market, followed by several case studies<br />
from active players in this market.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
Integrated Sensing Systems for Health and Safety<br />
Kiyoshi Itao<br />
Tokyo University of Science<br />
2-6 Kagurazaka, Shinjuku-ku<br />
Tokyo, 162-0825, Japan<br />
Toshihiro Ito<br />
Advanced Industrial Science and Technology<br />
1-2-1 Namiki, Tsukuba,<br />
Ibaraki, 305-8564, Japan<br />
Abstract- Information and communication services have been<br />
developed from the first generation of fixed phone era to the<br />
second generation of cellular phone era, and eventually the<br />
third generation will be developed as the era of keyboard-less<br />
and wireless sensor terminals. Recently, with the technology<br />
weaving of microsensors, wearable computers and wireless<br />
systems, various micro sensing systems have been developed. In<br />
the near future, with more progress in nanotechnologies,<br />
flexible sensors, flexible power supplies, flexible computers and<br />
flexible displays will be developed, and every terminal will<br />
become wearable. From this viewpoint, we would like to explain<br />
the novel technologies and services in the third generation<br />
network. In details, we introduce Human Recorder System<br />
development in our NPO-WIN (Wearable Information<br />
Network http://www.npowin.org/j/)<br />
I. INTRODUCTION<br />
With the recent technology weaving of microsensors,<br />
wearable computers and wireless systems, various<br />
microsensing systems have been developed. The<br />
nanotechnologies will lead to the rapid progress of flexible<br />
sensors, flexible power supplies, flexible computers and<br />
flexible displays. Every terminal will become wearable in the<br />
near future, and textile in the future on the basis of<br />
nanotechnologies as shown Fig. 1.<br />
The nature interfacer is core device of such a new sensor<br />
communication system that can capture various kinds of<br />
information using microsensors, process data recognition by<br />
one clip computers, and transmit information by wireless<br />
technologies. The key technology here is the software to<br />
automatically decide whether to stay active or to become<br />
sleep to save energy, as well as the relevant hardware<br />
technologies. The smaller these nature interfacers become,<br />
the more their application range will be expanded, for<br />
example from an airplane to a small bird.<br />
In this paper, we will present three fundamental ideas,<br />
firstly the concept of nature interface, secondly wearable<br />
information technologies, and thirdly examples of typical<br />
applications, such as Human Recorder System by NPO-WIN<br />
(Wearable Information Network http://www.npowin.org ).<br />
In conclusion, we will discuss future information<br />
communication systems utilizing human vital information.<br />
II. AIMING AT THE HARMONY BASED ON<br />
SENCER INFORMATION COMMUNICATION<br />
The present telecommunication network system is of<br />
human-centered, as shown in Fig.2. The nature system is<br />
completely separated from the artifacts system. Only a small<br />
portion of the huge nature (real world) information is<br />
incorporated info this system because of poor sensing<br />
technology.<br />
In contrast, what we have advocated is a new<br />
telecommunication system called “sensor communication<br />
system,” which could make it possible to gather a huge<br />
amount of information from nature, including both animals<br />
and plants, as well as information from artifacts (Fig.3). It<br />
would be an information communication system with a thick<br />
information input pipe, and various sensor groups. It could<br />
be also the system that can monitor the state of nature and<br />
artifacts concisely and widely. Such technologies could<br />
become the foundation for the progress of science to general<br />
and transdisciplinary science.<br />
Fig. 1 Mt. Fuji of Technology<br />
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<br />
transmission of results to a computer which could be<br />
connected via communication networks.<br />
Fig.2. Second generation network based on cellular phone<br />
Fig.3. Third generation network based on sensing terminals<br />
Fig.4. Wearable Computing toward Nature Interface<br />
III.<br />
WEARABLE COMPUTING TOWARD<br />
NATURE INTERFACE<br />
TABLE I<br />
From mobile communication to sensor communication<br />
The volume and weight of information communication<br />
devices should ideally be minimized to zero as the ultimate<br />
target. Micro system technologies are being developed from<br />
various fields of study, and the efforts continue proceeding.<br />
While computer development continues following the<br />
traditional path of human-operated machines, we are also<br />
facing the development of a “pervasive computer”, where<br />
computers run without human operators. Furthermore, the<br />
Internet has evolved to version 6 (IP. Ver. 6), and the time<br />
has come when all the devices on the earth could be<br />
connected to a network and identified uniquely. Moreover<br />
,the device for short-distance radio communications has been<br />
miniaturized to one chip. The technology that develops the<br />
interfaces among nature, human beings, and artifacts is<br />
achieved by information microsystem technology with the<br />
following two trends: technology of miniaturization, and<br />
pervasive computers. Wild animals, human beings, and<br />
mobile artifacts could be equipped with Nature Interfacers,<br />
which constantly monitor and evaluate sensor information<br />
input, process the monitoring information to recognize the<br />
state of objects, and perform control or diagnosis by wireless<br />
technologies. Nature Interfacers like this have been<br />
developed so far (Fig.4).<br />
The forms of computer terminals are classified as shown in<br />
Table 1. Basically, human beings operate conventional<br />
computers by giving instructions as digital inputs through<br />
keyboards. Computer downsizing has been accompanied by<br />
rapid advances in LSI technologies and micro-machine<br />
technologies, as well as revolutionary advances in the<br />
personalization and mobility of information. And, finally, the<br />
technologies have evolved to allow wearable computers.<br />
Furthermore, computer automation has been developed<br />
based on the following technologies: detection of analog<br />
information into digital quantity, further recognition of this<br />
information based on knowledge (the database), and<br />
IV. WEARABLE INFORMATION NETWORK USING<br />
SENSORS TERMINAL<br />
With the technology explained above, we should develop a<br />
terminal suitable for sensor communication. It should be<br />
equipped not only by human beings but also by animals, or<br />
artifacts, and should serve as a key device in detecting nature<br />
information, including health monitoring information,<br />
position of an animal or degradation of artifacts,<br />
environmental information, etc. This is the micro<br />
information terminal that I have proposed as the “Nature<br />
Interfacer”(Fig. 5)<br />
The Nature Interfacer is the device that captures various<br />
kinds of information using microsensors, processes data<br />
recognition by one clip computers, and transmits information<br />
by wireless technologies. The key technology here is the<br />
software to automatically decide whether to stay active or to<br />
become sleep to save energy, as well as the relevant<br />
hardware (mechatronics) technologies. The smaller these<br />
nature interfacers become, the more their application range<br />
will be expanded, for example from an airplane to a small<br />
213
ird (Fig. 6)<br />
The computer communication society will advance<br />
continuously due to the development of related technologies.<br />
Those technologies, however, should not be applied to tools<br />
only for people's pleasure and convenience. Instead, we<br />
should apply such advanced technologies to our<br />
environment, and should take urgent measures to prevent<br />
environmental destruction in the 21st century. By fusion of<br />
the Internet and cellular phone technology, the world of the<br />
Nature Interfacer will be realized in the near future. Fig. 7<br />
shows a conceptual figure.<br />
Fig. 5. Nature Interfacer<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
V. HUMAN RECORDER SYSTEM DEVELOPMENT<br />
FOR SENSING THE AUTONOMIC NERVOUS SYSTEM<br />
An airplane is equipped with a flight recorder, a ship with a<br />
voyage recorder, and recently drive recorder is being<br />
installed in cars. Sure, human also needs recording<br />
equipment for his vital data then. I named this "Human<br />
recorder". We defined “Human Recorder” as a wearable<br />
device that constantly detects and records vital signs 24<br />
hours, 365 days.<br />
Fig. 8 shows an image of health monitoring system<br />
enabled by the Human Recorder System. A micro energy<br />
powered chipped sensor composed with an<br />
electrocardiograph (ECG), an electroencephalograph (EEG),<br />
a thermometer for skin temperature, and an tri-axis<br />
accelerometer, sends data to a smartphone by wireless using<br />
weak radio signal. The smartphone linked to the Internet<br />
through seamless communication network will forward<br />
periodically stored data to a computer (memory storage) unit<br />
on which data display and analysis software is provided.<br />
In the technological progress of recent microdevices, we<br />
could achieve the development of the first stage of Human<br />
Recorder (Fig. 9). The main microdevice is miniature and<br />
lightweight electrocardiograph, 40×35mm in size, 7mm in<br />
thickness, and 11g in weight, (smallest, lightest in the world),<br />
combined with a tri-axis acceleration sensor, a thermometer<br />
for skin temperature, battery and wireless data transmission<br />
chip.<br />
Animals<br />
Artifact<br />
Sensor<br />
Watch-type computer<br />
Information (position<br />
physiology,chemicals)<br />
Micro generator<br />
Fig.8 Composition of Human Recorder system<br />
To network<br />
Fig. 6. Application of Nature Interfacer<br />
Wireless transmitter<br />
Fig.9 The first stage of Human Recorder:<br />
smallest and lightest ECG device in the world<br />
Fig. 7. Positioning of Nature Inter-facer in the third generation network<br />
This new device enables, noninvasive, precise, real-time<br />
detection and collection of ECG, which is the key for human<br />
autonomic nervous system detailed analysis, by the<br />
following process.<br />
• Human Recorder is attached to the chest. (user does<br />
not feel any discomfort with the device)<br />
• User can spend time as usual.<br />
• ECG can be measured continuously and recorded<br />
into a smartphone for about 120 hours.<br />
• The heart beat cycle is extracted from the<br />
214
electrocardiogram.<br />
• The heart beat cycle time variation computation<br />
• The heart beat cycle fluctuation frequential analysis<br />
Typical result of ECG heart beat cycle frequential<br />
analysis is presented on , where strength of frequency zone<br />
“L” reflects sympathetic nerve’s activity, and strength of<br />
frequency zone “H” reflects parasympathetic nerve’s activity<br />
(Fig. 10). Then, L/H ratio is an indicator of activity dominant<br />
nervous system (large: sympathetic nerve is dominant, small:<br />
parasympathetic nerve is dominant).<br />
Fig.10 Image of heart beat cycle fluctuation frequential analysis<br />
VI. FUTURE HEALTHCARE SYSTEM CONSIDERATION<br />
It has been recognized to be able to evaluate sleeping with<br />
not only brain wave but also HRV spectral analysis with data<br />
of Human recorder in case 1. The parasympathetic nervous<br />
system becomes dominant while sleeping and the<br />
sympathetic nervous system becomes chiefly dominant<br />
while awaking. Therefore, the circadian rhythm of the<br />
sleeping-awaking can be observed by analyzing the heart<br />
beat change. That is, it was regarded that the distinction<br />
between REM sleeping and non REM sleeping and<br />
evaluation of quality while sleeping was possible, too. It<br />
means “excellent sleep” by considering of the sleeping<br />
rhythm in this case. Additionally, it is going to be also<br />
possible to observe the change of a physiologic index that<br />
influences the autonomic nervous system such as<br />
perspiration, physical change, and body temperature by<br />
HRV spectral analysis.<br />
In case 2, under the high stress level, the sympathetic<br />
nerve becomes dominant. To the contrary, the<br />
parasympathetic nerve activity becomes dominant with<br />
comical video screening, and during acupunctural treatment.<br />
It has been recognized that a voluntary laugh, when comical<br />
images are selected by oneself, has influence on autonomic<br />
nervous system reaction. In this study, it would not be<br />
possible to confirm the hypothesis that combining<br />
acupuncture and laugh by comical video has an effect on<br />
relaxation. Moreover, the change of autonomic nervous<br />
system has been captured at various time span. It is useful<br />
and needed to evaluate autonomic nervous system as a<br />
simple tool to clarify the physiology mechanism.<br />
Former studies reported that HRV spectral analysis was<br />
effective to evaluate the physical and mental loads by<br />
quantifying respectively the activity level of sympathetic<br />
and parasympathetic nerves [2,3]. However, HRV spectral<br />
analysis index is known to be different according to the age,<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
sex and the individual variation. Not only the examination of<br />
measurement condition, H and L/H, but also the<br />
examinations expressing the bio-rhythm, such as "1/f<br />
fluctuation" might be necessary to improve the accuracy of<br />
autonomic nervous system evaluation in future research.<br />
From the results of this study, we propose a Human<br />
Recorder System using ECG device and adapted<br />
visualization software as an effective tool for stress<br />
management.<br />
In our aging society, we are sure that a lot of elderly<br />
people hope they could have good information on their own<br />
vital (physical and mental) conditions and, if necessary, they<br />
could be cared and cured timely with minimum of helping<br />
power. We’ve been now challenging to build such a sensor<br />
networking society in the near future. Fig. 11 shows an ECG<br />
transmitter, which is expected as a micro-bio telemetry.<br />
As the application of ECG data analysis was examined in<br />
this study, in the future the Human Recorder system could be<br />
composed also with some functional sensors such as EEG,<br />
and the software for processing and display could be adapted<br />
according to various needs. The Human Recorder system,<br />
combined with a set of warning functions such as sound<br />
alarm, vibrations will be useful in a lot of application in the<br />
healthcare field.<br />
The evidence of human being’s vital information is<br />
potentially too rich and sensitive to be measured by Human<br />
recorder. Human being’s vital information remains<br />
undeveloped so far, and we are now beginning to recognize a<br />
human being as a sensor itself united with mind and body.<br />
The evidence of human being’s vital information is<br />
potentially too rich and sensitive to be measured by digital<br />
devices.<br />
Fig.11 Micro-bio telemetry system using Human Recorder<br />
VII. FUTURE INFORMATION COMMUNICATION SYSTEMS<br />
It We envision that Future Information Communication<br />
Systems are going toward the building of a new generation<br />
of sensor networking society by firstly checking up human<br />
vital information as well as knowledge information using<br />
sensing technologies and terminals.<br />
Let me describe further the feature of such a society by<br />
using the analogy of the past communication by physical<br />
mail vs. telephone: our past communication way used to be<br />
mainly a letter, which implies his or her own character and<br />
mindset. Hand writing letter takes days and hours to be<br />
delivered by a postman to the addressee, But I believe such a<br />
215
letter communication way is meaningful and will remain.<br />
Since then information communication has been mainly<br />
shifted from letter to telephone, from telephone to keyboard,<br />
from analog to digital.<br />
However, in the society of rapid progress in science and<br />
technology, I’m afraid that we have been facing the fact that<br />
we may not be able to communicate with others<br />
satisfactorily and properly by using a portable mailing<br />
equipment. It’s difficult for modern communication to<br />
convey human being’s vital information such as why your<br />
heart is so beating today, or your face is blushing when you<br />
send e-mail to your boy friend.<br />
Differently speaking, that is evidence of human being’s<br />
vital information is potentially too rich and sensitive to be<br />
measured by digital devices. Human being’s vital<br />
information remains undeveloped so far, and we are now<br />
beginning to recognize a human being as a sensor itself<br />
united with mind and body.<br />
I envision the future information communication way as<br />
follows; Apart from inputting your vital information into the<br />
keyboard by the help of manpower, all you need is to put a<br />
small sensing –chip on your chest. Thus it can read your<br />
potential vital information and automatically send it to other<br />
people by way of wearable computer. As the next step, I’m<br />
planning to build a sensor-networking world that can weave<br />
every vital information sent by each wearable computer. In<br />
our aging society, I’m sure that a lot of elderly people hope<br />
they could have good information on their own vital<br />
(physical and mental) conditions and, if necessary, they<br />
could be cared and cured timely with minimum of helping<br />
power.<br />
We’ve been now challenging to build such a sensor<br />
networking society in the near future.<br />
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<br />
[4] H.Kubota,“Collective Theory of Vital Signs” -To promote<br />
life-saving medicine-, Sinko Trading Co., LTD., Tokyo,<br />
Japan,pp.8,40,83-90,113-115.2006.<br />
[5] Kiyoshi Itao, “Wearable sensor network connecting artifacts, nature,<br />
and human being,” KEYNOTE PRESENTATION, in Proc. of the<br />
Sixth IEEE SENSORS 2007 Conference, Atlanta, pp. 1120–1123,<br />
2007<br />
[6] Kiyoshi Itao, “Human Recorder System Development for Sensing<br />
the Autonomic Nervous System,” in Proc. of the Seventh IEEE<br />
SENSORS 2008 Conference, Lecce, pp. 423–426, 2008<br />
VIII. CONCULUSION<br />
All creations send out information. Up to the current<br />
information technology and communication service stage,<br />
however, all information is not yet received nor<br />
communicated well.<br />
In this paper, we propose the formation of society where<br />
all information is received and communicated well.<br />
Going up to the next stage from where we are now, I<br />
believe that “wearable technology”, “sensing technology”,<br />
and “sensor communication” could be the key. And we,<br />
human beings, will play a crucial role in what I call “sensor<br />
network.” As the development of wearable technology and<br />
wearable devices, we could be an interface to all creations.<br />
That is the world of Nature Interface.<br />
REFERENCES<br />
[1] K.Itao,“Micromechatronics Technology for Wearable Information<br />
and Communication Equipment”, Sens. and<br />
Materials,vol.10,No.6,pp.325-335,1985.<br />
[2] K.Itao,“Next-Generation Information and Communication<br />
Equipment based on Microsystems Technologies”, International<br />
Journal of The Japan Society for<br />
Engineering,vol.31,No3,pp.167-171,Sep.1997.<br />
[3] K.Itao,“Micromechatronics for wearable Information systems”,<br />
Journal of Micromechatronics,vol.1, No.1,pp.5-13,2000.<br />
216
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<br />
Design, Fabrication, and Integration of<br />
Piezoelectric MEMS Devices for Applications in<br />
Wireless Sensor Network<br />
Jian Lu*, Yi Zhang, Toshihiro Itoh, Ryutaro Maeda<br />
Research Center for Ubiquitous MEMS and Micro Engineering (UMEMSME),<br />
National Institute of Advanced Industrial Science and Technology (AIST), Namiki 1-2-1, Tsukuba, Ibaraki, 305-8564, Japan<br />
Abstract- One of the competitive solutions to expand the<br />
function of microelectromechanical system (MEMS) is the<br />
integration of piezoelectric lead zirconate titanate (PZT) thin<br />
films for device self-actuation at low driving voltage, device<br />
self-sensing with low power consumption, as well as for energy<br />
harvesting. However, up-to-date, difficulties still exist not only<br />
in PZT film preparation but also in PZT film integration with<br />
other MEMS components and ICs. This paper therefore<br />
presents our recent progress on large area deposition, fine<br />
pattern etching, and low temperature bonding of PZT thin<br />
films for wafer scale PZT film integration and piezoelectric<br />
MEMS application. The energy dissipation mechanism in<br />
piezoelectric MEMS devices was also discussed to optimize the<br />
device structure for the pursuit of better performance.<br />
Ultra-sensitive micro cantilever and disk resonator with<br />
on-chip piezoelectric PZT transducers were presented herein as<br />
an exploratory application of piezoelectric MEMS devices in<br />
distributed wireless sensor network.<br />
I. BACKGROUND<br />
To deal with population aging, environmental pollution,<br />
global warming, and other modern society problems,<br />
microelectromechanical system (MEMS) is significantly<br />
important because various applications, such as human<br />
healthcare, food safety, environmental monitoring, animal<br />
watching, green manufacturing, mechanical structure<br />
monitoring, and smart living, can be realized by MEMS and<br />
distributed wireless sensor network (WSN) technology with<br />
high sensitivity, low cost, and low power consumption<br />
[1]-[4].<br />
To expand the function of MEMS for applications in<br />
WSN, the integration of various materials or components for<br />
device actuation and sensing is essential. One of the most<br />
competitive materials is the piezoelectric lead zirconate<br />
titanate (Pb(Zr x ,Ti 1-x )O 3 , PZT) thin film because PZT is a<br />
high energy density material which scales very favorably<br />
upon miniaturization [5]. The piezoelectric coefficient d 33<br />
and dielectric constant ɛ of the PZT film was reported as<br />
high as 143 pC/N and 1310 respectively, which is one order<br />
higher than that of zinc oxide (ZnO) film (d 33 =11 pC/N,<br />
ɛ =11) and aluminum nitride (AlN) film (d 33 =3.4 pC/N,<br />
ɛ =10.4). Besides, the well-integrated PZT film on MEMS<br />
devices can be used not only for device self-actuation at low<br />
driving voltage and device self-sensing with low power<br />
consumption, but also can be used for energy harvesting.<br />
However, up-to-date, difficulties still exist in PZT<br />
preparation and PZT integration with other MEMS<br />
components and ICs. It is mostly due to the high temperature<br />
annealing process and the residual stress of the film, as well<br />
as the energy dissipation of the piezoelectric MEMS devices<br />
[6]-[8]. Moreover, the high volume mass production and<br />
commercialization of MEMS have been expected for many<br />
years since IC industry went to the well-developed stage. The<br />
bottlenecks, which discourage MEMS industry to advanced<br />
steps, are the difficulties in integration of MEMS<br />
components with ICs, the yields, and the cost. Especially to<br />
piezoelectric MEMS, the fabrication, integration of the<br />
piezoelectric PZT thin film, and the design, optimization of<br />
the device structure need great efforts not only from the<br />
technical point of view but also through innovative academic<br />
research.<br />
We have engaged in large-area deposition, fine pattern<br />
etching, low temperature bonding of PZT thin films for<br />
MEMS application, and energy dissipation mechanism of the<br />
piezoelectric MEMS devices for the pursuit of better device<br />
performance for many years. This paper thus presents our<br />
recent progress of above work. Moreover, the design,<br />
fabrication and evaluation of ultra-sensitive micro<br />
cantilevers and disk resonators, which has on-chip<br />
PZT-electrode stacks as the transducer, were presented in<br />
this paper as an exploratory application of piezoelectric<br />
MEMS devices in distributed wireless sensor network for<br />
human healthcare, environmental monitoring, and other<br />
applications.<br />
II.<br />
RESULTS AND DISCUSSION<br />
2.1 Large area PZT film deposition by sol-gel process<br />
For PZT preparation in large area, residual stress<br />
frequently leads to wafer-warpage, PZT film delaminating,<br />
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<br />
Fig. 1 AES results indicate the inter-diffusion between different layers along<br />
depth (sputter time) of the PZT/Pt/Ti/SiO 2/Si wafer.<br />
Fig. 2 The balance layer not only enables large wafer scale thick crack-free<br />
PZT preparation but also reduces wafer-warpage.<br />
or PZT film cracking. Wafer-warpage results in large<br />
miss-alignment between photomask and the wafer in<br />
following photolithography process. PZT delaminating or<br />
cracking places a limitation on film size, thickness, and<br />
likely to deteriorates film properties. It is an intractable<br />
problem for large area integration and mass production of<br />
piezoelectric MEMS devices.<br />
Fig. 1 shows element distribution along the depth of the<br />
PZT/Pt/Ti/SiO 2 /Si wafer measured by Auger Electron<br />
Spectroscopy (AES). The PZT film was prepared by sol-gel<br />
process with the thickness of 1 μm (8 layers, annealed by<br />
RTA at the temperature of 650°C for 3 min after each layer<br />
deposition). Pt/Ti with the thickness of 200/50 nm was used<br />
as bottom electrode of the PZT film. The results in Fig. 1<br />
suggested that the large residual stress is mainly caused by<br />
the diffusion of Ti into Pt in Pt/Ti bottom electrode and the<br />
migration of Pb from PZT into substrate during PZT high<br />
temperature annealing process.<br />
Our results also indicated that another reason for PZT<br />
delaminating or cracking is the thermal stress due to thermal<br />
expansion coefficient mismatch between PZT, Pt/Ti bottom<br />
electrode, and SiO 2 buffer layer (thermal expansion<br />
coefficient of PZT: 4.03×10 -6 /K; thermal expansion<br />
coefficient of Pt: 14.2×10 -6 /K; thermal expansion coefficient<br />
of SiO 2 : 0.4×10 -6 /K). Therefore, we successfully prepared<br />
high quality crack-free piezoelectric film in large area by<br />
sol-gel process. In this process, another Pt/Ti film, which was<br />
deposited on backside of the wafer by the same process as<br />
Pt/Ti bottom electrode, was used to balance the residual<br />
stress cause by the inter-diffusion of Ti into Pt in Pt/Ti<br />
bottom electrode and the thermal expansion coefficient<br />
mismatch between Pt/Ti bottom electrode and SiO 2 . The<br />
process details were published elsewhere [9]. As<br />
demonstrated in Fig. 2, wafer-warpage was dramatically<br />
reduced by this approach, which is essential for large area<br />
fabrication of piezoelectric MEMS devices [10].<br />
2.2 PZT fine pattern fabrication by ICP-RIE<br />
Fabrication of the PZT-electrode fine pattern with the<br />
feature size of less than 10 microns is the next essential step<br />
for piezoelectric MEMS application after large area PZT film<br />
deposition. Wet chemical etching is often used as a low-cost<br />
process in MEMS fabrication. However, as reported in<br />
literature, ferroelectric and piezoelectric properties of the<br />
PZT film will be markedly degraded by the diffusion of<br />
hydrogen atoms from HNO 3 /HF etchant into PZT film [11].<br />
Moreover, because of the intrinsic large undercutting defect<br />
of the wet chemical etching process, it is hard to be used<br />
when feature size of the PZT-electrode stack is less than 10<br />
microns, which are expected by most MEMS devices for the<br />
pursuit of high integration density, low cost, and high device<br />
performance.<br />
Fig.3 Etching-rate of PZT, Pt and photoresist (S1830) by ICP-RIE at various<br />
Ar concentrations in Ar/SF 6 mixed gas. The insert shows an obtained<br />
PZT-electrode fine pattern with the feature size of 2 μm.<br />
218
In our work, we proposed an innovated dry process for<br />
PZT-electrode fine pattern fabrication [12]. This process use<br />
conventional ICP-RIE system and Ar/SF 6 mixed gas as the<br />
etchant. AFM images of the PZT film before and after the<br />
dry-etching revealed that the grain-boundary and the shape<br />
of the PZT crystallites were more identifiable when using<br />
lower Ar concentration in Ar/SF 6 mixed etchant. It suggested<br />
that the etching is reactive-physical combined process. As<br />
shown in Fig. 3, the highest PZT etching-rate (58 nm/min)<br />
and the best etching-selectivity (PZT:Pt=1.14) was achieved<br />
at 66.7% of Ar in Ar/SF 6 mixture. We also confirmed that the<br />
remanent polarization and the dielectric constant of the PZT<br />
film were 16.5 μC/cm 2 and 1019 before the etching, 15.5<br />
μC/cm 2 and 1013 after the etching. It demonstrated that the<br />
proposed dry-etching process did not degrade the properties<br />
of the PZT film. As shown in insert of Fig. 3, a<br />
PZT-electrode fine pattern with the feature size of 2 μm was<br />
successfully obtained by this process.<br />
2.3 Low temperature bonding of PZT film<br />
For most of the MEMS and IC devices, process<br />
temperature of more than 400 °C is fatal. However, to obtain<br />
well-crystallized and (100)-oriented PZT film by sol-gel<br />
method, high annealing temperature in the range of 600–750<br />
°C is usually required. Even the transforming temperature of<br />
PZT perovskite-phase (~530 °C) is beyond that of most<br />
MEMS and ICs can withstand. Although adding modifiers<br />
into PZT solution [13], or using seeding layers to enhance<br />
PZT nucleation [14] has been reported effective to reduce the<br />
PZT annealing temperature, their benefits to piezoelectric<br />
MEMS fabrication are limited. Z.Wang et al. proposed the<br />
bonding of bulk PZT with silicon wafer, and then thin down<br />
the PZT to less than 10 μm by using chemical mechanical<br />
polishing [15]. This method is complicated, time-consuming,<br />
and high-cost. Therefore, it is hard to be used for MEMS<br />
mass production.<br />
Ti: 50 nm<br />
Cr: 50 nm<br />
PZT film: 1 μm<br />
Pt/Ti: 200/50 nm<br />
SiO 2 : 2 μm<br />
Si: 500 μm<br />
Si: 400 μm<br />
4-inch<br />
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<br />
Au film 100/300 nm<br />
Au film: 300 nm<br />
film low temperature integration on silicon substrate. We use<br />
Au film as the intermediate layer. Fig. 4 shows schematic<br />
view of the sample and its SAM image after bonded at 150<br />
°C. Die-shear test revealed that the bonding strength is ~75<br />
MPa at the bonding pressure of 325 MPa [16]. This<br />
technique effectively avoids the damage of PZT high<br />
temperature annealing to other MEMS components and ICs.<br />
It also enables the fabrication of piezoelectric material,<br />
piezoelectric MEMS components, and ICs separately by<br />
different wafers, and then bonded them together for<br />
integrated MEMS/ICs mass production.<br />
2.4 Energy dissipation in piezoelectric MEMS device<br />
Piezoelectric MEMS device offers great self-actuation<br />
and self-sensing capabilities, but it always suffers from low<br />
device performance in sensitivity. To investigate the energy<br />
dissipation mechanism, various cantilevers with different<br />
layers are designed and fabricated, which includes SiO 2<br />
elastic layer, Pt/Ti bottom electrode layer, PZT film, Ti/Pt/Ti<br />
upper electrode layer, and SiO 2 top electric passivation layer.<br />
The measured quality-factors (Q-factor) of those cantilevers<br />
were analyzed by theoretical calculation. Fig. 5 summarized<br />
the difference between measured Q-factors and theoretical<br />
calculated Q-factors. It clearly revealed that the energy<br />
dissipation by the PZT film and the multi-layered structure is<br />
extremely large. It is comparable to air dumping under<br />
atmospheric pressure and becomes dominating under<br />
reduced pressures [17].<br />
Q-factor decreases to calculated value (%)<br />
10<br />
0<br />
-10<br />
-20<br />
-30<br />
-40<br />
-50<br />
-60<br />
-70<br />
SiO 2<br />
SiO 2<br />
+Ti/Pt<br />
SiO 2<br />
+Ti/Pt+PZT+Ti/Pt/Ti+SiO 2<br />
SiO 2<br />
+Ti/Pt+PZT<br />
Length: 200 μm<br />
Length: 250 μm<br />
Length: 300 μm<br />
1 2 3 4 5<br />
Number of structure layer<br />
Fig.5 Q-factor of the cantilever decreased with the increasing of structure<br />
layers, especially after PZT film integration.<br />
Fig. 4 PZT bonding on silicon substrate by surface activated bonding (SAB)<br />
using Au film as the intermediate layer: schematic view of the sample (left)<br />
and SAM images of the sample bonded at 150 °C (right).<br />
To promote the PZT application in case >400 °C<br />
temperature cannot been used, surface activated bonding<br />
(SAB) was introduced in our work for the first time for PZT<br />
Based on above results, a piezoelectrically-actuated<br />
micro cantilever [18] and a piezoelectrically-transduced disk<br />
resonator [19] were designed and fabricated in our work as<br />
resonant-based ultra-sensitive mass sensor for human<br />
healthcare and other applications. The device fabrication was<br />
done by 4-inch SOI wafers using above large area PZT film<br />
219
deposition and PZT fine pattern dry-etching techniques. Fig.<br />
6 shows SEM images of the fabricated cantilever (Fig. 6 (a))<br />
and the disk resonator (Fig. 6 (b)).<br />
In cantilever as shown in Fig. 6 (a), two PZT actuators<br />
were arranged symmetrically on both sides of the silicon<br />
cantilever and connected to the cantilever via thin beams<br />
near to substrate for cantilever excitation. Then the<br />
piezoelectric PZT actuator could be separated from the<br />
resonant structure to compress the energy dissipation from<br />
PZT film and the multi-layered structure. To compress<br />
negative effects from residual stress, support beams were<br />
designed at the front end of the actuator to reduce actuator’s<br />
initial bending. Another purpose of the support beam is to<br />
limit actuator’s vibration amplitude at the resonant frequency<br />
to suppress energy dissipation. A piezoresistive<br />
Wheatstone-bridge-gauge was integrated at the fixed end of<br />
the cantilever to detect its vibration.<br />
(a)<br />
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<br />
The PZT-electrode stacks were kept far from support beams<br />
to avoid clamped energy loss. To reduce the effects of<br />
support beam on disk vibration as well as to compress<br />
clamped energy loss, support beams were designed as thin as<br />
possible (7 μm×15 μm) after finite element analysis (FEA)<br />
using ANSYS ® . An Au/Cr layer with the thickness of<br />
200/50 nm was deposited on both cantilever and disk surface<br />
as the mass adsorption area to demonstrate its application as<br />
a mass sensor.<br />
2.5 Device evaluation and application<br />
After fabrication, the devices were wire-bonded and<br />
then packaged for evaluation. The results demonstrated that<br />
the cantilever (length: 100 μm; width: 30 μm) has excellent<br />
Q-factor of 1113 in air, which is several times higher than<br />
latest reported Q-factor of other integrated micro cantilevers<br />
[20][21]. Under reduced pressure of about 30 Pa, Q-factor of<br />
the cantilever was as high as 7279. Fig. 7 shows measured<br />
equivalent capacitance Cs values of the PZT film on disk<br />
resonator (Fig. 6 (b)). Clearly, the Cs variation was 0.2~0.3%<br />
at the resonant frequency owing to its vibration-induced<br />
piezoelectric charge. The disk shows great signal to noise<br />
ratio besides its high Q-factor (~1300 in air). It is also<br />
noteworthy that an electric voltage of 0.2~1 volt was proved<br />
sufficient for cantilever and disk actuation, which improves<br />
its integration capability from the viewpoints of power<br />
supplies and power consumption.<br />
(b)<br />
PZT-electrode<br />
stack<br />
Silicon Disk<br />
Fig. 7 Piezoelectric induced output (equivalent capacitance Cs) of a<br />
fabricated piezoelectric disk resonator.<br />
Fig. 6 SEM images of the fabricated (a) micro cantilevers actuated by PZT<br />
thin film and (b) disk resonator transduced by PZT thin film.<br />
In disk as shown in Fig. 6 (b), PZT-electrode stacks with<br />
limited size to reduce its energy dissipation were integrated<br />
on surface of the disk for both disk actuation and sensing.<br />
Various piezoelectric MEMS devices are expected to be<br />
integrated in sensor network for ubiquitous applications due<br />
to its self-actuation at low driving voltage, device<br />
self-sensing with low power consumption, as well as its<br />
energy harvesting capabilities. Fig. 8 explains concept of a<br />
human healthcare system by wireless sensor network<br />
technology. Although lots of work must be done, we believe<br />
it will come to reality soon.<br />
220
Fig. 8 Concept of the in-home personal healthcare system using<br />
piezoelectric MEMS resonator as the ultra-sensitive gas sensor.<br />
III. CONCLUSIONS<br />
Piezoelectric PZT is expected as one of the key<br />
functional materials for MEMS industry. Although much<br />
knowledge is available from state-of-the-art studies for wafer<br />
scale PZT film deposition, etching, characterization, and<br />
large-scale piezoelectric devices fabrication, the mass<br />
production and the commercialization of piezoelectric<br />
MEMS are still complex and difficult at present.<br />
This paper reviewed our recent progress on large area<br />
deposition, fine pattern etching, and low temperature<br />
bonding of PZT thin films for wafer scale fabrication of<br />
piezoelectric MEMS devices from both academic and<br />
technical point of view. The energy dissipation mechanism<br />
in piezoelectric MEMS device and the structure optimization<br />
are also discussed and clarified for the pursuit of better<br />
device performances.<br />
For practical mass production and commercialization,<br />
further works are still undergoing. It includes (1) developing<br />
standard processes for piezoelectric materials fabrication on<br />
8~12-inch wafer; (2) 3D wafer level packaging between<br />
piezoelectric processed wafers and other MEMS processed<br />
wafers for the promotion of large area integration and mass<br />
production; and (3) improving the reliability, stability as well<br />
as performances of the piezoelectric MEMS devices. We will<br />
report the latest results and other details in our following<br />
publications soon.<br />
ACKNOWLEDGMENT<br />
This research is partially supported by the Japan Society<br />
for the Promotion of Science (JSPS) through its “Funding<br />
Program for World-Leading Innovative R&D on Science and<br />
Technology (FIRST Program)."<br />
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<br />
REFERENCES<br />
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potential impact on surgery, World Journal of Surgery 2001, 25(11),<br />
1458-1466.<br />
[2] S.Junnila, H.Kailanto, J.Merilahti et al.: Wireless, multipurpose in-home<br />
health monitoring platform: two case trials, IEEE Transactions on<br />
Information Technology in Biomedicine 2010, 14(2), 447-455.<br />
[3] R.K.Das, R.K.Garg: Global environmental microelectromechanical<br />
systems sensors: Advanced weather observation system, Defence<br />
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[4] M.Hautefeuille: C.O’Mahony; B.O’Flynn et al., A MEMS-based<br />
wireless multisensor module for environmental monitoring,<br />
Microelectronics Reliability 2008, 48(6), 906-910.<br />
[5] P.Muralt: Ferroelectric thin films for micro-sensors and actuators: a<br />
review, Journal of Micromechanics and Microengineering 2000, 10,<br />
136–146.<br />
[6] D.L.Polla, P.J.Schiller: Integrated ferroelectric microelectromechanical<br />
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[7] H.Raeder, F.Tyholdt, W.Booij et al.: Taking piezoelectric microsystems<br />
from the laboratory to production, Journal of Electrocermics 2007,<br />
19(4), 357-362.<br />
[8] S.Tadigadapa, K.Mateti: Piezoelectric MEMS sensors: state-of-the-art<br />
and perspectives, Measurement science & Technology 2009, 20(9),<br />
092001.<br />
[9] J.Lu, T.Kobayashi, Y.Zhang et al.: Wafer scale lead zirconate titanate<br />
film preparation by sol-gel method using stress balance layer, Thin<br />
Solid Films 2006, 515(4), 1506-1510.<br />
[10] J.Lu, Y.Zhang, T.Kobayashi et al.: Preparation and characterization of<br />
wafer scale lead zirconate titanium film for MEMS application,<br />
Sensors and Actuators A: Physical 2007, 139, 152-157.<br />
[11] T.Kobayashi, M.Ichiki, R.Kondou et al.: Degradation in the<br />
ferroelectric and piezoelectric properties of Pb(Zr,Ti)O3 thin films<br />
derived from a MEMS microfabrication process, Journal of<br />
Micromechanics and Microengineering 2007, 17(7), 1238-1241.<br />
[12] J.Lu, Y.Zhang, T.Ikehara et al.: Inductively coupled plasma reactive ion<br />
etching of lead zirconate titanate thin films for MEMS application,<br />
IEEJ Transactions on Sensors and Micromachines 2009, 129(4),<br />
105-109.<br />
[13] W.G.Zhu, Z.H.Wang, C.L.Zhao et al.: Low temperature processing of<br />
nanocrystalline lead zirconate titanate (PZT) thick films and ceramics<br />
by a modified sol-gel route, Japanese Journal of Applied Physics 2002,<br />
41, 6969-6975.<br />
[14] H.Suzuki, S.Kaneko, K.Murakami et al.: Low-temperature processing<br />
of highly oriented Pb(ZrxTi1-x)O3 thin film with multi-seeding<br />
layers, Japanese Journal of Applied Physics 1997, 36, 5803-5807.<br />
[15] Z.H.Wang, J.M.Miao, C.W.Tn et al.: Fabrication of piezoelectric<br />
MEMS devices-from thin film to bulk PZT wafer, Journal of<br />
Electroceramics 2010, 24(1), 25-32.<br />
[16] Y.H.Wang, J.Lu, T.Suga: Low Temperature Wafer Bonding Using<br />
Gold Layers, In Proc. 2009 International Conference on Electronics<br />
Packaging Technology & High Density Packaging (ICEPT-HDP),<br />
pp.15A-2-1, Beijing, China, Aug. 2009<br />
[17] J.Lu, T.Ikehara, Y.Zhang et al.: Energy Dissipation Mechanisms in<br />
Lead Zirconate Titanate Thin Film Transduced Micro Cantilevers,<br />
Japanese Journal of Applied Physics 2006, 45(11), 8795-8800.<br />
[18] J.Lu, T.Ikehara, Y.Zhang et al.: High Quality Factor Silicon Cantilever<br />
Driven by PZT Actuator for Resonant Based Mass Detection,<br />
Microsystem Technologies 2009, 15(8), 1163-1169.<br />
[19] J.Lu, T.Suga, Y.Zhang et al.: Micromachined Silicon Disk Resonator<br />
Transduced by Piezoelectric Lead Zirconate Titanate Thin Films,<br />
Japanese Journal of Applied Physics 2010, 49, 06GN17.<br />
[20] D.Jin, X.Li, J.Liu et al.: High-mode resonant piezoresistive cantilever<br />
sensors for tens-femtogram resoluble mass sensing in air, Journal of<br />
Micromechanics and Microengineering 2006, 16, 1017-1023.<br />
[21] Z.Shen, W.Y.Shih, W.H.Shih: Self-exciting, self-sensing<br />
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femtogram/Hertz ssensitivity, Applied Physics Letters 2006, 89,<br />
023506.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
Novel MEMS digital temperature sensor for wireless<br />
avian-influenza monitoring system in poultry farm<br />
Yi Zhang*, Hironao Okada, Takeshi Kobayashi, Toshihiro Itoh<br />
National Institute of Advanced Industrial Science and Technology (AIST)<br />
1-2-1, Namiki, Tsukuba<br />
Ibaraki 305-8564, Japan<br />
Abstract- This paper presents our recent progress on the<br />
development of sensor node for wireless avian-influenza<br />
monitoring system in poultry farm. It was found that the<br />
influenza infection could be detected based on the temperature<br />
and activity monitoring at a very early stage. In order to meet<br />
the requirements of lower power consumption and high<br />
sensitivity, new micro temperature sensor technology was<br />
developed in this paper. The new temperature sensor consists of<br />
array of bimorphs that response to different temperature by<br />
the mechanism of event-driven on/off. Difference kinds of<br />
bimorph configuration and structure were developed and<br />
experimentally examined. It was found that 3D bimorph<br />
showed attractive advantages relative to traditional planar<br />
configuration including easy-to-package, compact and<br />
easy-to-fabrication. Wafer-scale 3D microfabrication is also<br />
established for the 3D bimorph.<br />
should give the alarm at the very beginning of the infection;<br />
otherwise, there would be still high risks of influenza<br />
pandemic. The modern poultry farms in Japan usually have<br />
more than 50,000 chickens so that only several percentages<br />
of them could be affordable with wireless sensor nodes.<br />
Considering the limitation of system cost, it is also difficult<br />
to integrate those advanced functions such as direct virus<br />
detects in to the. Therefore, first of all, it is necessary to<br />
determine the minimum functions but enough for the<br />
practical monitoring. It is well known that the health state of<br />
chicken and other animals could be monitored by the<br />
variation of body temperature and individual activity. It is<br />
thus interesting to investigate whether it is possible to<br />
establish a high performance avian influenza monitoring<br />
system only based on the monitoring of body temperature<br />
and activity.<br />
I. INTRODUCTION<br />
With rapidly growing threats from avian-influenza in<br />
recent years [1], there is considerable interest in the<br />
development of wireless health-monitoring system<br />
technology for poultry farm. It is in particular important for<br />
Asia and other areas where the safety of the poultry and<br />
products are important for daily living of the local society. It<br />
is thought that the wireless health-monitoring system can<br />
reduce the public risk and thereby economic lost of the<br />
avian-influenza to the least. However, there are few<br />
literatures on the application of wireless network and sensor<br />
nodes in poultry farm. It is necessary to establish basic<br />
knowledge and requirements on the sensor node and the<br />
wireless system. For example, the wireless<br />
health-monitoring system consists of large number of sensor<br />
nodes because a modern poultry farm has several tens<br />
hundreds of poultry animals and even more. In addition, the<br />
poultry animals are small so that those sensor nodes should<br />
be very tiny and light weight. This paper is aiming to present<br />
our recent progress on the development of the wireless<br />
sensor node.<br />
II.<br />
DESIGN AND EXPERIMENTALS<br />
A. Prototype wireless sensor node<br />
We thought that an avian influenza monitoring system<br />
Fig. 1 Photograph of the prepared prototype wireless sensor node.<br />
We have developed prototype wireless sensor node and<br />
used it for a small scale animal experiments in order to<br />
collect those required information. Figure 1 show its<br />
photography. A thermistor is used for measuring the body<br />
temperature. A 3-axis accelerometer (H34C, Hitachi Metals<br />
Ltd.) is used for the activity sensing. The weight and size of<br />
the chip is 1.2 g and φ18×3 mm, respectively. The sensor<br />
node is protected by a plastic case (φ24×9 mm, 2.2 g). The<br />
total weight of the prototype node is only about 5.2 g<br />
including a button battery. The work distance is about 20 m.<br />
More details could be found in our recent reports [2-3]. The<br />
experimental and simulation results indicated that the<br />
influenza infection including the highly pathogenic avian<br />
influenza (HPAI) could be detected by the wireless sensor<br />
node at 10 hours and earlier before the death. It was also<br />
222
found that the early-stage diagnosis and alarm could be<br />
realized through the health monitoring of the 5% of the<br />
chickens. It could be drawn the conclusion that that the avian<br />
influenza monitoring could be realized by the temperature<br />
and activity sensing. However, owing the capacity limitation<br />
of available battery, the work lifetime of the prototype sensor<br />
node does not meet the practical application, in which<br />
two-year lifetime is required. Much effort had been made on<br />
the development of ultra-low power consumption circuit and<br />
event driven on/off accelerometer technologies but few<br />
detailed literatures are available on the development of<br />
MEMS-based temperature sensor with ultra-low power<br />
consumption. This paper would give a detailed report on our<br />
recent progress on the development.<br />
B. Digitally sensing of temperature<br />
Conventional temperature sensors require high power<br />
consumption for sensing and transferring analog signals into<br />
digital ones so that they do not meet the low-power<br />
requirements [4-6]. Among those sensing structures of<br />
temperature, bimorph structure is attractive because of its<br />
passive sensing mechanism. As the bimorph structure is<br />
directly driven by the thermal expansion mismatch of<br />
constitution layers, its power consumption is intrinsically<br />
zero. Therefore, up to now, it has been widely utilized as a<br />
powerless switch to protect system from overheating and<br />
thereby failures. Since the response temperature of one<br />
bimorph can be directly controlled by the adjusting of the<br />
beam length or other structure parameters, it is possible to<br />
realize the temperature sensing in a wide range by using a<br />
group of bimorph structures. If there is an array of several<br />
bimorphs that could be response to different temperature<br />
variation, i.e. event driven on/off, a passive but high<br />
sensitivity temperature sensor could be possible. Therefore,<br />
we have suggested two kinds of new bimorph structures and<br />
examined their thermal behavior with the temperature<br />
variation.<br />
Figure 2 (a) is schematic of triple-beam bimorph. It would<br />
deflect up with the increasing of temperature and contact the<br />
counter pads so that other application circuits can be<br />
switched on. The lengths of the side beam and the middle<br />
beam were determined for different response temperatures<br />
through finite element simulation by using ANSYS software.<br />
The simulation shows that the side beams assure 0.5 o C<br />
sensitivity and the middle beam assists in achieving 0.1 o C<br />
sensitivity. Figure 3 shows the simulation results. It is<br />
noteworthy that the length of the middle beam can be<br />
adjusted by the range of 20-30 μm for achieving about 1 o C<br />
difference. The variation is about 10% of the total beam<br />
length so that the microfabrication of the bimorph array<br />
would be feasible.<br />
Figure 2 (b) is schematic of 3D-bimorph. The 3D bimorph<br />
is beneath the substrate surface and its bending direction is<br />
not perpendicular to the top surface. With temperature<br />
changing, the bimorphs bends to or away from each other for<br />
switch on and off, respectively. The sensitivity can therefore<br />
be determined by the gap increment instead of the length<br />
increment. Calculation shows that a sensitivity of about<br />
0.5 o C could be achieved by using 3 μm Si/0.3 μm Au<br />
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May 2011, Aix-en-Provence, France<br />
<br />
bimorph with a gap increment of 0.5 μm. The gap increment<br />
could be realized by using conventional MEMS process.<br />
Compared with the triple-beam layout, the 3D bimorph can<br />
be easily packaged and integrated with other chip<br />
components but its microfabrication needs more effort.<br />
Beam Length, μm<br />
Bimorph deflects up and is switched on.<br />
Fig. 2 (a) Schematic of triple-beam bimorph and its work mechanism.<br />
Fig. 2 (b) Illustration of 3-D metal/Si bimorph.<br />
280<br />
270<br />
260<br />
250<br />
240<br />
230<br />
220<br />
210<br />
200<br />
190<br />
38 40 41 42<br />
Middle beam, L 1<br />
Side beam, L 2<br />
C. Fabrication and characterization<br />
Figure 4 is the fabrication sequence of the triple-beam<br />
43<br />
Response temperature, o C<br />
Fig. 3 Length of the side and middle beams determined by ANSYS<br />
simulation for different temperature. The displacement of bimorph<br />
and the thickness of metal layer are assumed to be 5 μm and 500 nm,<br />
respectively.<br />
45<br />
223
imorph. Bimorph array of single beam was also prepared<br />
for direct comparisons. The fabrication is mainly involved of<br />
commonly-used surface micromachining and deep reactive<br />
ion etching technology. The top and bottom layer of the<br />
bimorph was Mo (500 nm-thick) and Au (500 nm-thick),<br />
respectively. The former has the thermal expansion<br />
coefficient of about 4.9 ppm/ o C. The latter has the thermal<br />
expansion coefficient of about 14.4 ppm/ o C. Figure 5 is<br />
fabrication process of the 3D bimorph, in which SOI wafer<br />
was used and its device layer has high resistivity (> 1000<br />
Ω⋅cm). Spray coating method was used for the resist coating<br />
in order to get good coverage on non-planar surface and the<br />
fabrication of isolation gap. Other key processes include the<br />
conformal deposition of thin metal film and the formation of<br />
the isolation structure between the adjacent bimorph.<br />
Because the beam height was 35 μm, the sputtering method<br />
was utilized for the deposition of thin metal film.<br />
Fig 4 Fabrication sequence of the triple-beam bimorph<br />
Fig. 5 Fabrication sequence of the 3D bimorph.<br />
The displacement measurement of the triple-beam<br />
bimorph was carried out in a home-made quartz stage by<br />
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May 2011, Aix-en-Provence, France<br />
<br />
using a confocal scanning laser microscopy (OPTELICS<br />
S130, LASERTEC). The tip displacements were in-situ<br />
measured. The 3D bimorph was examined on a Cascade<br />
9000 analytical probe station with a heater (Temptronic Co.)<br />
by using Agilent 4284A LCR meter.<br />
III. RESULTS AND DISCUSSION<br />
Figure 6 are the SEM images of the prepared triple-beam<br />
bimorphs. Traditional single-beam bimorph was also<br />
prepared for direct comparisons of thermal response<br />
behavior. The bi-metal layer structure is visible in Fig. 6. It<br />
is noteworthy that all the prepared bimorphs have large<br />
initial deflection because of the residual stress that resulted<br />
from the deposition and micromachining of the thin films.<br />
The initial deflection was within 10 ~ 30 μm. The<br />
triple-beam bimorph had almost same initial bends as the<br />
single-beam or traditional bimorph. It was also found that<br />
the initial bends of the triple-beam bimorph was not the<br />
average value of those of the side and middle beam. The side<br />
beam was dominantly determined the initial bends of the<br />
triple-beam bimorph. Figure 7 representative Z-images for<br />
the displacement measurement of the triple-beam bimorph.<br />
The triple-beam bimorph had tilted during the upward<br />
deflection upon temperature increasing, which possibly<br />
resulted from alignment error during the fabrication. The<br />
initial bends and tilts of the triple-beam bimorph would be<br />
barriers to the practical application. As well, the package of<br />
the triple-beam bimorph would be difficult.<br />
Figure 8 are the SEM images of the prepared 3D bimorphs,<br />
respectively. It consisted of ten 5 μm-thick Si/0.4 μm-thick<br />
Au bimorph in five pairs. Good leading connections were<br />
formed across the cavity edge. The isolation gaps were well<br />
formed, too. Figure 9 was optical photo of the chip after the<br />
dicing. The chip was 1.2 mm square. The 3-D bimorph was<br />
robust and can stand for conventional dicing process with<br />
simple protection by polymer sheet. We could draw the<br />
conclusion that the new 3D layout could simplify the<br />
microfabrication process and thus the cost.<br />
Figure 10 is the measured displacements of the prepared<br />
bimorph upon the increasing of temperature from 25 o C to<br />
38 o C, 41 o C and 44 o C. The triple-beam bimorph exhibited<br />
different behaviors from the traditional single-beam<br />
bimorph upon the increasing of temperature. Firstly, the<br />
former had larger tip displacements than the latter. For<br />
example, the tip displacement was about 2 μm for the<br />
triple-beam bimorph with the temperature increasing from<br />
41 to 44 o C while the other was only about 0.5 μm. The larger<br />
tip displacement makes bigger increment of dimension and<br />
therefore higher sensitivity possible. For example, the<br />
triple-beam thermometer could be consisted of 20 bimorphs<br />
at the displacement increment of 100 nm for temperature<br />
sensing between 41 and 44 o C. Larger displacement and<br />
beam-length increment also suggested that the fabrication<br />
process was more feasible. We could draw the conclusion<br />
that the triple-beam bimorph is prior to the single-beam one<br />
in the views of better temperature sensing and easier<br />
fabrication.<br />
224
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May 2011, Aix-en-Provence, France<br />
<br />
Fig. 8 SEM images of the prepared Au/Si bimorph devices before<br />
dicing. Good leading connections and isolation structure were<br />
successfully achieved.<br />
Fig. 6 SEM views of the prepared single-beam, triple-beam, and the<br />
bi-metal layer of Mo and Au.<br />
38 o C<br />
41 o C<br />
Fig. 7 Representative Z-images of the prepared triple-beam bimorph<br />
by the confocal system.<br />
Fig. 9 Photos of as-prepared chips after dicing.<br />
225
Tip Displacement, μm<br />
Tip Displacement, mm<br />
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May 2011, Aix-en-Provence, France<br />
<br />
increasing to and above about 70 o C, suggesting that the 3D<br />
Thermometer of triple-beam bimorph<br />
bimorph began to response. The practical response<br />
temperature is much higher than the required one that is<br />
22<br />
around 42 o C. One main reason is that the thickness of gold<br />
20<br />
film is only about 400 nm. Another noteworthy phenomenon<br />
is that the temperature dependence of the measured tip<br />
18<br />
displacement showed little hysteresis during the thermal<br />
cycles. Although the present prototype did not meet the<br />
16<br />
44 o C<br />
requirement of the practical application, though the<br />
14<br />
41 o C<br />
improvement of process and design, its commercial<br />
application could be expected.<br />
12<br />
10<br />
0 10 20 30 40 50<br />
Bimorph Numbering<br />
38 o C<br />
Thermometer of single-beam bimorph<br />
22<br />
20<br />
18<br />
16<br />
14<br />
12<br />
44 o C<br />
41 o C<br />
38 o C<br />
10<br />
0 1 2 3 4 5 6 7 8 9 10<br />
Bimorph Numbering<br />
Fig. 10 Measured tip displacement upon temperature increasing from<br />
25 o C.<br />
Capacitance, fF<br />
34<br />
32<br />
30<br />
28<br />
26<br />
24<br />
22<br />
20<br />
18<br />
16<br />
14<br />
12<br />
10<br />
8<br />
6<br />
4<br />
2<br />
0<br />
3A-3B pair<br />
Increase; A sample<br />
Decrease; A sample<br />
Increase; B sample<br />
Decrease; B sample<br />
20 30 40 50 60 70 80 90 100 110<br />
Temperature, o C<br />
III. CONCLUSIONS<br />
This paper presented a wireless sensor node prototype for<br />
the avian influenza monitoring system of modern poultry<br />
farm. It was found that the influenza infection could be<br />
detected by the developed sensor node at very early stage<br />
through the monitoring of body temperature and activity.<br />
Bimorph-based thermometers were also developed. The 3D<br />
bimorph shows great potential in the practical manufacture<br />
and application while more efforts are still needed for high<br />
sensitivity.<br />
REFERENCES<br />
[1] H. Pilcher, “Increasing virulence of bird flu threatens mammals”,<br />
Nature 430 (4), 4(2004)<br />
[2] H. Okada, T. Itoh, K. Suzuki, T. Tatsuya, K. Tsukamoto,<br />
“Simulation study on the wireless sensor-based monitoring system<br />
for rapid identification of avian influenza outbreaks at chicken<br />
farms”, in Proc.9 th Annual IEEE Conference on Sensors (IEEE<br />
Sensors 2010), pp 660-663, Hawaii, US, November 1-4, 2010.<br />
[3] H. Okada, T. Itoh, T. Tatsuya, K. Tsukamoto, “Wireless sensor<br />
system for detection of avian influenza outbreak farms at an early<br />
stage”, in Proc. 8 th Annual IEEE Conference on Sensors (IEEE<br />
Sensors 2009), pp 1374-1377, Christchurch, NZ, October 25-28,<br />
2009.<br />
[4] S. Scott, F. Sadeghi, D. Peroulis, “An inherently-robust 300 o C<br />
MEMS temperature sensor for wireless health monitoring of ball<br />
and rolling element bearings”, in Proc. 8 th Annual IEEE<br />
Conference on Sensors (IEEE Sensors 2009), pp 975-978,<br />
Christchurch, NZ, October 25-28, 2009.<br />
[5] A. DeHennis and K. D. Wise, “A Wireless microsystem for the<br />
remote sensing of pressure, temperature, and relative Humidity”,<br />
Journal of Microelectromechanical Systems, vol. 14, pp. 12-22,<br />
2005.<br />
[6] H. Y. Ma, Q. A. Huang, M. Qin, T. T. Lu, “A micromachined<br />
silicon capacitive temperature sensor for wide temperature range<br />
applications”, J. Micromech. Microeng. 20(2010) 055036.<br />
Fig. 11 Measured capacitance vs. temperature of the Au/Si bimorph<br />
pairs. Higher sensitivity can be expected through reducing the distance<br />
between the bimorphs.<br />
Figure 11 is measured capacitance vs. temperature of one<br />
pair of bimorph shown in the insert at top right. There is<br />
abruptly increasing of capacitance with the temperature<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
226
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<br />
Application of Wireless Sensor Nodes to<br />
Commercial Power Consumption Monitoring<br />
Toshihiro Itoh, Jun Fujimoto and Ryutaro Maeda<br />
National Institute of Advanced Industrial Science and Technology (AIST) & JST, CREST<br />
Namiki 1-2-1<br />
Tsukuba, Ibaraki 305-8564, Japan<br />
Abstract- We developed prototypes of miniaturized wireless<br />
sensor nodes for monitoring the power consumption of<br />
electrical devices and have been experimentally applying the<br />
networked monitoring system using them to power monitoring<br />
of the “convenience stores”. Nowadays, carbon dioxide (CO2)<br />
emission in the ICT field is increasing enormously due to the<br />
high electric power consumption of ICT devices, e.g., in internet<br />
data centers as well as offices and homes. In order to reduce<br />
ICT’s CO2 emission, it is indispensable to introduce energy<br />
management systems (EMS) taking the advantages of wireless<br />
sensor network technology. Therefore, we have been<br />
developing wireless clamp-meter probes integrated with a<br />
thermosensor and a monitoring system that enables<br />
simultaneous power consumption measurement of 50 power<br />
lines. Using the sensor nodes, power consumption monitoring<br />
of “convenience stores” was demonstrated and it was found<br />
that the obtained power “profile” of equipments can effectively<br />
present the key features of their-usage.<br />
social views such as technology distribution and usage of<br />
technology in society. Therefore, we propose the concept of<br />
“Meso”-level that provides a link between “Macro” and<br />
“Micro” levels, as shown in Fig. 1. On the “Meso”-level, we<br />
can address issues related to the effect of technology on<br />
social structures and human behavior with wide-ranging<br />
implications, such as education, lifestyle, business, and<br />
global economy. This paper is about an experiment related<br />
to “Meso”-level, utilizing prototype of wireless sensor nodes<br />
on a “Micro”-level. Through field experiments of sensor<br />
nodes in commercial areas, for instance, it could be<br />
considered how to integrate technology with social issue,<br />
such as power-saving in Japan.<br />
Overall goal, Social Needs<br />
e.g. CO2 70% reduction in 2050<br />
Macro<br />
I. INTRODUCTION<br />
Recently, the ICT impact on CO2 emission has attracted a<br />
great deal of attention with regard to global warming<br />
problems. This raises two issues. One is the problem of<br />
increasing power consumption by ICT diffusion. Several<br />
kinds of ICT services have been disseminated widely and<br />
rapidly throughout the society, and have changed our daily<br />
life. These services are supported by complicated hidden<br />
ICT networks, which consume much electric power.<br />
Furthermore, we have noticed that ICT diffusion indirectly<br />
accelerates economic development in developing countries,<br />
e.g. “Offshoring.” ICT diffusion may contribute to an<br />
increase in power consumption globally due to this economic<br />
development. The other issue, in contrast, is the hope to<br />
rescue climate change. The positive impact of ICT diffusion<br />
is a reduction in resource and power consumption through<br />
“dematerialization” and “efficiency improvement” in the<br />
society. Dematerialization refers to the replacement of<br />
conventional materials and human mobility, which were<br />
once needed to carry information, with electrons.<br />
By the way, current approaches to achieving a low carbon<br />
society have focused on linking “Micro” to “Macro” directly.<br />
“Macro” means overall goal of creating low-carbon society.<br />
“Micro” means technological innovations and approaches<br />
supported by institutions. However, this approach lacks<br />
Social perspective<br />
(How to integrate technology with social issues?)<br />
Technologies Development<br />
Meso<br />
Micro<br />
Technology A Technology B Technology C<br />
Fig. 1. Concept of “Meso”-level that provides a link between<br />
“Macro” and “Micro” levels.<br />
The project, “Applications of Wireless Sensor Nodes to<br />
Control Electric Power Consumption from ICT (Information<br />
Communication Technology) System” in Core Research for<br />
Evolutional Science and Technology (CREST) program<br />
Japan, started in October 2007 [1]. This project aims at<br />
reducing electricity power consumption in Japan, especially<br />
focusing on power consumption from ICT system, such as<br />
Data-centre, due to diffusion of wireless sensor nodes for<br />
monitoring electronic current of equipment. Our previous<br />
study included in this project, presented assessment results of<br />
electricity consumption from ICT in future states based on<br />
“2025 ICT Society Scenarios”[2]. These results reveal that<br />
the total power consumption from ICT was over 49.2 TWh in<br />
2008, which was around 5% of total electricity demand in<br />
Japan, 2008. Furthermore, we depicted a future ICT society<br />
based on scenario-planning and brainstorming methods, and<br />
227
estimated power consumption in 2025 utilizing these<br />
scenarios. The results suggest that power consumption from<br />
ICT reached to around 100 TWh in 2025, without<br />
considering technological progress [2]. After this study, we<br />
began to examine the new technology of wireless sensor<br />
nodes and its application of power monitoring in residential,<br />
commercial, and business areas in order to make our<br />
previous study more concrete.<br />
This paper presents the experiment of power monitoring<br />
using wireless sensor nodes in “convenience stores”. From<br />
these experiments, benefits and issues in practical use of our<br />
monitoring system were discussed.<br />
Diode Temp. Sensor<br />
MCU<br />
RF-IC<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
II. WIRELESS SENSOR NODES FOR POWER MONITORING<br />
Antenna<br />
Fig. 2. Wireless sensor node integrated with<br />
a current clamp probe (smallest type) and thermometer.<br />
TABLE I<br />
KEY COMPONENTS OF PROTOTYPE OF A SMALL WIRELESS SENSOR NODE.<br />
Fig. 2 shows a prototype of the wireless sensor node for<br />
power monitoring system. The prototyped sensor node<br />
consists of a wireless communication module, sensor<br />
interface circuits, a diode temperature sensor, and a clamp-on<br />
type current transformer having two jaws which open to<br />
allow clamping around an electrical conductor [3]. Table I<br />
shows specifications of the key components used in the node.<br />
The sensor node can detect the flow of current with the help<br />
of the current transformer and the circumambient<br />
temperature. The smallest sensor node can detect the power<br />
of 1 – 1500 W, since the small clamp-on type current<br />
transformer (CTL-6-S32-8F-CL, U.R.D., Ltd.) can be<br />
applied to the current range of 0.01 – 15 Arms. In case of<br />
monitoring the larger current, large clamp-on type current<br />
transformers, shown in Fig. 3, were utilized for high power<br />
application. The wireless communication module includes a<br />
low-voltage and low-power microcontroller unit<br />
(C8051F921, Silicon Laboratories) and single-chip 2.4 GHz<br />
transceiver (nRF24L01, Nordic Semiconductor ASA). Since<br />
the microcontroller can be operated with 0.9 V at minimum<br />
and has a built-in dc-dc converter, the module can work with<br />
one 1.5 V button-type battery. When using a battery of 100<br />
mAh, the sensor node with transmission once a second was<br />
working continuously throughout two months. If the<br />
transmission frequency is set to be once a minute, the sensor<br />
node could work throughout 10 years and be described as a<br />
“maintenance-free” node.<br />
Clamp-on Type AC<br />
Current Sensor<br />
(Smallest Type)<br />
CTL-6-S32-8F-CL [4]<br />
MCU<br />
C8051F921 [5]<br />
Transceiver IC<br />
nRF24L01 [6]<br />
Receiver<br />
100x60x17mm 3<br />
-Micro-SD storage (battery-powered receiver)<br />
-Transmitting to PC via USB<br />
- Dimensions (mm): 18W x 25H x 18t<br />
- Windng (Turn): 800<br />
- Current Range (Recommended): 10 mA<br />
– 15 A<br />
- Supply Voltage: 0.9 – 1.8 V (One-cell<br />
mode operation)<br />
- Built-in dc-dc converter with 1.8 – 3.3 V<br />
output (65 mW max)<br />
- Typical sleep mode current < 0.1 μA<br />
- 10-Bit Analog to Digital Converter<br />
- 2.4-2.5 GHz ISM band<br />
- Minimum supply voltage: 1.9 V<br />
- Supply current in TX mode @ 0dBm<br />
output power: 11.3 mA<br />
- Supply current in Power Down mode:<br />
900 nA<br />
Clamp (S):18X25X18mm:12g<br />
Wireless module,<br />
Sensor interface circuits<br />
19X14x14mm 3<br />
-Working over 12 month<br />
by SR-44 button battery<br />
Clamp (M):23X38.5X26mm:45g<br />
Clamp (L):29X44.5X31mm:70g<br />
Fig. 3. Wireless sensor nodes used for commercial power<br />
consumption monitoring experiment.<br />
III. COMMERCIAL POWER CONSUMPTION MONITORING<br />
10 “convenience stores” (CVSs) were chosen as an<br />
experiment. CVSs are like grocery stores open 24 hours and<br />
7 days a week. They have been deeply rooted in Japanese<br />
culture, there are around 42,000 stores in Japan and around<br />
35 million people use stores every day. Individual stores<br />
provide services while consuming large amount of electricity<br />
around 500 kWh/day. As shown in Table II, there are several<br />
kinds of equipment in CVSs. This equipment can be<br />
classified into two types by the supplied type of electricity,<br />
single-phase AC and three-phase AC.<br />
TABLE II<br />
TYPICAL EQUIPMENT IN CONVENIENCE STORES<br />
• Single-phase AC (200V)<br />
– Lighting<br />
– Name board<br />
– Sign stand<br />
– ATM<br />
– Copy machine<br />
– Microwave<br />
– Water heater<br />
– Heated Food (Oden,<br />
Steamed bread, Fried<br />
food)<br />
– Coffee machine<br />
• Three-phase AC (200V)<br />
– Display cooler and<br />
freezers<br />
• Reach-in, walk-in<br />
– Air-conditioning<br />
– Chilled case<br />
– Fryers<br />
– Drink cooler<br />
– Ice-cream case<br />
Fig.4 shows wireless sensor nodes attached to power lines.<br />
228
Power monitoring data detected by wireless sensor nodes<br />
was analyzed in individual stores. The result obtained in one<br />
store was summarized in Fig. 5, showing the result of power<br />
profiling of CVS. Horizontal axis presents day’s average<br />
temperature in Tachikawa area (in the center in terms of<br />
north and south and a little west of Tokyo Metropolitan area).<br />
Power consumption of display cooler and freezers and<br />
air-conditioning changed largely with temperature. More<br />
detail, power consumption of display cooler increases in a<br />
parabolas fashion with increasing temperature. That of<br />
air-conditioning shows minimum around 15 degree in spring<br />
and autumn season. These changes cause this trend of total<br />
power consumption of CVS against average temperature.<br />
On the contrary, power consumption of equipment worked<br />
by “single-phase AC” and some equipment worked by<br />
“three-phase AC” kept almost constant with temperature<br />
changes. There is a lot of equipment worked by single-phase<br />
AC. As shown in Fig.6, the equipment used on shop counter,<br />
such as heating food and coffee machine, and lighting shows<br />
large contribution to total power consumption. It can be<br />
assumed that power consumption of single-phase AC shows<br />
similar values due to day’s temperature.<br />
Single-phase AC (200V)<br />
Wireless<br />
Sensor Nodes<br />
Three-phase AC (200V)<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Power Consumption kWh/day<br />
Though the equipment of store H is older than that of store J,<br />
power consumption of store H shows similar changes with<br />
the store J. This is because store H has good installation<br />
environment of outdoor-unit. On the contrary, the oldest<br />
equipment of store F shows different changes with other<br />
stores. This may be caused by degradation of equipment and<br />
characteristic with non-invertor.<br />
Store G<br />
Microwave<br />
Lighting<br />
(shop, name board)<br />
Water heater<br />
Lighting<br />
(equipment)<br />
Heating food,<br />
Coffee-machine etc.<br />
Fig. 6. Power consumption of individual equipment in single-phase AC.<br />
.<br />
300<br />
250<br />
200<br />
150<br />
100<br />
50<br />
Display Cooler and Freezers (Walk-in Type)<br />
2000/8(6horsepower,6doors)<br />
Deteriorated equipment<br />
Different environment “outdoor-unit”<br />
2005/8(6horsepower,6doors)<br />
Energy-saving equipment<br />
2006/8(6horsepower,7doors)<br />
ATM<br />
2005/4<br />
(6horsepower,7doors)<br />
東 D文 化<br />
武 F蔵 砂 川<br />
多 G摩 関 戸<br />
砂 J 川<br />
矢 H野 口<br />
2010/3(8horsepower,7doors)<br />
Fig. 4. Example of sensor installation.<br />
0<br />
0 5 10 15 20 25 30<br />
Average Temperature ℃/ day<br />
Fig. 7. Estimate cause of power difference between stores.<br />
Power Consumption kWh/day<br />
900<br />
800<br />
700<br />
600<br />
500<br />
400<br />
300<br />
200<br />
100<br />
0<br />
Winter Spring/Autumn Summer<br />
For heating<br />
Air-conditioning<br />
Display cooler and freezers<br />
Ice-cream, Fryer, etc.<br />
Single-Phase AC(200V)<br />
1 6 11 16 21 26 31<br />
Average Temperature ℃/ day<br />
Fig. 5. Schematic power consumption in CVS.<br />
Three-phase AC (200V)<br />
Fig. 7 shows both fitted curve of individual stores and<br />
features of equipment which are installation date and<br />
performance. The latest equipment in store J shows low<br />
power consumption. As the equipment gets older in store D,<br />
F & G, curves shift direction to larger values without store H.<br />
IV. SUMMARY<br />
We have applied prototypes of wireless sensor nodes and<br />
network system to monitoring the power consumption of<br />
equipment in convenience stores as one of the important<br />
applications for commercial power monitoring. The<br />
prototype of sensor node is a wireless current clamp-on type<br />
probes integrated with a thermometer and the system enables<br />
simultaneous monitoring of 50 power lines. Using the sensor<br />
nodes and system, power consumption monitoring of 10<br />
convenience stores has been successfully demonstrated. It<br />
has been found that for this application, besides the low cost<br />
of sensor systems, ease of installation and undisturbed<br />
environment in setting monitoring system were strongly<br />
required. For achieving these properties and the low cost of<br />
sensor systems, it would be necessary to have wireless and<br />
non-battery system, small size sensor nodes.<br />
ACKNOWLEDGMENT<br />
The authors thank Seven-Eleven Japan Co., Ltd. for the<br />
assistance of the power monitoring experiment in the stores.<br />
229
REFERENCES<br />
[1] http://www.jst.go.jp/kisoken/crest/en/area04/5-02.html<br />
[2] J. Fujimoto and T. Hata, Assessment of Power Consumption from<br />
ICT in Future States based on “2025 ICT Society Scenarios”,<br />
EcoDesign2009, Dec. 9, 2009, pp.837-842.<br />
[3] T. Itoh, Y. Zhang, M. Matsumoto and R. Maeda, Wireless Sensor<br />
Nodes for Monitoring the Power Consumption of Information and<br />
Communication Devices, EcoDesign2009, Dec. 9, 2009,pp853-856.<br />
[4] http://www.u-rd.com/english/products/ac/ac_3.html.<br />
[5] https://www.silabs.com/products/mcu/lowvoltagelowpower/<br />
Pages/default.aspx.<br />
[6] http://www.nordicsemi.com/files/Prod_brief_RFSilicon_<br />
nRF24L01.pdf<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
230
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Developing MEMS DC Electric Current Sensor<br />
for End-use Monitoring of DC Power Supply<br />
Kohei Isagawa 1 , Dong F. Wang 1 , Takeshi Kobayashi 2 , Toshihiro Itoh 2 and Ryutaro Maeda 2<br />
1<br />
Micro Engineering & Micro Systems Laboratory, Ibaraki University (College of Eng.), Hitachi, Ibaraki 316-8511 Japan<br />
(Tel: +81-294-38-5024; Fax: +81-294-38-5047; E-mail: dfwang@mx.ibaraki.ac.jp)<br />
2 Ubiquitous MEMS and Micro Engineering Research Center (UMEMSME), AIST, Tsukuba, Ibaraki 305-8564, Japan<br />
Abstract- A non-drive, non-contact MEMS DC electric current<br />
sensor to satisfy the increasing needs of DC power supply for<br />
monitoring the electricity consumption by either one-wire or<br />
two-wire appliance cord has been proposed. A micro magnet is<br />
integrated into the MEMS-scale DC sensor and the appropriate<br />
position for locating the micro magnet has been theoretically<br />
pinpointed. A macro-scale prototype DC sensor was therefore<br />
fabricated, and an impulse piezoelectric voltage output can be<br />
clearly detected out when a DC electric current was applied to a<br />
two-wire electrical appliance cord. A linear relationship<br />
between the detected peak value of the impulse output voltage<br />
and the applied DC electric current was further obtained based<br />
on the empirical measurements. After the demonstration of the<br />
macro-scale prototype DC sensor, the MEMS-scale DC sensor<br />
has been then theoretically designed from the view point of<br />
reasonable output voltage measurements, and preliminarily<br />
micro-fabricated for further characterizations.<br />
Keywords- MEMS DC sensor; Electricity end-use monitoring; DC<br />
power supply; PZT; Non-drive; Non-contact; Two-wire cord<br />
I. INTRODUCTION<br />
The energy consumption by factories, automobiles and<br />
even human’s daily life make the increase of CO 2 exhaust,<br />
which subsequently aggravates the green-house effect. The<br />
total amount of CO 2 exhaust in Japan in 2008 was 1.21 billion<br />
ton, and about one of thirds was caused by residential section<br />
and commercial section. About 40% of the amount of CO 2 is<br />
caused by electrical consumption of household equipment and<br />
Information and Communication Technology (ICT) devices.<br />
Moreover, in Japan, the electricity consumption of ICT devices<br />
will increase by about 4.2 times by the year of 2025. The<br />
electricity consumption of internet data center (IDC) is also<br />
rapidly increasing with the increasing amount of the data traffic<br />
on the internet. It is estimated to grow by two order of its<br />
present value by the year of 2025 [1]. In addition, IDC have<br />
been anticipated to achieve a decrease in AC to DC conversion<br />
loss. Something similar is being conducted at DC houses<br />
consisting of a solar battery or a storage cell. Therefore, it is<br />
essential to monitoring the DC electricity consumption so as to<br />
establish an effective electricity management system.<br />
Although Hall element based direct current sensor is the<br />
main stream at the present day. However, since the household<br />
equipment and ICT device use two-wire appliance cord, the<br />
Hall element based direct current sensors can not be applied<br />
directly without a line separator to first separate the two-wire<br />
appliance so as to measure the current. In addition, a drive<br />
voltage is physically necessary for the Hall element based<br />
sensors, which is inconvenient to monitor electrical<br />
consumption at anytime and anywhere without a power supply.<br />
In this work, a novel MEMS DC sensor, which is<br />
self-driven and applicable to both one-wire and two-wire<br />
appliance cord, has been proposed, designed, and preliminarily<br />
investigated. The proposed MEMS DC sensor is believed to<br />
be very useful to various kinds of DC systems in the near<br />
future.<br />
II.<br />
PORPOSING MEMS-SCALE DC SENSOR AND ITS<br />
APPLICATION TO TWO-WIRE ELECTRICAL APPLICANCE CORD<br />
The proposed MEMS DC sensor, as shown in Fig. 1, is<br />
expected to be utilized for monitoring the electricity<br />
consumption by one-wire or two-wire appliance cord. The<br />
critical component, which is encapsulated in the green shell, is<br />
a cantilever made up of piezoelectric film and a permanent<br />
micro-magnet fixed on the end. When a direct current from DC<br />
power supply is flowed via a two-wire appliance cord, the<br />
piezoelectric film coated cantilever is bended by the created<br />
magnetic force acted on the micro magnet, and the output<br />
voltage is generated by the piezoelectric film and the applied<br />
DC current is therefore detected out.<br />
Fig. 1. A newly proposed non-drive and non-contact MEMS DC sensor<br />
for end-use monitoring of DC power supply.<br />
231
Compared with commercially available current sensors, the<br />
proposed MEMS DC electric current sensor has several<br />
advantages such as non-drive, non-contact, and capability in<br />
sensing one-wire to two-wire appliance cord. In this study, a<br />
macro-scale prototype DC sensor was also fabricated and<br />
demonstrated to confirm whether the impulse (momentary)<br />
output voltage signal from DC current could be clearly detected<br />
out or not.<br />
III.<br />
MAGNET’S POSITION WITH A RESPECT TO THE<br />
GENERATED MAGNETIC FILED<br />
The magnetic force on a permanent magnet in a magnetic<br />
field is proportional to the integral of the field gradient over the<br />
magnet’s volume [2 - 4]. If the magnet comes close to a long<br />
current-carrying cord, the magnetic forces in the plane normal<br />
to the wire can be described by Equation (1).<br />
( )<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
pervious study [5].<br />
d<br />
d<br />
= BF ry ∫<br />
d = BF VH<br />
( )<br />
dy<br />
∫ d (1)<br />
rz VH<br />
dz<br />
dH z<br />
dz<br />
⎛<br />
⎜<br />
π ⎝<br />
( − )<br />
a + ⎞<br />
+<br />
yi<br />
ay )(<br />
⎟ (4)<br />
+−<br />
zay<br />
))( ++<br />
zay<br />
⎠ (<br />
−= 222 ))(<br />
222<br />
Fig. 3 plots the z-direction gradient of the y-component of<br />
the magnetic field surrounding a two-wire appliance cord.<br />
The dash lines in Figs. 2 and 3 mean the maximum value of<br />
the z-direction gradient of the z-component of the magnetic<br />
field for a single wire or a two-wire appliance cord,<br />
respectively.<br />
In the above Equation, y and z are horizontal and vertical<br />
direction, respectively, F is the magnetic force on the magnet,<br />
H y and H z are horizontal and vertical components of magnetic<br />
field in amperes pre meter, B r is the remanence of permanent<br />
magnet in Tesla, and V is the magnet’s volume. Assuming that<br />
the remanence of the permanent magnet is uniform and aligned<br />
in the positive z-direction.<br />
In order to analyze the magnetic field gradient all around an<br />
electric power cord, the magnetic field surrounding a single<br />
current-carrying cord should be considered and is described by<br />
Equation (2).<br />
Fig. 2. Plot of the magnitude of the vertical component of<br />
z-direction magnetic field gradient around a single wire, 10 A current assumed.<br />
i<br />
H = (2)<br />
2πr<br />
H is the magnetic field in amperes per meter, i is the current<br />
in wire (A), and r is radial distance from the wire to point of<br />
interest. The direction of H is determined using the ‘right-hand<br />
rule’, aligning the thumb of the right hand with direction of<br />
flowing current. Equation (3) can thus be inducted from<br />
Equation (2) for z-direction gradient of the z-component of the<br />
magnetic field surrounding a single wire.<br />
dH z iyz<br />
−= (3)<br />
dz π + zy<br />
222<br />
)(<br />
Fig. 2 plots the y-direction gradient of the y-component of<br />
the magnetic field surrounding a single wire.<br />
In case of a two wire appliance cord, we define a as the<br />
distance between the center of the appliance cord and the center<br />
of right cord or lift one. Then, we induct the z-direction gradient<br />
of the y-component of the magnetic field surrounding a<br />
two-wire appliance cord by Equation (4) which reported in our<br />
Fig. 3. Plot of the magnitude of the vertical component of<br />
z-direction magnetic field gradient around a two-wire appliance cord,<br />
10 A current assumed.<br />
IV.<br />
DEMONSTRATION MEASUREMENTS BY USING A<br />
MACRO-SCALE DC SENSOR DEVICE<br />
Generally speaking, when a direct current is applied to the<br />
appliance cord, the output signal from the proposed cantilever<br />
based MEMS DC sensor is supposed to be very impulsive.<br />
Therefore a macro-scale device, as shown in Fig. 4, has been<br />
fabricated to demonstrate whether the output signal arising<br />
from direct current can be measured or not. The macro-scale<br />
232
device consisted of a bimorph cantilever with a permanent<br />
magnet located at the cantilever tip. The size of the bimorph<br />
and the permanent magnet is 28 × 13.4 × 9 (mm 3 ) and 5 × 5<br />
× 2 (m 3 ), respectively. The remanence magnetization B r is<br />
assumed as 1.2 T. Fig. 5 shows a simple measurement set-up<br />
for preliminary demonstration by the macro-scale device. The<br />
output voltage, arising from the direct current supplied by a DC<br />
power supply (AND Co. AD-8735A, JPN), is measured by an<br />
oscilloscope (Tektronix Co. TDS2014, USA). The applied<br />
direct current is measured by a current probe (Tektronix Co.<br />
TCP312, USA). The output voltage and the applied direct<br />
current were measured with a 2msec sampling interval.<br />
However, the center of the permanent magnet is located at 25.5<br />
mm from the base of cantilever and 4.1 mm from the center of<br />
the appliance cord.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
impulse signal and the applied DC current. The sensitivities of<br />
the macro-scale device were derived as around 7 mV/A and 5<br />
mV/A for the cases of turning on and turning off the DC power<br />
supply, respectively.<br />
Fig. 6. A typical measurement showing the output voltage impulses when a<br />
direct current of 2A was applied to a two-wire appliance cord (ON), and when<br />
the applied current was discontinued (OFF), respectively, as shown in Fig. 5.<br />
The output voltage from the current probe was also drawn for comparison.<br />
Fig. 4. A macro-scale prototype DC current sensor with a permanent magnet<br />
on the tip of the piezoelectric bimorph cantilever was fabricated<br />
for demonstration measurements.<br />
Fig. 7. The peak values (in error bar) of the output voltage impulse as a function<br />
of the applied direct current, as measured in Fig. 6.<br />
V. DESIGN OF MEMS-SCALE DC SENSOR DEVICES<br />
Fig. 5. Measurement set-up for demonstration measurements by the fabricated<br />
macro-scale device shown in Fig. 4.<br />
As a result, we succeeded in measuring the impulsive values<br />
of the output voltage by the macro-scale device for the first time.<br />
Fig. 6 typically shows that the output voltage impulse signal was<br />
clearly detected out when 2A was applied to a two-wire<br />
appliance cord. The peak value was measured as -10 mV when<br />
turning on the DC power supply, while that was measured as 17<br />
mV when turning off the DC power supply. The output voltage<br />
impulse signal converged to zero within 0.07 sec. Such<br />
measurement can be conducted from a lower applied current of<br />
0.5 A to a higher one of 3 A. Fig. 7 shows a linear relation<br />
between the absolute peak values (in error bar) of output voltage<br />
A. StructuralDesign with An Applicable Approach<br />
We assume that the measurement system for the<br />
MEMS-scale DC sensor employ a microcomputer built-in A/D<br />
converter with the resolution capability of 12 bit and the<br />
reference voltage of 3 V. In the case of future DC houses, it is<br />
also reasonable to further assume that a direct current supplied<br />
to the home electrical appliance is in the range of 0.04 A to 10<br />
A. It is therefore very crucial to carry on such kind of a<br />
structural design to make the MEMS-scale DC sensor not only<br />
measure a detectable impulse signal of over 0.74 mV even<br />
when a very lower direct current of 0.04A is applied, but also be<br />
able to bear off the stronger bending when a higher direct<br />
current of 10 A is applied. Generally, the piezoelectric sensors<br />
work with a charge amplifier driven by an electrical power. It is<br />
thus necessary to design a novel sensor device which can<br />
generate enough high voltage by itself so as to meet our future<br />
powerless working requirement.<br />
Fig.8 gives a schematic showing such kind of design<br />
233
approach. In order to increase the output voltage, the logic is to<br />
fabricate more PZT plates on the substrate and connect them in<br />
series with each other. As shown in Fig.8, l and L m are the<br />
length of the PZT plate and space for magnet, respectively. w<br />
and φ are the width of the cantilever and the space between<br />
neighboring PZT plates. However, a novel device design aimed<br />
to accomplish a high output voltage will be theoretically and<br />
quantitatively discussed in detail in the following Session.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
where F z is the magnetic force when counterpoises restoring<br />
force of cantilever. E i and E p are Young’s moduli of each layer<br />
and PZT thin film, respectively. A i is the z-y cross-section area<br />
of the each layer and is expressed as follows.<br />
i<br />
= φ−− }) hnwA<br />
1({<br />
(9)<br />
where h i is the thickness of each layer, while φ = 0 in case of z-y<br />
cross-section of the substrate. In Equation (8), Z p is the distance<br />
between the position z p of center of the PZT thin films and the<br />
position z N of neutral axis parallel to the length direction of<br />
cantilever and can be expressed as Equation (10). Z i is the<br />
distance between the position z i of center of the each layer and<br />
the neutral axis parallel to the length direction of cantilever and<br />
expressed as Equation (11),<br />
i<br />
−= zzZ (10)<br />
Npp<br />
−= zzZ (11)<br />
Nii<br />
Fig. 8. A schematic figure showing the PZT plates and definitions of various<br />
parameters used in the following theoretical derivation.<br />
B. Theoretical and Quantitative Derivation of the output<br />
voltage<br />
We define z axis, y axis, x axis and the origin of the z axis as<br />
the axis vertical to the surface (thickness direction), the width<br />
direction, the length direction and the bottom of the cantilever<br />
in Fig. 8, respectively. Also, the width of individual PZT plate<br />
w E is expressed by the following Equation (5).<br />
where the neutral axis z N is expressed as follows [7],<br />
∑i<br />
∑<br />
AEz<br />
iii<br />
z<br />
N<br />
=<br />
(12)<br />
AE<br />
i<br />
ii<br />
The accumulated charge of individual PZT plate can then be<br />
expressed as follows by Equations (6), (7), (8).<br />
− nw<br />
w − φ) 1(<br />
E<br />
= (5)<br />
n<br />
The accumulated charge in individual PZT plate is<br />
expressed as follows from Gauss’s law:<br />
Q<br />
ind<br />
∫∫<br />
= Ddwdl<br />
(6)<br />
l w<br />
where D is the electrical-field displacement in z-direction. For<br />
piezoelectric sensing, the electrical-field displacement D<br />
without applying any external electrical field is described by<br />
Equation (7):<br />
= dD<br />
31σ p<br />
(7)<br />
where d 31 and σ are transverse piezoelectric constant<br />
(-50pmV -1 ) and x-direction stress when applied bending<br />
moment on the tip of cantilever, respectively. Also, x-direction<br />
stress σ is expressed as follows [6],<br />
i<br />
E<br />
1<br />
31<br />
1<br />
Vtotal<br />
= 2<br />
(<br />
mpp<br />
+ ) FlLZE<br />
z<br />
∑ +<br />
i<br />
σ = 2<br />
(8)<br />
2<br />
∑ ( + ZAIE<br />
iiii<br />
)<br />
Q<br />
ind<br />
=<br />
31<br />
1<br />
∑<br />
2<br />
+<br />
i<br />
+ )(<br />
l wl<br />
F (13)<br />
ZAIE<br />
)(<br />
mpp<br />
E<br />
2 z<br />
iiii<br />
When n pieces of PZT plates are electrically connected in<br />
series, the total accumulated charge Q total is equivalent to the<br />
accumulated charge Q ind in individual PZT plate.<br />
ind<br />
= QQ (14)<br />
total<br />
Therefore, the wider the width of the PZT plate, the more<br />
increment the charge of the device.<br />
The output voltage can be expressed by Equation (14),<br />
+ )(<br />
l wlL<br />
1<br />
F (15)<br />
ZAIE<br />
)(<br />
CC<br />
mpp<br />
E<br />
2<br />
z<br />
iiii<br />
total<br />
+<br />
others<br />
234
where C others is the capacitance of the measurement system.<br />
C total is the total capacitance of the device and is expressed by<br />
Equation (16).<br />
Cind<br />
Ctotal = (16)<br />
n<br />
The capacitance of individual PZT plate C ind is given by<br />
S<br />
E<br />
Cind<br />
= εε<br />
(17)<br />
0<br />
hp<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Fig. 12 shows a typical design of our MEMS DC sensor<br />
consisting of ten-PZT plates. Electrical connects in series by<br />
wire bonding for signal amplification will be operated. The<br />
square space Pt/Ti is the space for positioning the permanent<br />
magnet. The total output voltage V total in Fig. 12 generates 3.3<br />
mV when the appliance cord is supplied by 0.04 A. On the<br />
other hand, the tensile stress acted in the substrate is calculated<br />
as around a critical value of 9 MPa if applied a current of 10 A.<br />
This fact implies that substrate materials should have a<br />
sufficient tensile yield stress higher than the above mentioned<br />
critical value.<br />
where ε and ε 0 are the vacuum dielectric constant (8.85×<br />
10 -12 Fm -1 ) and relative dielectric constant (1000) of PZT thin<br />
films, respectively. S E is the area of the electrode and h p is the<br />
thickness of the PZT thin films. Therefore, in order to obtain a<br />
high V total in Equation (15), we need to decrease the capacitance<br />
C ind and increase PZT plates at the same time.<br />
We thus turn to calculate Q total , C total , V total to design the<br />
MEMS DC sensor by Equations (5)-(17). The size of cantilever<br />
is chosen as 4300 μm×4000 μm to withstand the magnetic<br />
force when a current of 10A is applied. The space between two<br />
PZT plates φ is 190 μm. The magnetic force is calculated under<br />
the current of 1A and the center of magnet is located at 3.8 mm<br />
from the center of the appliance cord. In this calculation, we<br />
ignore C others in Equation (15). Table 1 and Table 2 show values<br />
of h i , E i , z i for each layer of the MEMS DC sensor and values of<br />
w E , z N , Z p with respect to number of PZT plates, respectively.<br />
Fig. 9, Fig. 10, and Fig. 11 show the results of Q total , C total , and<br />
V total , respectively. It can be noted from Fig. 9 and Fig. 10, the<br />
more the PZT plates, the more decrease the values of Q total and<br />
C total . However, since the decrease of C total is greater than that of<br />
Q total , an increasing behavior can be observed for V total from Fig.<br />
11. Therefore, we consider it would be effective to<br />
simultaneously increase number of PZT plates and decrease w E ,<br />
φ by micromachining to achieve a high output voltage DC<br />
sensor.<br />
Fig. 9. Calculated Q total as a function of number of PZT plates.<br />
Table 1. Values of h i , E i , z i for each layer of the MEMS DC sensor.<br />
Fig. 10. Calculated C total as a function of number of PZT plates.<br />
i material h i (μm) E i (μm) z i (μm)<br />
top 5 top Pt/Ti 0.2 168 6.6<br />
4 PZT 2 72.5 5.5<br />
3 bottom Pt/Ti 0.2 168 4.4<br />
2 thermal SiO 2 0.3 73.1 4.15<br />
bottom 1 structure Si 4 190 2<br />
Table 2. Values of w E , z N , Z p with respect to number of PZT plates.<br />
Number of PZT plates w E (μm) z n (μm) Z p (μm)<br />
10 220 2.54 2.96<br />
8 323 2.60 2.90<br />
6 473 2.63 2.87<br />
4 835 2.71 2.79<br />
2 1860 2.76 2.74<br />
1 3910 2.78 2.72<br />
Fig. 11. Calculated C total as a function of number of PZT plates<br />
235
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Fig.12. A typical design of MEMS DC sensor consisted of ten beam-type PZT<br />
plate array for further verification.<br />
VI.<br />
MICROFABRICATION OF PROTOTYPE MEMS-SCALE DC<br />
SENSOR DEVICES<br />
As shown in Fig. 13, the prototype devices are fabricated<br />
from multilayer of Pt/Ti/PZT/Pt/Ti/SiO 2 deposited on<br />
silicon-on-insulator (SOI) wafers using a 5-mask<br />
micromachining process [8][9]. Deposition of the multilayer is<br />
started from thermal oxidation of SOI wafers followed by Pt/Ti<br />
bottom electrodes deposition. After (100)-oriented PZT thin<br />
film is deposited by a sol-gel process [10], Pt/Ti top electrodes<br />
are then sputtered.<br />
The etching process is as follows. Pt/Ti top electrodes are<br />
first etched by an Ar ion through mask No.1. PZT thin films are<br />
then wet-etched with an aqueous solution of HF, HNO 3 , and<br />
HCl through mask No.2. After Pt/Ti bottom electrodes are<br />
etched by an Ar ion through mask No.3, both the thermal SiO 2 ,<br />
structural Si and buried oxide (BOX) are etched by RIE with<br />
CHF 3 gas (SiO 2 ) and SF 6 gas (Si) through (mask No.4). Finally,<br />
the substrate Si is etched from the backside to release the<br />
cantilever through mask No.5.<br />
Fig. 14 shows the fabricated MEMS DC sensor device. It<br />
can be noted that the cantilever are warped due to the BOX<br />
layer remained on the backside.<br />
Fig.13. Schematic diagram of the fabrication process chart. (a)<br />
Pt/Ti/PZT/Pt/Ti/SiO2 deposition, (b) Pt/Ti/PZT/Pt/Ti/SiO2 etching, (c)<br />
cantilever patterning, (d) substrate etching from backside to release cantilever.<br />
Fig. 14. The optical micrograph (Top view) of the fabricated MEMS DC sensor<br />
device.<br />
VII. CONCLUSIONS<br />
A non-drive, non-contact prototype MEMS-scale DC<br />
electric current sensor has been theoretically studied,<br />
geometrically designed and preliminarily fabricated for a<br />
measurement range of 0.04 A to 10 A. It would be effective to<br />
simultaneously increase number of PZT plates and narrow both<br />
electrode width and space between neighboring PZT plates by<br />
micromachining to achieve a high output voltage DC sensor.<br />
Based on a handmade macro-scale prototype DC sensor, we<br />
also succeeded in detecting out the impulsive values of the<br />
output voltage for a current range of 0.5 A to 3 A for the first<br />
time.<br />
ACKNOWLEDGEMENT<br />
Part of this work was supported by MEMS Inter University<br />
Network and performed in the Ubiquitous MEMS & Micro<br />
Engineering Research Center (UMEMSME) of National<br />
Institute of Advanced Industrial Science & Technology<br />
(AIST).<br />
REFERENCES<br />
[1]Y. Zhang, S.Uchiyama, D.K.Lee, H.Hiroshima, T.Itoh, R.Maeda<br />
International Workshop on Green Device and Micro Systems GDMS<br />
2011(10pp)<br />
[2]E S Leland, P K Wrigth and R M White J.Microelectromech. Microeng. 19<br />
(2009) 094018 (6pp)<br />
[3]M.V.Shutov, E.E. Sandoz, D.L.Howard, T.C. Hsia, R.L. Smith, S.D. Collins<br />
Sensors and Actuators A 121 (2005) 566-575<br />
[4]H.H Yang, N.V.Myung, J.Yee, D.-Y. Park, B.-Y. Yoo, M. Schwartz, K.<br />
Nobe, J.W.Judy Sensors and Actuators A 97-98 (2002) 88-97<br />
[5]Kohei Isagawa, Dong F.Wang, Takeshi Kobayashi, Toshihiro Itoh, Ryutaro<br />
Maeda JCK MEMS/NEMS 2010 (P-05)<br />
[6]Qiang Zou, Wei Tan, Eun Sok kim and Gerald E.Loed J.Microelectromech<br />
Syst., vol.17, no.1, pp45-57,Feb.2008<br />
[7]T Kobayashi, H Okada, T Masuda and R Meada, T Itoh Smart mater.<br />
Struct.19(2010) 105030 (8pp)<br />
[8]M.S.Weinbeng, J.Microelectromech. Syst., vol.8, no.4,<br />
pp529-533,Dec.1999.<br />
[9]Kobayashi T, Ichiki M, Kondou R, Nakamura K and Maeda R 2007 j.<br />
Micromech. Microeng. 17 1238<br />
[10]Kobayashi T, Ichiki M, Tsaur J and Maeda R 2005 Thin Solid Films 489 74<br />
236
11-13 <br />
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<br />
Low Power Analog to Digital Convertor with Digital<br />
Calibration for Sensor Network<br />
Tsukasa Fujimori, Hiroshi Imamoto, Hideaki Kurata, Yasushi Goto, Toshihiro Ito, Ryutaro Maeda<br />
BEANS Laboratory G device Center<br />
G Device Research Body Ubiquitous MEMS and Micro Engineering R.C.<br />
1-2-1, Namiki, Tsukuba, Ibaraki 305-8564, JAPAN<br />
Abstract- A low-power analog-front-end (AFE) LSI for<br />
sensor networks—based on an analog-to-digital convertor<br />
(ADC) with digital calibration—was developed. Power<br />
consumption of the ADC in the AFE LSI was reduced by<br />
applying digital calibration. As a result, the proposed<br />
successive approximation register (SAR) ADC achieves both<br />
high effective resolution (11.7 bits) and extremely low power<br />
consumption (2.5 mW) at 1 Msps. Moreover, average power<br />
consumption of the AFE LSI (including the ADC) is about 5<br />
μW, which is low enough for sensor networks.<br />
I. INTRODUCTION<br />
Energy-saving technologies for CO2 emission reduction<br />
have become ever more important in recent years. Sensor<br />
networks have been expected to provide one of the<br />
most-promising solutions for energy saving. However, the<br />
large physical size and high power consumption of sensor<br />
nodes have been major constraints on the widespread usage of<br />
sensor networks.<br />
Sensor nodes generally consist of batteries, wireless circuits,<br />
a microcomputer, sensors, and analog-front-end (AFE) circuits.<br />
The wireless circuits and the microcomputer utilize<br />
deep-submicron CMOS technology to reduce power<br />
consumption. On the other hand, the power consumption of<br />
the AFE circuits has not been reduced. It is not easy to apply<br />
deep-submicron CMOS technology for the AFE circuits,<br />
which include amplifiers and analog-to-digital converters<br />
(ADCs), because the analog circuits that compose the AFE<br />
circuits are sensitive to process variation. It is therefore<br />
essential to develop low-power AFE circuits for sensor<br />
networks.<br />
Responding to the above-mentioned need, in the present<br />
study, we developed a low-power AFE LSI for a sensor<br />
network by utilizing an ADC with digital-calibration, which<br />
are generally used for developing high speed and high<br />
accuracy ADC circuits. Power consumption of the ADC in the<br />
AFE circuits was reduced by applying digital-calibration<br />
techniques. As a result, an ADC with high effective resolution<br />
of 11.7 bits and extremely low power consumption of 2.5 mW<br />
at 1 Msps (mega samples per second) was developed.<br />
Moreover, average power consumption of the AFE LSI<br />
(including the ADC) is about 5 μW, which is low enough for<br />
sensor networks.<br />
II. TARGET OF AFE LSI FOR SENSOR NETWORK<br />
Figure 1 shows a block diagram of the AFE LSI, which<br />
consists of an ADC, a programmable gain amplifier (PGA), a<br />
voltage-reference (Vref) generator, a clock generator, and<br />
digital interface circuits. The ADC is a successive<br />
approximation register (SAR) type [2]. The AFE LSI<br />
amplifies analog signals from the sensors and converts them<br />
into digital signals. Note that most of the required periphery<br />
circuits for these circuits are also built-in.<br />
Table 1 lists the performance targets for the developed AFE<br />
LSI. Resolution of the ADC of 14 bits, effective resolution of<br />
ADC of 11 bits or more, and sampling rate of the ADC of 1<br />
Msps all exceed the performance of the ADC generally used<br />
by sensor nodes, and the target power consumption is below<br />
10 mW (which is the approximate power consumption at the<br />
time of operation). Furthermore, the target average power<br />
consumption when performing one sampling per second is 10<br />
μW or less. These power consumptions are one half or less<br />
than those of present AFE circuits.<br />
Sensor A<br />
Sensor B<br />
Sensor C<br />
Sensor D<br />
M<br />
UX<br />
M<br />
UX<br />
Digital interface circuits MCU<br />
Figure 1.<br />
PGA<br />
Vref generator<br />
SAR ADC with<br />
digital calibration<br />
Clock generator<br />
Block diagram of AFE LSI<br />
Table 1. Targets of the AFE LSI for sensor network<br />
Targets<br />
Resolution<br />
14 bits<br />
(effective resolution >11 bits)<br />
Sample rate<br />
1 Msps<br />
Power (active)<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
If these power consumptions of the AFE LSI were achieved<br />
<br />
6<br />
in practice in the future, the frequency of replacing batteries of<br />
sensor nodes could be decreased; moreover, sensor nodes with<br />
5<br />
a low-power AFE LSI could be operated by energy-harvesting<br />
devices [3]. However, if the area of the chip increases, its cost<br />
4<br />
will also increase, even though the chip may be highly<br />
efficient, and it would not be practical. The target for the<br />
3<br />
effective area was therefore set at less than 2.25 mm 2 .<br />
III. DESIGN OF LOW-POWER SAR ADC WITH DIGITAL<br />
CALIBRATION<br />
We investigated reducing supply voltage as an approach to<br />
lower the power consumption without losing performance.<br />
This is because the power consumption of CMOS circuits is<br />
generally proportional to the square of the voltage.<br />
According to the International Technology Roadmap for<br />
Semiconductors (ITRS), supply voltages and leakage<br />
currents of the CMOS circuits are shown in figure 2 [4].<br />
The horizontal axis is the gate length of the CMOS<br />
transistors, the left vertical axis is supply voltage, and the<br />
right vertical axis is leakage current. Low supply voltage has<br />
become standard with the development of CMOS processes.<br />
For example, 1.2 V is used for gate lengths less than 130 nm.<br />
However, the smaller the gate length of the transistors<br />
becomes, the more the leakage current increases.<br />
In this paper, we chose the 130-nm process for AFE LSI,<br />
because it provides the lowest leakage current in supply<br />
voltage of 1.2V.<br />
At present, the power-supply voltage of sensor nodes is in<br />
the range of 2 to 5 V. If sensor nodes could be operated at 1.2<br />
V, power consumption would be 30 to 90% lower than the<br />
present level. However, it is not easy to reduce the supply<br />
voltage of the AFE LSI. It is especially difficult to reduce the<br />
supply voltage of the ADC circuits.<br />
Figure 3(a) shows a block diagram of conventional SAR<br />
ADC [2]. The ADC consists of an analog-circuits part and a<br />
digital-circuits part. The voltage of incoming signals is<br />
compared with the voltage generated by the digital-to-analog<br />
circuit (DAC), and the result of the comparison is set to the<br />
successive approximation register (SAR). The voltage<br />
generated by the DAC is changed in order and the<br />
comparison is repeated, so data of analog-to-digital<br />
conversion is obtained.<br />
The supply voltage of the digital circuits can be reduced<br />
easily. However, if the supply voltage decreases, the<br />
operational margin of the analog circuits also decreases; the<br />
accuracy of the DAC is degraded. As a result, the resolution<br />
of SAR ADC degrades at low supply voltage. As a<br />
countermeasure to this difficulty, huge capacitors are usually<br />
used for standard SAR ADC to suppress the capacitor<br />
mismatch derived from process variation. This leads to large<br />
area of SAR ADC for 1.2-V operation. To solve this problem,<br />
we applied digital calibration technique to SAR ADC for<br />
sensor networks.<br />
Figure 3(b) shows the SAR ADC with digital calibration,<br />
which is performed by using digital circuits and software<br />
algorithms to cancel errors that occur in analog operation [1].<br />
Supply voltage (V)<br />
2<br />
1<br />
0<br />
1000 500 350 230 180 130 90 65 45<br />
Gate length (nm)<br />
Figure 2. supply voltages and leakage currents of<br />
the CMOS circuit<br />
(a) without digital calibration technique<br />
(b) with digital calibration technique<br />
Figure 3. Block diagram of SAR ADC<br />
Figure 4. Chip layout of SAR ADC<br />
120<br />
100<br />
80<br />
60<br />
40<br />
20<br />
0<br />
Comparator<br />
Leakage current (pA)<br />
Digital<br />
circuits<br />
238
11-13 <br />
May 2011, Aix-en-Provence, France<br />
One complete conversion process consists of two conversion<br />
<br />
phases, in which a single SAR quantizer performs the same<br />
digitization twice, perturbed by two added offsets, +Δ and -Δ,<br />
and resolves to two non-binary codes, D+ and D-,<br />
SAR ADC<br />
respectively. These codes are subsequently converted to<br />
1.6 m m 2<br />
binary ones, shown as d+ and d-, by a weighted sum of the<br />
individual bits. If the conversion process is ideal, the<br />
DAC<br />
difference e between d+ and d- less 2Δ must be zero. In other<br />
words, a nonzero e will provide information to infer the<br />
PG A LOGIC<br />
unknown weighting vector W and Δ; and this idea leads to<br />
the adaptive learning algorithm to calibrate the SAR ADC in<br />
which e is gradually forced to zero. After the learning<br />
procedure converges, the mean of d+ and d- will yield the<br />
Vrefgenerator<br />
correct digital output with the effect of Δ cancelled.<br />
Since errors of the DAC are canceled, digital calibration<br />
Clock generator<br />
technique decreases capacitance required for DAC. As a<br />
result, both small area and low power consumption of SAR<br />
Figure 5. Chip photo graph of our AFE LSI<br />
ADC is achieved.<br />
Figure 4 shows the layout of the SAR ADC with digital<br />
calibration. The ADC can be operated at 1.2 V and is<br />
calibrated to sufficient accuracy. Nevertheless, the size of the<br />
ADC is very small, namely, about 1.6 mm 2 , which is reduced<br />
SNDR = 62.4 dB (10.1 bit)<br />
by about 90% of the standard size.<br />
To sum up, a low-power-consumption ADC with very high<br />
performance was achieved by the digital calibration.<br />
IV. AFE LSI FOR SENSOR NETWORK<br />
Figure 5 shows a chip photograph of the AFE LSI, which<br />
was fabricated by using the 0.13-μm CMOS process. The chip<br />
size is 4 by 4 mm, and the effective chip size (including PGA,<br />
DAC, SAR ADC, clock generator, reference voltage generator,<br />
and logic circuits) is about 2.3 mm 2 . The AFE LSI's clock<br />
frequency is 20 MHz.<br />
Figure 6 shows measured spectra of the ADC of the AFE<br />
LSI signals, namely, a sine wave of 1 kHz and 90% of<br />
full-scale voltage. Figure 6(a) is the spectrum obtained without<br />
digital calibration, and Figure 6(b) is the spectrum obtained<br />
with digital calibration. From these spectra, signal-to-noise<br />
and distortion ratios (SNDRs) were estimated as 62.4 dB (10.1<br />
bits) in the case without digital calibration and 72.0 dB (11.7<br />
bits) in the case with digital calibration. That is, the accuracy<br />
of the ADC with digital calibration is increased about three<br />
times compared to that of the ADC without digital calibration.<br />
Evaluation results of effective size and power consumption<br />
for each circuit block are listed in Table 2. The power<br />
consumption of the ADC is about 2.5 mW, and the total power<br />
consumption of the AFE LSI is about 6.4 mW. Moreover,<br />
average power consumption of the AFE LSI at sampling rate<br />
of one sample per second was estimated to be about 5 μW,<br />
which is small enough compared to that of AFE circuits of<br />
general sensor nodes. In other words, the total of power<br />
consumption of the AFE LSI is considerably decreased while<br />
small chip size was maintained. The developed AFE LSI is<br />
thus considered suitable for sensor nodes.<br />
(a) without digital calibration<br />
SNDR = 72.0 dB (11.7 bit)<br />
(b) with digital calibration<br />
Figure 6. Spectrum of the SAR ADC. (Voltage of<br />
input signals are 90 % of full scale voltage.)<br />
239
Table 2. Evaluation results of the AFE LSI<br />
power<br />
consumption<br />
(mW)<br />
effective<br />
chip size<br />
(mm 2 )<br />
Logic 0.0 0.27<br />
Vref generator 1.2 0.11<br />
Clock generator 0.2 0.0081<br />
PGA 2.5 0.30<br />
ADC 2.5 1.6<br />
Total 6.4 2.3<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
V. SUMMARY<br />
A low-power-consumption and highly accurate<br />
analog-front-end (AFE) LSI for sensor nodes was<br />
developed. Digital calibration expanded the operational<br />
margin of the analog circuits of the successive<br />
approximation register (SAR) analog-digital converter<br />
(ADC) in the LSI at low supply voltage. As a result, the<br />
proposed SAR ADC achieves both high effective resolution<br />
(11.7 bits) and extremely low power consumption (2.5<br />
mW); accordingly, it is suitable for sensor networks.<br />
ACKNOWLEDGMENT<br />
I am grateful to Dr. Takashi Oshima, who provided his<br />
carefully considered feedback and valuable comments. I<br />
also acknowledge the staff of the LSI Design and<br />
Fabrication Dept. of Hitachi ULSI Systems Co., Ltd. This<br />
work was supported by the New Energy and Industrial<br />
Technology Development Organization.<br />
REFERENCES<br />
[1] W. Liu, P. Huang, and Y. Chiu, “A 12b 22.5/45 MS/s 3.0 mW<br />
0.059 mm 2 CMOS SAR ADC Achieving Over 90dB SFDR,”<br />
ISSCC Dig. Tech. Papers, pp. 380-381, Feb. 2009.<br />
[2] J. L. McCreary and P. R. Gray, “All-MOS charge redistribution<br />
analog-to-digital conversion techniues—Part I,” IEEE J.<br />
Solid-State Circuits, vol. SC-10, no. 6, pp. 371–379, Dec. 1975.<br />
[3] R. Elfrink, et al., “First Autonomous Wireless Sensor Node<br />
Powered by a Vacuum-Packaged Piezoelectric MEMS Energy<br />
Harvester,” IEEE International Electron Devices Meeting,<br />
pp. 543-546, Dec. 2009.<br />
[4] International Technology Roadmap for Semiconductors 2009<br />
240
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Large Area Adaptative Fluidic Lens<br />
Solon Mias [1,2], Aurélien Bancaud [1,2], Henri Camon [1,2]<br />
1 CNRS-LAAS, 7 avenue du colonel Roche, F-31077 Toulouse<br />
2 University of Toulouse; UPS; INSA; INP; ISAE; LAAS; F-31077 Toulouse, France<br />
Abstract- We have developed a large area (30mm diameter)<br />
adaptive fluid-lens using simple fabrication procedures that do<br />
not require elaborate micro-fabrication techniques. The lens<br />
structure consists of a 2mm thick liquid reservoir which is<br />
sandwiched between a solid borosilicate glass substrate, a thick<br />
PDMS-elastomer side-wall spacer and a more flexible PDMScoated<br />
PET membrane. The reservoir is connected via a 0.3<br />
mm tube to a commercially available micro-pump used to alter<br />
the pressure within the reservoir, thus altering the surfacecurvature<br />
of the PET membrane and at the same time the<br />
optical power of the lens. The lens focal length can be changed<br />
from infinity to 0.5m.<br />
I. INTRODUCTION<br />
Ophthalmic glasses are one of the oldest portable devices.<br />
Even so, they have remained mostly the same over hundreds<br />
of years consisting of a rigid structure with fixed focal<br />
length. Recently, lenses with variable focal distances<br />
(bifocal and progressive lenses) have been developed<br />
particularly for people suffering from presbyopia [1].<br />
Presbyopia is a condition mostly occurring due to aging of<br />
the human eye and where the eye’s ability to accommodate<br />
is reduced [2]. Other recent developments include the<br />
creation of microstructures on a lens in order to produce a<br />
customized lens [3]. Even so, the above structures are rigid<br />
and their focal length and wavefront correction abilities are<br />
determined by the fabrication.<br />
Adaptive lenses on the other hand have the ability to tune<br />
their focal length according to the needs of the user. Many<br />
such devices have been developed using a variety of<br />
technologies such as liquid crystal devices [4-11],<br />
electrowetting devices [12, 13] and micro-fluidic devices<br />
[14-17]. Liquid crystal adaptive lenses rely on the<br />
birefringence of the liquid crystal mixtures used in order to<br />
create a refractive index modulation within the lens which<br />
then translates into a phase change of the propagating light.<br />
These lenses are limited by the birefringence and the<br />
thickness of the liquid crystal layer. In addition, they exhibit<br />
diffraction effects due to the electrodes used to spatially<br />
address the liquid crystal layer. Different electrode shapes<br />
[18] and variable resistivities [6] have been developed in<br />
order to reduce the diffraction effects but in the expense of<br />
greater fabrication complexity. Finally, the greatest problem<br />
with liquid crystal lenses is their polarization sensitivity.<br />
Therefore two liquid crystal layers are required in order to<br />
modulate all light-polarizations and this increases<br />
fabrication complexity even further, particularly due to the<br />
need of fine alignment between corresponding pixels on<br />
each layer.<br />
Electrowetting devices on the other hand do not suffer<br />
from polarization sensitivity. Two companies have been<br />
recently involved in the fabrication of electrowetting lenses,<br />
namely Varioptic [13] and Philips [12]. Both companies use<br />
two non miscible liquids; an aqueous conducting solution<br />
along with insulating oil of the same density. The liquids<br />
are inserted within a closed cell with appropriately placed<br />
electrodes. The use of two liquids instead of one (e.g. water<br />
in air) is necessary for suppressing any optical distortion of<br />
the gravity on the liquid-liquid interface. The angle of the<br />
conducting fluid with the cell wall changes when an electric<br />
field is applied to the cell. Hence, as the voltage changes,<br />
the curvature of the interface between the two liquids is also<br />
changed. Hence a lens is formed which can be tuned from<br />
convex to concave using appropriate voltage levels.<br />
Unfortunately, devices produced are limited to small active<br />
areas due to distortions on the interface between the two<br />
liquids. Therefore both Varioptic and Philips lenses are<br />
mostly destined for the mobile phone industry rather than<br />
ophthalmic optics.<br />
Fluid-lens devices are also polarisation insensitive. The<br />
tuning of their focal length is achieved by altering the<br />
pressure within a liquid reservoir which has at least one wall<br />
made out of a flexible membrane (usually PDMS). When<br />
the pressure inside the cell is below the atmospheric<br />
pressure, the device turns into a concave lens. As the<br />
pressure increases above the atmospheric pressure then the<br />
lens turns into a convex lens. The change of the pressure<br />
within the lens can be tuned using a micro-pump, a syringe<br />
or a volume changing material [15]. The diameter of the<br />
lenses produced are relatively small (200µm [19] to 20mm<br />
[17]). The uniformity of the devices can be a problem. For<br />
example, PDMS non-uniformity during the membrane<br />
fabrication can cause defects of up to 4µm [19]. Also if the<br />
membrane of the lens is too flexible then the weight of the<br />
liquid can distort the normally-spherical shape of the<br />
membrane when the lens is used in a non-horizontal<br />
position.<br />
In this publication we describe the fabrication of a large<br />
area adaptive fluid-lens using simple fabrication procedures<br />
that do not require elaborate micro-fabrication techniques.<br />
The membrane of the lens is made out of PDMS-coated<br />
241
11-13 <br />
May 2011, Aix-en-Provence, France<br />
PET layer which is solid relative to a single PDMS<br />
<br />
PDMS<br />
membrane of the same thickness. Therefore the lens<br />
membrane exhibits good uniformity despite its large area.<br />
spacer wall<br />
II. DEVICE FABRICATION<br />
Across section of the produced adaptive lens is shown in<br />
Fig.1. Firstly a 2mm thick PDMS layer is fabricated using a<br />
mixture of PDMS base 184 from Dow Corning and crosslinking<br />
agent at a mass ratio of 10 to 1. The largest part of<br />
the mixture is degassed within a cylindrical tank and leave<br />
more than one hour to allow a natural spread by gravity of<br />
the material into the tank. Thus we can get a very flat<br />
surface. Then the material is cured for an hour at 70°C<br />
within the cylindrical container for reticulation. The<br />
thickness is calculated by knowing the diameter of the<br />
cylindrical tank and the weight difference before and after<br />
filling. After reticulation the PDMS is removed from the<br />
container. A circular hole of 30mm is then opened through<br />
the spacer to form the reservoir side-walls. A 6mm channel<br />
long is also formed from the inner to the outer wall of the<br />
spacer in order to accommodate two tubes as shown in<br />
Fig.1(a). The tubes have inner diameter of 0.3mm and outer<br />
diameter of 1.5mm. These manufacturing steps can be<br />
easily enhanced by the completion of a resin mold (in SU-8<br />
for example) on a silicon or glass wafer. The remaining<br />
non-cured PDMS mixture is also degassed and then spin<br />
coated on to a 75µm thick PET layer at a speed of 1200 rpm<br />
for 20 seconds. (PDMS thickness on PET 50 µm) The<br />
PDMS-coated PET is also placed at 70°C for an hour in<br />
order to cure the PDMS coating. The reason for using the<br />
PET layer is for increasing the rigidity of the flexible<br />
membrane of the lens. This eliminates any non-uniformity<br />
on the membrane surface due to the fabrication and due to<br />
the weight of the liquid as the lens is placed vertically. The<br />
PDMS coating of the PET is used for adhesion purposes. To<br />
elaborate further, the PDMS spacer and the PDMS-coated<br />
PET are etched using oxygen plasma. The oxygen plasma of<br />
PDMS results in the creation of free radicals on the PDMS<br />
surface. After the plasma etching, the PDMS spacer is<br />
placed on top of the PDMS-coated PET so as to be in<br />
contact with the spin-coated PDMS layer. The present of the<br />
free radicals ensures a strong adhesion between the two<br />
PDMS surfaces.<br />
The Young’s modulus of PDMS is in the range of MPa<br />
(0,2 Mpa more precisely) while that of PET is about 3000<br />
Mpa, ie three decade above. As the thicknesses have the<br />
same order of magnitude (50 µm for PDMS and 75 µm for<br />
PET), the deformation is overwhelmingly governed by the<br />
properties of PET.<br />
The sealed reservoir is then filled with de-ionized water<br />
using a commercially available micro-pump, produced by<br />
Bartels Mikrotechnik GmbH. The micro-pump is connected<br />
to one of the two tubes and the second tube is used for<br />
allowing the air to exit from the sealed reservoir. The<br />
introduction of the liquid in the reservoir is shown in Fig.2.<br />
Fig.2 (d) shows that no air remains within the reservoir at<br />
the end of the process. The internal dimensions of the<br />
system are large enough not to ask problems due to<br />
capillary forces.<br />
Glass<br />
Substrate<br />
PDMS<br />
coating<br />
PDMS<br />
spacer wall<br />
Glass Substrate<br />
Tubes<br />
(a)<br />
Liquid<br />
reservoir<br />
PET<br />
(b)<br />
Liquid reservoir<br />
Fig. 1. (a) Top view of the fabricated fluid-lens. (b) Cross-section of the<br />
fluid-lens along the dotted line<br />
(a)<br />
(c)<br />
(b)<br />
(d)<br />
Fig. 2. The introduction of DI-water in the sealed reservoir. The DI-water<br />
is pumped out of a bottle (shown at the bottom-left of the images) using a<br />
micro-pump. The water is introduced into the reservoir using the one of the<br />
two tubes (water inlet tube) while the air already present inside the<br />
reservoir is allowed to exit the lens via the second tube (air outlet tube).<br />
Images (a) to (d) show the reservoir being progressively filled with DIwater.<br />
Image (d) shows that no air remains into the reservoir.<br />
III. DEVICE OPERATION AND CHARACTERIZATION<br />
After the spacer and the membrane have been bonded<br />
together, the combined spacer/membrane layer is again<br />
etched using oxygen plasma along with a borosilicate glass<br />
substrate. After the plasma etching, the tubes are placed<br />
within the 6mm channel and the spacer/membrane layer is<br />
then bonded onto the glass-substrate thus forming the liquid<br />
reservoir. Non-cured PDMS mixture is then used to fill the<br />
gaps between the channel and the tubes and the complete<br />
device is then heated at 70°C for an hour in order to cure the<br />
PDMS and seal the channel - and thus the reservoir.<br />
The adaptive fluid-lens was used under different<br />
operational modes. In the first mode the micro-pump was<br />
activated in order to raise the water pressure inside the<br />
reservoir. The air-outlet tube was left open and hence water<br />
was allowed to exit through the outlet tube. In this mode of<br />
242
operation, the pressure inside the reservoir is increased due<br />
to the continuous on-flow of the liquid, but remains below<br />
the maximum possible pressure due to the open outlet tube.<br />
Thus the membrane surface curvature is not the maximum<br />
possible. This mode of operation was used as a “safe mode”<br />
in order to observe if the reservoir was well sealed and<br />
avoid any damage on the PDMS-sealed channel. In the<br />
second mode of operation the outlet tube was sealed using a<br />
plastic clip. This mode of operation was used in order to<br />
obtain the maximum pressure within the reservoir and hence<br />
the minimum focal length possible. In both modes of<br />
operation the micro-pump flow was maintained at 5ml/min<br />
which is the maximum available with this type of micropump.<br />
The adaptive fluid-lens was characterized in both modes<br />
of operation either qualitatively (using direct optical<br />
observation via a digital camera) or quantitatively (using a<br />
specially designed optical laser set-up). The results of the<br />
direct optical observation under “safe mode” operation are<br />
shown in Fig.3. The images show clearly an increase in the<br />
lens optical power. Fig.4 shows how the optical power of<br />
the lens is increased when the device operates at maximum<br />
pressure.<br />
A specially designed laser set-up was used in order to<br />
obtain the focal length range of the adaptive lens. In the setup<br />
(shown in Fig.5) a 632nm He-Ne laser beam was<br />
allowed to propagate through the lens. The beam was then<br />
collected and observed via a screen placed at 1,06m away<br />
from the lens. The laser was mounted on a fix frame while<br />
the lens was allowed to move perpendicularly to the<br />
incoming laser beam. The calculation of the focal length<br />
range was made by observing the deviation of the laser spot<br />
on the screen as the lens is moved from right to left. When<br />
the laser beam coincides with the optical axis of the lens<br />
then the spot on the screen remains fixed no matter what the<br />
pressure in the liquid reservoir. The lens is then moved<br />
perpendicularly to the laser beam. Hence the laser beam<br />
remains parallel to the optical axis of the lens. If the lens is<br />
not pressurized, then the spot on the screen remains at the<br />
same position as before. When the pressure in the fluidreservoir<br />
is changed then the spot on the screen deviates due<br />
to the refractive properties of the pressurized lens. The focal<br />
length of the lens can thus be calculated using the diagram<br />
of Fig.6, where f is the focal length, e is the deviation of the<br />
spot; d is the distance traveled by the lens and r is the<br />
distance from the lens to the screen. Using the similar<br />
triangles OAB and OCD we derive that,<br />
(d-e)/d=(f-r)/f. (1)<br />
Using simple mathematical manipulation we find that the<br />
focal length is given by,<br />
f=dr/e. (2)<br />
(a)<br />
(b)<br />
Fig. 3. Image magnification using the adaptive fluid-lens in “safe mode”<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
operation.<br />
Fig. 4. Image magnification using the adaptive fluid-lens at maximum<br />
pressure (d). Note that the image axis is not parallel to the optical axis of<br />
the lens.<br />
Fig. 5. Optical set-up used for calculating the focal length range of the<br />
adaptive lens. Part (b) is a close-up image of the laser-lens set-up. MP is<br />
the micro-pump, and MPD is the micro-pump driver.<br />
Fig. 6. Diagram used for calculating the focal length range of the adaptive<br />
lens. f is the focal length, r is the distance from the lens to the screen, d is<br />
the distance traveled by the lens and e is the deviation of the spot on the<br />
screen.<br />
For example, when the non-pressurized lens was moved<br />
by d=8mm then the spot on the screen remained unmoved.<br />
Therefore e=0mm and f→∞. When the lens was operated<br />
under “safe mode” operation, the spot deviation on the<br />
screen was e=8mm. Therefore, the focal length was<br />
f=1.06m. On the other hand when the lens was operated<br />
under maximum pressure, then the deviation of the spot was<br />
as high as 15mm. In this case the focal length was 0.565m.<br />
243
It is important to note that even in the later case the pressure<br />
in the reservoir was maintained below the maximum<br />
possible in order to avoid damaging the lens. Fig.7 shows<br />
how the laser spot moves on the screen as the pressure in<br />
the liquid reservoir is increased above the atmospheric<br />
pressure.<br />
(a)<br />
(b)<br />
(c)<br />
(d)<br />
(e)<br />
Fig. 7. The laser spot movement on the screen as the pressure in the<br />
liquid reservoir is increased from atmospheric (a) to the maximum pressure<br />
(f) sustained via the micro-pump. The incident laser beam is parallel to the<br />
optical axis of the lens but is fired 8mm from the center of the lens.<br />
IV. CONCLUSIONS<br />
We have demonstrated the fabrication and<br />
characterization of a large area adaptive fluid-lens. The<br />
fabrication process is extremely simple as no nanofabrication<br />
techniques are required. The focal-length range<br />
of the adaptive lens varies from infinity to 0.565m<br />
depending on the pressure within its fluid reservoir. Such<br />
adaptive lenses can find use in ophthalmic optics (digital<br />
eye-glasses for example), astronomical optics (because the<br />
gravity will not affect the membrane-curvature of the lens)<br />
or telecoms applications (beam steering and fiber to fiber<br />
connections). In future, the fabrication steps could be easily<br />
enhanced and the reliability of some aspects have to be<br />
investigated like delaminating between layers, the possible<br />
evolution of the mechanical properties of PDMS and the<br />
tightness of the structure because the PDMS is known to be<br />
porous.<br />
REFERENCES<br />
[1]. K. Krause, “Acceptance of progressive lenses”, Klin Monatsble<br />
Augenheilkd. 209 (2-3), 1996, pp 94-99.<br />
[2]. G. Li et al., “Switchable electro-optic diffractive lens with high<br />
efficiency for ophthalmic applications”, in Proceeding of the<br />
National Academia of Sciences of U S A, vol. 103(16), 2006, pp.<br />
6100-6104.<br />
[3]. Y. Liu, L. Warden, K. Dillon, G. Mills, A. Dreher, “Z-View<br />
diffractive wavefront sensor: principle and applications”, In<br />
Proceeding of SPIE, 6018, 2005, pp. 78-86.<br />
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[4]. T. Nose, S. Masuda, and S. Sato, “A Liquid Crystal Microlens<br />
with Hole-Patterned Electrodes on Both Substrates”, Japanese<br />
Journal of Applied Physics, vol. 31(Part 1, No. 5B), 1992, pp.<br />
1643-1646.<br />
[5]. S. Sato, “Liquid-Crystal Lens-Cells with Variable Focal Length”,<br />
Japanese Journal of Applied Physics, vol. 18 (9), 1979, pp. 1679-<br />
1684.<br />
[6]. M. Y. Loktev, V. N. Belopukhov, F. L. Vladimirov, G. V.<br />
Vdovin, G. D. Love, and A. F. Naumov, “Wave front control<br />
systems based on modal liquid crystal lenses”, Review of<br />
Scientific Instruments, vol. 71(9), 2000, pp. 3290-3297.<br />
[7]. S. Sato, A. Sugiyama, and R. Sato, “Variable-Focus Liquid-<br />
Crystal Fresnel Lens”, Japanese Journal of Applied Physics, vol.<br />
24 (8, Part 2), 1985, pp. 626-628.<br />
[8]. S. T. Kowel, P. Kornreich, and A. Nouhi, “Adaptive spherical<br />
lens”, Applied Optics, vol. 23, 1984, pp. 2774-2777.<br />
[9]. Y. Takaki and H. Ohzu, “Liquid-crystal active lens: a<br />
reconfigurable lens employing a phase modulator”, Optics<br />
Communications, vol. 126(1-3), 1996, pp. 123-134.<br />
[10]. L. N. Thibos and A. Bradley, “Use of liquid-crystal adaptiveoptics<br />
to alter the refractive state of the eye”, Optometry and<br />
Vision Science, vol. 74(7), 1987, pp. 581-587.<br />
[11]. V. Presnyakov, K. Asatryan, T. Galstian, and A. Tork, “Polymerstabilized<br />
liquid crystal for tunable microlens applications”,<br />
Optics Express, vol. 10(17), 2002, pp. 865-870.<br />
[12]. S. Kuiper and B. H. W. Hendriks, “Variable-focus liquid lens for<br />
miniature cameras”, Applied Physics Letters, vol. 85(7), 2004, pp.<br />
1128-1130.<br />
[13]. C. Gabay, B. Berge, G. Dovillaire, and S. Bucourt, “Dynamic<br />
study of a Varioptic variable focal lens”, In Proceeding of SPIE<br />
4767, (2002), pp. 159-165.<br />
[14]. K.-H. Jeong, G. L. Liu, N. Chronis, and L. P. Lee, “Tunable<br />
micro-doublet lens array”, Optics Express , vol. 12(11), (2004),<br />
pp. 2494-2500.<br />
[15]. L. Dong, A. K. Agarwal, D. J. Beebe, and H. Jiang, “Adaptive<br />
liquid microlenses activated by stimuli-responsive hydrogels”,<br />
Nature, vol. 442, 2006, pp. 551-554.<br />
[16]. M. Agarwal, R. A. Gunasekaran, P. Coane, and K. Varahramyan,<br />
“Polymer-based variable focal length microlens system”, Journal<br />
of Micromechanics and Microengineering , vol. 14(12), 2004, pp.<br />
1665-1673.<br />
[17]. D.-Y. Zhang, N. Justis, and Y.-H. Lo, Fluidic adaptive lens of<br />
transformable lens type, Applied Physics Letters, vol. 84(21),<br />
2004, pp. 4194-4196.<br />
[18]. H. Ren, Y.-H. Fan, S.-T. Wu, “Adaptive liquid crystal lenses”,<br />
US patent No: 6,859,333 B1 , February 22, 2005.<br />
[19]. N. Chronis, G. Liu, K.-H. Jeong, and L. Lee, “Tunable liquidfilled<br />
microlens array integrated with microfluidic network”,<br />
Optics Express, vol. 11(19), 2003, pp. 2370-2378.<br />
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<br />
Fabrication and Characteristics of a Fused<br />
Silica-Based Optical Waveguide with Femtosecond<br />
Fiber Laser Pulses<br />
Ting-Chou Chang 1 , Chien-Hsing Chen 2 , Wei-Hung Shih 3 , Jian-Neng Wang 4 , Chai-Yu Lee 1 , Jaw-Luen Tang 2 , Shau-Chun<br />
Wang 1 , Lai-Kwan Chau 1 , Wei-Te Wu 5*<br />
1 Department of Chemistry and Biochemistry, National Chung Cheng University<br />
168 University Road, Minhsiung, Chiayi 621, Taiwan<br />
2 Department of Physics, National Chung Cheng University<br />
168 University Road, Minhsiung, Chiayi 621, Taiwan<br />
3 Department of Mechanical Engineering, National Chung Cheng University<br />
168 University Road, Minhsiung, Chiayi 621, Taiwan<br />
4 Department of Construction Engineering, National Yunlin University of Science and Technology,<br />
123 University Road, Section 3, Douliou, Yunlin 640, Taiwan<br />
5* Department of Biomechatronics Engineering, National Pingtung University of Science and Technology<br />
1, Shuefu Road, Neipu, Pingtung 912, Taiwan<br />
Tel: +886-8-770-3202 Ext. 7599; Fax: + 886-8-774-0420; weite@mail.npust.edu.tw<br />
Abstract<br />
This study investigates the fabrication characteristics<br />
of a femtosecond fiber laser on a fused-silica-based optical<br />
waveguide. The wavelength and repetition rate of the<br />
femtosecond fiber laser are 532 nm and 1 MHz,<br />
respectively. We selected three main fabrication<br />
parameters for systematic adjustment: laser power (E),<br />
scanning speed ( v s<br />
) and focus depth (d = 0 at the surface<br />
of substrate). We succeeded in fabricating a waveguide<br />
layer inside the silica subtracts. By analyzing the light<br />
translation path and the net fluence in the waveguide, the<br />
range of fabrication energy of the waveguide on the fused<br />
silica was kept within 0.973 - 1.438 KJ/cm 2 .<br />
I. Introduction<br />
Recently, developments in nanotechnology have led to<br />
a proliferation of electro-optic system applications. To<br />
minimize system size, industries including communications,<br />
construction and biomedical detection have widely applied<br />
optical waveguides such as photonic crystal fibers [1], fiber<br />
interferometers [2], surface plasma resonance (SPR) sensors<br />
[3], localized plasma resonance (LPR) sensors [4] and<br />
guided-mode resonance (GMR) sensors [5].<br />
Waveguide device are fabricated through techniques<br />
including ion bombardment, laser machining,<br />
photolithography, and mechanical stamping [6], commonly<br />
using fused silica as a substrate. Laser machining is a low cost,<br />
high speed and high yield method for the localized heat<br />
treatment of fused silica. However the linear absorption of<br />
fused silica depends on the laser source. Using an ultraviolet<br />
laser requires a process to bind oxygen to the fused silica to<br />
increase light sensitivity [7]. Using CO 2 laser [8] results in a<br />
greatly increased linear absorption of the fused silica which<br />
makes precise machining more difficult and can cause<br />
damage around the machining area. High-power density<br />
femtosecond fiber lasers with a pulse of 10 -15 seconds are an<br />
appropriate tool for the fabrication of optical waveguides due<br />
to their independence in the linear absorbing effect of fused<br />
silica.<br />
This study investigates the fabrication characteristics of<br />
femtosecond fiber lasers on fused-silica-based optical<br />
waveguides. We selected three main fabrication parameters,<br />
laser power (E), scanning speed ( v s<br />
) and focus depth (d = 0 at<br />
the surface of substrate) which are systematically adjusted to<br />
investigate the differences of post-machining light waveguide<br />
characteristics, transmission loss rate and the relation<br />
between the net influence and light waveguide.<br />
II. Experimental section<br />
1. Waveguide principles<br />
As shown in Fig. 1, the light waveguide is composed of<br />
a layer of Media 1 (i.e. a media different from the substrate)<br />
sandwiched between two layers of Media 2 (i.e. the<br />
substrate).<br />
One of two application phenomena of light waveguides<br />
is the refraction within these media with different refraction<br />
indices. Based on the Snell’s law, the refraction angle, φ , is<br />
smaller than the incident angle, θ , as light is incident into<br />
Media 1. The other application phenomenon is total reflection<br />
for keeping and transmitting all laser energy within the Media<br />
1 layer. This means that Snell’s law requires the refraction<br />
index of Media 1, n 1 , to be larger than that of Media 2.<br />
The numerical aperture (NA), (i.e., the maximum<br />
acceptable energy of light wave guide), is defined as.<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
245
NA sinθ c<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
2 2<br />
<br />
1<br />
−=≡<br />
nn (1) rotating the half-wavelength and polarization slides. The laser<br />
2<br />
is focused by an objective lens (Mitutoyo-10X,NA=0.28).<br />
The charge-coupled device (CCD) is used to help aim the laser<br />
on the machining area to ensure beam quality.<br />
The machining path and machining rate are controlled by<br />
programming the X-Y micro-positioning platform. An optical<br />
microscope is used to inspect the machined products.<br />
where θ c is the maximum acceptance angle.<br />
The light among the incident light increases with NA. If<br />
the incident angle is larger than θ c , some light is refracted<br />
into Media 2. Therefore, the incident angle must be smaller<br />
thanθ c to satisfy the total reflection and forming guide mode.<br />
Air, n air<br />
n 2<br />
media 2<br />
n 1<br />
media 1<br />
n 2<br />
III. Results and Discussion<br />
1. Waveguide fabrication<br />
In this study we selected three main fabrication<br />
parameters, laser power (E), scanning speed ( v s<br />
) and focus<br />
depth (d = 0 at the surface of substrate). By fixing the laser<br />
power at 170 mW and the focus depth at 0 μm, the fused silica<br />
was modified at scanning speeds.<br />
media 2<br />
Fig. 1 Waveguide translation principle<br />
In general, the fused silica is homogeneous with the<br />
constant refraction index. However, the refraction index of<br />
fused silica increases with the annealing rate [9]. The<br />
femtosecond laser’s pulse characteristic makes it appropriate<br />
for decreasing the annealing rate. The pulse energy does not<br />
integrated easily in the working area and results in a lower<br />
annealing rate.<br />
5.1μm<br />
(a)<br />
1mm/s<br />
4.1μm<br />
3.0μm<br />
2.7μm<br />
(b) (c) (d)<br />
2mm/s 3mm/s 4mm/s<br />
2. Experimental Setup<br />
The specifics of the apparatus used in this study are<br />
shown in Table 1 and Fig. 2. The central wavelengths of the<br />
laser are 532 and 1064 nm, the pulse duration is less than 400<br />
fs and the repetition rate is 1 Hz – 1 MHz. The laser beam is a<br />
Gaussian beam.<br />
Fig. 2 Femtosecond fiber laser machining system schematic<br />
Table 1 Femtosecond fiber laser machining system<br />
specification<br />
Wavelength 1064 nm & 532 nm<br />
Repetition rate<br />
1 Hz~1 MHz<br />
Pulse duration<br />
increased.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Power meter<br />
LD@1553nm<br />
collimator<br />
3.6μm<br />
2.5μm<br />
MMF-fiber<br />
(a)<br />
1mm/s<br />
(b)<br />
2mm/s<br />
(c)<br />
3mm/s<br />
Fig. 4 Fabrication with various scanning speeds at E = 170 mW and<br />
d = 10 μm<br />
3.6μm<br />
(a)<br />
5mm/s<br />
3.7μm<br />
(d)<br />
8mm/s<br />
3.3μm<br />
(b)<br />
6mm/s<br />
3.3μm<br />
(e)<br />
9mm/s<br />
2.3μm<br />
(c)<br />
7mm/s<br />
2.5μm<br />
(f)<br />
10mm/s<br />
Fig. 5 Fabrication with various scanning speeds at E = 170<br />
mW and d = 0 μm<br />
The laser power was increased to 230 mW to modify fused<br />
silica 10 μm in depth. Scanning speed should be increased to<br />
avoid surface ablation. The results in Fig. 5 show that the<br />
fused silica is ablated given scanning speeds between 5 mm/s<br />
and 7 mm/s; and modified widths of3.7μm, 3.3μm and 2.5μm<br />
correspond to scanning speeds of 8 mm/s, 9 mm/s and 10<br />
mm/s, proving that the fused silica can be modified at different<br />
depths through focusing and tuning the laser power and<br />
scanning speed.<br />
2. Waveguide propagating loss measurement<br />
Fig. 6 shows the system for measuring waveguide<br />
propagating loss. The system conducts the laser diode (LD,<br />
center wavelength = 1553 nm) to the waveguide layer on the<br />
XYZ-rotation stage. In the end of waveguide layer, the<br />
collimator couples the multi-mode optic fiber to the power<br />
meter for acquiring and analyzing signal. The results show<br />
that the propagating loss are 4.6 dB/cm、4.8 dB/cm、6.2<br />
dB/cm as the scanning velocities are 8 mm/s, 9 mm/s and 10<br />
mm/s, respectively, with 230 mW and 10 μm of depth. It<br />
indicates that the increased scanning velocity causes the<br />
larger energy loss due to the low absorbing energy of fused<br />
silica.<br />
XYZ-rotation<br />
stage<br />
Transmission (dBm)<br />
0<br />
-5<br />
-10<br />
-15<br />
-20<br />
-25<br />
-30<br />
-35<br />
-40<br />
LD<br />
-45<br />
1553.2 1553.4 1553.6 1553.8 1554.0 1554.2<br />
Wavelength (nm)<br />
Fig. 6 The system for measuring waveguide propagating loss<br />
Table 2 Fabrication parameters of waveguide using<br />
femtosecond laser<br />
Laser<br />
power<br />
E<br />
(mW)<br />
170<br />
230<br />
Scanning focusing<br />
velocity depth<br />
NF<br />
results<br />
v<br />
s d<br />
(KJ/cm 2 )<br />
(mm/s) (μm)<br />
1<br />
ablation 7.191<br />
2 ablation 3.596<br />
3 ablation 2.397<br />
4 0 ablation 1.798<br />
5 waveguide 1.438<br />
6 waveguide 1.198<br />
7 waveguide 1.027<br />
5<br />
ablation 1.946<br />
6 ablation 1.621<br />
7 ablation 1.390<br />
10<br />
8 waveguide 1.216<br />
9 waveguide 1.081<br />
10 waveguide 0.973<br />
3. Waveguide discussion<br />
The laser power, the diameter of the laser beam, the<br />
scanning speed and the rate of repetition are all very<br />
influential factors in laser machining. This study analyzes the<br />
machining performance with NF factor [10], as shown below.<br />
2ω0<br />
PRF<br />
NF =<br />
(5)<br />
vs<br />
where ω is the minimal radius of the laser beam, R = 1<br />
0<br />
MHz is the repetition rate, and E<br />
F p<br />
= is the average<br />
2<br />
Rπω 0<br />
fluence per lasing. In this study, the laser wavelength, λ , is<br />
532 nm. The focus distance of the lens, f , is 20 mm. The<br />
diameter of the incident laser, D, is 5 mm. The ω is 1.5 μm,<br />
0<br />
estimated by the 1.05 of the measured laser beam quality<br />
factor ( D<br />
M<br />
2 πω0<br />
= ). E is the laser power. Substituting these<br />
2λf<br />
parameters into Eq.(5), the range of the NF value is 1.438 –<br />
0.973 KJ/cm 2 , indicating the range of fabrication energy of<br />
247
the waveguide on the fused silica, as shown in Table 2 and Fig.<br />
7. Furthermore, based on Table 2, the laser power should be<br />
increased with the scanning speed, thus increasing machining<br />
speed.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
silica is kept within 1.438 – 0.973 KJ/cm 2 . Future research<br />
could study parameters such as NA, NF and transmission loss<br />
in optimal methods to develop new design applications for<br />
biochemistry sensors and micro-optic systems.<br />
Modification type<br />
Waveguide<br />
No change<br />
Damage<br />
d=0μm<br />
d=10μm<br />
1 2 3 4 5 6 7<br />
NF (kJ/cm 2 )<br />
Fig. 7 The modification type with NF variation<br />
In this study, a waveguide fabricated with 170 mW laser<br />
power, 5mm/s scanning speed and 0 μm focus depth<br />
successfully conducted light as shown in Fig. 8, proving that<br />
femtosecond lasers can be used to fabricate waveguides on<br />
fused silica.<br />
Fig. 8 Waveguide fabricated with 170 mW laser power,<br />
5mm/s scanning speed and 0 μm focus depth<br />
IV. Conclusions<br />
This study describes the successful fabrication of a<br />
fused-silica-based optical waveguide using a femtosecond<br />
fiber laser, and an investigation of the fabrication<br />
characteristics of femtosecond fiber lasers on<br />
fused-silica-based optical waveguides. The results show that<br />
the modified width decreases with increasing scanning speed,<br />
regardless of machining depth. By analyzing the light<br />
translation path and the net fluence in the waveguide, the<br />
range of fabrication energy of the waveguide on the fused<br />
References<br />
1. C. H. Chen, S. C. Chen, Y. C. Chen, H. T. Hu, T. H.<br />
Wei, W. T. Wu, J. N. Wang and J. L. Tang, Research<br />
on laser-induced long-period fiber grating sensor<br />
modified with gold nano-rods, The 8th Pacific Rim<br />
Conference on Lasers and Electro-Optics, Shanghai,<br />
2009<br />
2. Chien-Hsing Chen, Yi-Chun Chen, Jian-Neng Wang,<br />
Lai-Kwan Chau, Jaw-Luen Tang and Wei-Te Wu,<br />
“Multimode fiber Mach–Zehnder interferometer for<br />
measurement of refraction index”, IEEE Sensors<br />
2010 Conference - the 9th Annual IEEE Conference<br />
on Sensors, 2010/11/1-2010/11/4, USA.<br />
3. Y. Liu, J. Kim, Numerical investigation of finite<br />
thickness metal-insulator-metal structure for<br />
waveguide-based surface plasmon resonance<br />
biosensing, Sens. and Actu. B, Vol. 148, pp. 23-28,<br />
2010.<br />
4. L. K. Chau, Y. F. Lin, S. F. Cheng, and T. J. Lin,<br />
Fiber-optic chemical and biochemical probes based<br />
on localized surface plasmon resonance, Sens. and<br />
Actu. B, Vol. 113, pp. 100–105, 2006.<br />
5. Ian D. Block, Nikhil Ganesh, Meng Lu, and Brian T.<br />
Cunningham, ”Bulk-Micromachined Optical Filer<br />
Based on Guided-Mode Resonance in Silicon-Nitride<br />
Membrane,” IEEE Sens. J., Vol. 8, pp.274-280, 2008.<br />
6. C. S. Ma, W. B. Guo, D. M. Zhang, K. X. Chen, Y.<br />
Zhao, F. Wang, Z. C. Cui, S. Y. Liu, Analytical<br />
modeling of loss characteristics of a polymer arrayed<br />
waveguide grating multiplexer, Vol. 34 PP. 621-630,<br />
2002.<br />
7. C. Chen, X. Sun, D. Zhang, Z. Shan, S. Y. Shin, D.<br />
Zhang, Dye-doped polymeric planar waveguide<br />
devices based on a thermal UV-bleaching technique,<br />
Optics & Laser Technology, Vol. 41 , pp. 495–498,<br />
2009.<br />
8. A. M. Vengsrlar, P. J. Lemaire, et al. “Long-Period<br />
Fiber Gratings as Band-Rejection Filters,” Journal of<br />
Lightwave Technology, vol. 4, pp. 58-65, 1996.<br />
9. J. W. Chan, T. R. Huser, S. H. Risbud, J. S. Hayden, D.<br />
M. Krol, Waveguide fabrication in phosphate glasses<br />
using femtosecond laser pulses, APPLIED PHYSICS<br />
LETTERS, Vol. 82, pp. 2371-2373, 2003.<br />
10. L. Shah, Y. A. Arai, S. M. Eaton, P. R. Herman,<br />
Waveguide writing in fused silica with a femtosecond<br />
fiber laser at 522 nm and 1 MHz repetition rate,<br />
Optics Express, Vol. 13, pp. 1999-2006, 2005.<br />
248
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
A Multilevel Polymer Process for Liquid Direct<br />
Encapsulation for Opto-Fluidic Application<br />
Remy Bossuyt [1, 2, 3], Laurent Mazenq[1, 2], Véronique Conédéra [1, 2], Jérôme Ballet [1], Anne-Marie Gué [1, 2],<br />
Jean-Paul Cano [3] and Henri Camon [1, 2]<br />
1<br />
CNRS-LAAS, 7 avenue du colonel Roche, F-31077 Toulouse<br />
2<br />
University of Toulouse; UPS; INSA; INP; ISAE; LAAS; F-31077 Toulouse, France<br />
3<br />
Essilor International, rue Pierre et Marie Curie, 31670 Labège, France<br />
Abstract- This paper describes a polymer manufacturing<br />
technique for the realization of sealed liquid tanks with a high<br />
filling ratio. It is based on a lithographic process combined<br />
with lamination and can be performed on any type of<br />
substrate: Si, glass or polymer films. We demonstrated the<br />
encapsulation of liquid in multilevel tanks on large surface<br />
area (>118 cm 2 ) without any damage of the stored liquid and<br />
good flatness. This micro-fabrication technology proposes a<br />
significant breakthrough for opto-fluidic applications.<br />
We propose a one step process for the filling and sealing of<br />
tanks which are performed simultaneously. Tanks can be<br />
integrated at any level in the micro-fluidic system with a<br />
perfect filling rate and an excellent tightness.<br />
I. INTRODUCTION<br />
The encapsulation of liquids in miniaturized systems is a<br />
major challenge for a wide range of applications [1, 2, 3, 4].<br />
In most of cases, the encapsulation process necessitates a<br />
final sealing step which cannot be performed collectively<br />
and which imposes drastic limits on integration possibilities<br />
and performances (architectures, number and localization of<br />
reservoirs, fabrication cost …).<br />
The sealing technology previously developed at LAAS<br />
has been designed to close empty channels separated by<br />
distances greater than 500µm. The photo resist used was the<br />
well known SU-8. That allowed making channel networks<br />
in three dimensions (3D) by level superposition in a quite<br />
simple way [5]. In our research works, it came the need to<br />
trap a high liquid volume on multilevel, quickly and without<br />
any specific filling through a channel networks which<br />
consume the surface area and consequently decrease the<br />
surface filling ratio.<br />
In this paper, we propose process with one step which<br />
overcomes these major drawbacks. The filling and sealing<br />
of tanks are performed simultaneously. Tanks can be<br />
integrated at any level in the micro-fluidic system with a<br />
perfect filling rate and an excellent hermeticity. A new<br />
negative resist, epoxy based, is employed here. This photo<br />
resist is more transparent than the SU-8 and presents a good<br />
mechanical and chemical resistance once polymerized; this<br />
is important for a good liquid trapping. But we also take<br />
great advantage of the properties of this photo resist. The<br />
first one is a decrease of Soft-Bake (SB) and Post Exposure<br />
Bake (PEB) temperature (50°C instead of 90°C for Su-8)<br />
which reduces stress in structures and consequently<br />
deformations. The second advantage is a sticky surface after<br />
SB allowing an easier process for sealing.<br />
II. TECHNOLOGICAL PROCESS DESCRIPTION<br />
The goal of this study was to realize multilevel systems<br />
with high surface filling ratio. Each level would cover 50%<br />
of the visible surface by the liquid. The multi level then<br />
allows the rate to rise 100% in a checkerboard-like structure<br />
that will serve as an example for demonstration in this<br />
paper. The surface filling ratio or coverage rates of each<br />
level may vary depending on the intended applications. The<br />
final structure must be sufficiently transparent for<br />
observation of liquids or for use in transmission. In<br />
addition, it must be carried over to a host structure and thus<br />
be strippable.<br />
The general description of the complete technological<br />
process is given in Fig. 1. The micro fabrication process is<br />
described in four main parts: fabrication of the removable<br />
PET film onto a wafer, fabrication chessboard-like<br />
structure, filling and sealing, and finally fabrication of the<br />
second level. This process is performed on 6 inches glass<br />
wafer.<br />
Stack-up process description<br />
Step 1: Preparation of the<br />
removable plastic film<br />
Substrat<br />
Step 2: pillars of first level<br />
substrat<br />
Step 3: filling and sealing<br />
of first level by lamination<br />
liquid<br />
Step 4: fabrication of the second<br />
level<br />
liquid<br />
liquid<br />
Fig. 1. Schematic description of the main process steps for the fabrication<br />
of a two-level structure.<br />
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11-13 <br />
May 2011, Aix-en-Provence, France<br />
A. First step: preparation of the flexible substrate.<br />
<br />
The structures are formed on a glass substrate and must<br />
Step map<br />
be removable after manufacture. For this first step is to<br />
develop an interface allowing the realization of this<br />
operation. First and foremost, a combination of layers of<br />
4th step<br />
1st step<br />
glues with a high and low adhesive strength must be<br />
deposited on the glass substrate which will be used only as a<br />
mechanical support. This fabrication is operated by<br />
lamination. Then, a flexible sheet of plastic film (PET, 125<br />
µm) is laminated on the top of the glued glass wafer as<br />
illustrated in Fig. 2. The speed and pressure applied for the<br />
6 inch<br />
different laminations have to be chosen carefully. In order<br />
to limit wafer deformation, a low speed and a low pressure<br />
glass<br />
have to be applied. We also used a thick substrate (1,1 mm<br />
32 th step<br />
against 0,55 mm which is a common thickness for 6 inches<br />
wafer<br />
or less wafers).<br />
The plastic film surface is then activated by oxygen<br />
6 inches wafer<br />
plasma (P=200W, t=120s, V=200ml) in order to promote<br />
the adhesion of the photo resist during the following<br />
lithographic step and especially during and after the<br />
revelation of the photo resist. The activation surface process<br />
must be a low temperature process, firstly to avoid<br />
damaging definitively the surface of the plastic film and<br />
Substrat<br />
secondly to lower stress in the multilayer structure Fig. 3. Chessboard structures in top view (up) and a cross-section (down).<br />
(glass/paste/plastic film). Fig. 2 illustrates that first step.<br />
At the end, a cleaning with deionized water is performed<br />
to remove possible dusts at the surface. Now the wafer is<br />
ready to be micro structured, that is described in the next<br />
chapter.<br />
plastic film<br />
paste<br />
glass wafer 1,1 mm<br />
Fig. 2. After the deposition of a bi-layer of glue on a thick glass wafer, a<br />
plastic film of PET is then laminated.<br />
B. Second step: fabrication of the first structured level.<br />
Using a standard photolithographic process a first level of<br />
structures is realized on the plastic film. The microstructure<br />
design is a chessboard with a 600 µm period (Fig. 3). A 20<br />
µm thick resist is coated and soft baked at T= 50 °C during<br />
14 minutes in an oven. The photolithography is performed<br />
using a Canon FPA 3000 i4 stepper. With this tool, the<br />
mask design is scaled down by a factor of five. The figure is<br />
repeated 32 times on the wafer (Fig. 3.). Each figure is<br />
stitched to the next one in order to see only one figure at the<br />
end of the process. The stress increases during the spin<br />
coating of the resist, the UV exposition, the SB and PEB. It<br />
can induce a torsion in the structure that can make the<br />
alignment process impossible (in our example, the second<br />
level of microstructures will be aligned on the first level).<br />
This phenomenon is well known by the microlithography<br />
users and can be limited by several common ways such as<br />
low SB and PEB.<br />
Fig. 4. SEM view of the first level of the chessboard structure before filling<br />
and sealing.<br />
After exposure, a PEB is performed at 60°C during 8<br />
minutes in an oven. The revelation is achieved in a bath of<br />
isopropanol during 150 s. The figure 4 illustrates a top view<br />
of the first level before filling and sealing. At this stage, the<br />
first level is completely realized. It consists of a set of block<br />
(300 by 300 microns) over an area of 118 sq. cm ready for<br />
filling and sealing.<br />
C. Third step: filling and sealing of the first micro<br />
structured level.<br />
Tanks micro-structured at the 2 nd step (Fig. 3 and 4) have<br />
to be filled and sealed. These two operations are made at the<br />
same time. There are several conditions that the material to<br />
use to make the cover has to respect. It has to be strong<br />
enough to support (at least) a second level of<br />
microstructures, it has to be rigid enough to keep its shape<br />
during the covering process especially when large tanks are<br />
designed and it also has to have a thermal expansion<br />
250
coefficient compatible with the microstructure material. The<br />
photo resist used for the microstructures is a negative one<br />
epoxy based. After the polymerization, the photo resist is<br />
rigid enough to be used as a cover. In this way there is no<br />
mechanical incompatibility because the same material has<br />
been used for the microstructures and the fabrication of the<br />
cover.<br />
This 3 rd step is divided in two parts: the first one is the<br />
preparation of the cover which is called in this document the<br />
interface layer (IL) and the second one is the lamination of<br />
this IL onto the first level of the chessboard structure. The<br />
figure 5 illustrates the different operations realized for this<br />
step.<br />
In a first time, the cover is built. First of all, a PET film<br />
coated with a layer of glue (a weakly adhesive one) is<br />
laminated onto a thick glass wafer. Then the photo resist is<br />
spun-coated: it will be the IL. The thickness of the IL is set<br />
to 7 µm. The photo resist is soft baked at 50°C during 25<br />
minutes and then a room temperature stabilization delay is<br />
applied (at least 20 minutes long) (Fig. 5-1). We will take<br />
advantage that the photo-resist is still sticky after the SB to<br />
perform an efficient lamination onto the first level<br />
chessboard. The plastic film with the IL will be removed<br />
from the substrate (Fig. 5-2) just before performing the<br />
lamination onto the first level chessboard for the sealing.<br />
In a second time, the surface of the microstructures of the<br />
first level is activated by oxygen plasma (P=400W, t=30s,<br />
V=400ml). A drop of liquid is dripped down on the<br />
structure at the location where the lamination begins. In that<br />
way the liquid is pushed in the structure all along the<br />
lamination. This technique is interesting when only one type<br />
of liquid has to be filled in the tanks. At the end of the<br />
lamination, a full wafer UV exposition is done, associated<br />
to a PEB (still we the plastic film) at a low temperature<br />
(50°C) in order to fully polymerize the IL. The UV<br />
exposition has to be made quickly after the lamination in<br />
order to limit the solvent diffusion in the non polymerized<br />
resist at this time. After the PEB and a room temperature<br />
stabilization time, the plastic film is removed. Stages of this<br />
process step are illustrated by Fig. 5.<br />
1- Fabrication of<br />
IL on PET film<br />
Substrat<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Resist (7µm)<br />
PET<br />
Paste<br />
4- Removal of the PET film<br />
Fig. 5. Schematic representation of the filling and sealing step.<br />
The chessboard structure used in this experiment has<br />
sides of 300µm. The flatness of the IL (top side) has been<br />
measured after the lamination with a optical profilometer.<br />
The experimental profile variation is less than 400 nm high<br />
on 300 µm long as illustrated in figure 6.<br />
b)<br />
a)<br />
c)<br />
Fig. 6. Interferometric measurement over 1,2 x 0,9 mm area (a), horizontal<br />
profile line (b) and vertical one (c) of the cover of the level one after filling<br />
and sealing.<br />
In Fig. 6, it can be observed that the cover is shaped as a<br />
bump above tanks. A bump down was expected due to<br />
pressure applied during the lamination. However the<br />
amplitude of the deformation is low (350nm) and is not<br />
visible in the cross sectional view of figure 7<br />
2- Removal of PET<br />
film with the IL<br />
3- Lamination of the<br />
PET film onto the<br />
chessboard structure of<br />
the first level<br />
Fig. 7. SEM Cross-section view of tank with pillar and the cover (left) optical<br />
top view of filled and sealed strucrure (right).<br />
Fig. 7 shows also that the resist flew in the tank. It is<br />
visible on the right side of the pillar. There, the cover<br />
thickness is 9.37 µm (SEM measurement). Anywhere else,<br />
it is close to the expected value i.e 7 µm (physically<br />
measured with a mechanical profilometer). This is<br />
confirmed by the measurement above the pillar (6.94 µm)<br />
251
and far away from the pillar (7.47µm at 40µm from the left<br />
side pillar).<br />
The flow of the resist along the pillar during lamination is<br />
the major difficulty we have encountered during the<br />
development of this process. In order to avoid this problem,<br />
we have optimized process parameters. Figure 8 shows an<br />
example of flowing effect. This problem becomes really<br />
significant when microstructures have low dimensions as it<br />
reduces drastically the volume of encapsulated fluid. In our<br />
case it has been considered negligible regarding the final<br />
volume of each structure (1800µl). However, this<br />
phenomenon has been reduced by increasing the soft bake<br />
time (near twice the initial value) and by lowering the<br />
pressure and speed lamination to the minimum value<br />
allowed by the laminator. It has been tried to increase the<br />
thickness of the plastic layer which supports the resist but<br />
no improvement has been observed.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
2 – Liquid filling and sealing 3 – Removal of the top PET film<br />
Fig. 9. Realization of the second level.<br />
The figure 10 illustrates a SEM view of a quite complete<br />
two-level of microstructures showing tanks and pillars of<br />
the first level, IL covering the all first surface, and pillars of<br />
the second level.<br />
Pillar of the 2 nd level<br />
Upper tank level<br />
Interface layer<br />
lower tank level<br />
Pillar of the 1 st level<br />
Fig. 10. SEM view of empty structures without the final cover.<br />
Fig. 8. Example of a resist flow after the lamination (in the circle) before<br />
the optimization of the process.<br />
The first level of microstructures has been filled and<br />
sealed so the fabrication of the second level of<br />
microstructure can begin.<br />
D. Fourth step: fabrication of the second level and its<br />
sealing.<br />
In order to realize the second level of microstructures we<br />
have to repeat the same operations as described before (Fig.<br />
9). The first sealed level is treated with an oxygen plasma<br />
(P=200W, t=60s, V=200ml) in order to increase the<br />
wettability. The water drop angle is equal to 70° before<br />
plasma O 2 and falls down to 8° after treatment. The plasma<br />
O 2 increases also the adhesion of the spin coated resist on<br />
the IL surface of the first level. Then, the photo resist is<br />
spin-coated on the IL and soft backed as already before.<br />
Then using an alignment process on the Canon stepper, the<br />
second level is UV exposed. After a post exposure bake the<br />
second level is revealed. All parameters of these operations<br />
are the same those used in the fabrication of the first level.<br />
The second level has to be processed in a short delay (less<br />
than 2 hours) in order to avoid stress issues inducing a<br />
deformation of structures. The curvatures over the tanks can<br />
be increased from 300 nm to six microns if the time<br />
between the fabrication of level 1 and 2 is greater than 12<br />
hours.<br />
1- UV exposure and SB<br />
III. CONCLUSION<br />
We have developed a fabrication process to realize multilevel<br />
structures with encapsulated liquid. A chessboard<br />
structures have been chosen for demonstration but more<br />
complex structures could be realized as each level could be<br />
patterned. We stacked up two levels of reservoirs, but there<br />
is no limitation to go further and stack up other layers<br />
following the same process. The total area of the structure is<br />
118 cm² but it could be extended to larger area. We<br />
observed no formation of bubbles or liquid accumulation<br />
over pillars. This technology offers the possibility to easily<br />
implement optical functions at low cost and on large area<br />
and is therefore a basic step towards optofluidics<br />
applications.<br />
REFERENCES<br />
[1] D. Kohlheyer, J.Eijkel, S. Lenk, A. Floris, S. Staal, A. van den<br />
Berg, “Point-of-care lithium monitoring in whole blood using a<br />
disposable, prefilled and ready-to-use capillary electrophoresis<br />
microchip”, Micro Total Analysis Systems (µTAS), Jeju (Korea)<br />
2009, pp. 1731-1733.<br />
[2] E. Meng, C. Guttierrez,”Parylene-based encapsulated fluid MEMS<br />
sensors” 31th Int. Conf. IEEE EMBS, Minneapolis (USA), 2009,<br />
pp. 1039-1041.<br />
[3] T. Ninomiyaa, T. Okayamaa, Y. Matsumotoa, X. Arouettea, K.<br />
Osawaa, N. Miki, “MEMS-based hydraulic displacement<br />
amplification mechanism with completely encapsulated liquid”.<br />
Sensors & Actuators A, in press.<br />
[4] D. Psaltis, S.R. Quake, C.H. Yang, “Developeng optofluidic<br />
technology through the fusion of microfluidics and optics”, Nature,<br />
vol. 442, 2006, pp. 381-386.<br />
[5] P. abgrall, V. Conédéra, H. Camon, A.M. Gué, N.T. Nguyen, “SU-<br />
8 as a structural material for labs-On-Chips and MEMS”,<br />
Electrophoresis, Vol. 28, 2007, pp. 4539-4551.<br />
252
11-13<br />
<br />
May 2011, Aix-en-Provence, France<br />
<br />
Multiple-output MEMS DC/DC converter<br />
A. Chaehoi, M. Begbie, D. Weiland, S. Scherner.<br />
Institute for System Level Integration, Heriot Watt University Research Park, Research Avenue North, EH14 4AP Edinburgh, Scotland,<br />
www.isli.co.uk<br />
Contact: Aboubacar Chaehoi, tel: +44 (0)131 510 0681, fax: +44 (0)131 449 3141, aboubacar.chaehoi@sli-institute.ac.uk<br />
Abstract<br />
DC/DC converters are widely used in consumer<br />
electronic devices where usually a single power source is<br />
available while the electronic board of the device requires<br />
different voltage levels in order to power-up different block<br />
functions. In this paper we present the design of a MEMS<br />
single-input multi-output voltage level shifter. The lowvoltage<br />
to high-voltage conversion is based on the<br />
electrostatic transduction of variable capacitors built using<br />
interdigitated comb fingers. A 1mm2 MEMS prototype has<br />
been designed and fabricated using the SOIMUMPs process.<br />
In this study we present the co-design and co-simulation of<br />
the whole system (the MEMS device and its dedicated<br />
charge-pump-circuit) in a single <strong>EDA</strong> environment through<br />
MEMS+ [a Coventorware® tool that allows the cosimulation<br />
of MEMS and electronics in the Cadence Analog<br />
Design Environment]. We present analytical, FEM and<br />
MEMS+ models of the multi-output DC-DC converter and<br />
show that all our models converge towards the experimental<br />
results.<br />
Key words: MEMS, co-design, co-simulation, multi-output<br />
DC/DC converter.<br />
I. Introduction<br />
Electronic devices usually require multi level voltage supplies<br />
which must be derived from a single unique power source. For<br />
instance common handheld products are powered using one<br />
battery cell and at the same time different voltages level are<br />
required for different functions ranging from the microprocessor<br />
(a few volts) to the different the ASICs and memories to the<br />
screen display (up to 40 volts). Two types of purely electronic<br />
converter are dominant in the market: inductor-based DC/DC<br />
converter where a bulky inductor is needed and switch-mode<br />
DC/DC converters which suffer from switching losses and<br />
switching noise. Applying these architectures to multi-output<br />
systems inevitably leads to a large size system with the increase<br />
of inductor and/or capacitor number. We propose in this study a<br />
single-structure MEMS device that converts a single input DC<br />
voltage into two different DC output voltages. The principal of<br />
this MEMS device can easily be extended into multiple (more<br />
than two) output DC/DC converter. Voltage conversion is based<br />
on the electrostatic transduction of variable capacitors built<br />
using interdigitated comb fingers. The multiple-outputs are<br />
achieved by incorporating comb structures with different finger<br />
gap spacings into the same single structure.<br />
In a previous publication [1] the authors presented the design of<br />
a 1mm2 MEMS prototype designed and fabricated using the<br />
SOIMUMPs process [2]. The prototype exhibits 6.8V and 9V<br />
outputs from a supply driving voltage of 5V, with an initial rise<br />
time of 50ms to reach full output voltages. The development of<br />
the DC-DC converter was performed by separate simulations of<br />
the MEMS and electronics. In this study we propose and<br />
demonstrate a new design approach to the DC/DC converter.<br />
Designers usually develop their own models (VHDL, AHDL)<br />
for co-simulation of MEMS with CMOS or more commonly<br />
design and simulate MEMS and electronics separately, manually<br />
passing results from one simulation domain to another. This can<br />
lead to a number of problems including: incompatibility of<br />
interfaces, non-standard operation conditions and dynamic<br />
interaction between MEMS and electronics which cannot<br />
properly be simulated. Moreover this approach typically does<br />
not allow the IC designer any flexibility in the behaviour of the<br />
MEMS device. As an example the resonant frequency of<br />
vibrating structures is fixed by the mechanical design performed<br />
by the MEMS designer.<br />
In this new study we are able to co-simulate the whole system in<br />
both domains through MEMS+, a Coventorware® tool that<br />
allows the co-simulation of MEMS and electronics in the<br />
Cadence Analog Design Environment [2]. We have thus been<br />
able within the electronic design spaceto modify the mechanical<br />
structure and behavior within limits defined in the MEMS<br />
design space. The benefit we realize is the optimization of both<br />
the MEMS and the IC at the same time. The accurate MEMS<br />
behavioral model generated with MEMS+ is imported in<br />
Cadence Virtuoso as a schematic bloc in which parameters such<br />
as geometry and environmental variables can be changed, thus<br />
allowing the co-simulation of the MEMS device and its<br />
dedicated charge-pump-circuit in a single <strong>EDA</strong> environment. As<br />
part of this work both elements have been optimized based on<br />
their interaction (gap spacing, driving frequency) and the design<br />
space is explored in more detail than previously possible. We<br />
present analytical, FEM and MEMS+ models of the DC-DC<br />
converter and show that all our models converge towards the<br />
experimental results. We also analyse the influence of the<br />
resonance frequency of the mechanical device on the whole<br />
DC/DC converter efficiency.<br />
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11-13 May 2011, Aix-en-Provence, France<br />
<br />
II. The MEMS DC level converter principle <br />
The operating principle of the mechanical DC/DC converter is<br />
depicted on Figure 1. The MEMS is composed of a pair of<br />
interdigitated capacitive comb fingers used to drive laterally the<br />
whole structure at its resonant frequency, then on each side of<br />
the movable plate are attached further pairs of comb fingers<br />
forming two variable pump capacitors. Voltage conversion is<br />
based on the electrostatic transduction of the variable pump<br />
capacitors. The multiple-output is achieved by designing comb<br />
finger pairs with different gap spacings located in the same<br />
structure. The driving capacitors (Cd 1 and Cd 2 ) are connected<br />
such a way that with a driving sinusoidal voltage they alternately<br />
move the central plate along the lateral axis. An input voltage is<br />
applied across the pump capacitors (C 1 and C 2 ) which sets the<br />
initial charge Q 0 at their terminals. A decrease of the charge<br />
pump capacitor gaps leads to an increase of the capacitance<br />
value at a constant charging voltage which means that the charge<br />
in the capacitors increases. Then during the increase of the<br />
spacing gap, the accumulated charges are confined and<br />
transferred to an output storage capacitor at higher voltage in<br />
line with the reducing capacitance. Low leakage current diodes<br />
are used to prevent the charge cycling back to the supply and to<br />
the charge pump capacitors during the successive cycles thus<br />
ensuring the condition of constant charge.<br />
Where E is the young modulus of the beam, w b , is the width, t b<br />
is the and L a and L b are respectively the thigh length and the shin<br />
length of the suspension beam.<br />
The total mechanical damping D of the system is the sum of the<br />
squeeze film damping on the pump-charge capacitors (D 1 +D 2 ),<br />
the slide film damping on the driving comb-drive capacitors<br />
(D 3 ) and the slide film damping between the plate and the<br />
substrate (D 4 ).<br />
The squeeze film damping between the sensing comb fingers is<br />
calculated as:<br />
<br />
<br />
<br />
; <br />
<br />
(3)<br />
The slide film damping between the driving comb fingers is<br />
expressed as:<br />
<br />
2 <br />
(4)<br />
<br />
The slide film damping between the driving plate and the<br />
substrate is:<br />
<br />
(5)<br />
<br />
The damping ratio of the system is:<br />
<br />
<br />
(6)<br />
<br />
The quality factor of the system is obtained as:<br />
<br />
<br />
<br />
(7)<br />
From the electrical force on the comb drive, the stiffness of the<br />
beams and the quality factor, the static displacement of the<br />
movable plate and the displacement at the resonance can be<br />
calculated:<br />
<br />
<br />
(8)<br />
. (9)<br />
Figure 1: SEM picture of the dual-output DC/DC converter<br />
III. Analytical modeling<br />
The displacement of the central plate is due to the<br />
electrostatic force F elect generated by the voltage applied the<br />
driving capacitor Cd. The electrostatic force is calculated using<br />
the following formula [3]:<br />
<br />
<br />
(1)<br />
<br />
Where n d is the number of driving pair fingers, β is the fringe<br />
effect factor, ε r is the dielectric constant of the air, V d is the<br />
applied voltage and d RSd is the rotor to stator finger spacing gap.<br />
The movable plate is connected to four L-shaped beams. This<br />
kind of crab-leg beam suspension has an overall spring constant<br />
expressed as follows [4]:<br />
<br />
<br />
(2)<br />
From the equations developed above and the structure geometry<br />
describes in Table 2, the mechanical parameters and behavior of<br />
the MEMS device can be calculated. The table below<br />
summarizes the results of the analytical modeling.<br />
Variable Description Value<br />
Mechanical parameter<br />
F elect Electrostatic force 4.8E-8 N<br />
K Stifness 29.5 N/m<br />
M Mass of driving plate (kg) 1.0715E-8<br />
D Total mechanical damping (N.s.m -1 ) 5.2E-7<br />
ξ Damping ratio 4.61E-4<br />
Q Quality factor 1.0835E3<br />
F res Resonant frequency 8.35 kHz<br />
x static Static displacement of driving plate 1.63 nm<br />
x peak Displacement at resonance 1.761 μm<br />
Table 1: Analytical modeling – result summary.<br />
254
Variable Description Value<br />
Crab-leg suspension beam of driving plate<br />
E Material Young modulus 169 GPa<br />
w b Beam width 10 µm<br />
t b Beam thickness 10 μm<br />
L b Beam length 600 μm<br />
L a Beam length (thigh) 50 μm<br />
d p distance between plate and substrate 400 μm<br />
A Area of driving plate 2.5E-07<br />
M Mass of driving plate (kg) 1.0715e-8<br />
Driving capacitors Cd 1 Cd 2<br />
w Comb finger width 5 μm<br />
t Comb finger thickness 10 μm<br />
L d Comb finger length 70 μm<br />
h d Overlapping height of electrodes 40 μm<br />
n d Number of finger pairs 50<br />
d RSd Rotor to stator finger spacing gap 3 μm<br />
β Correction factor for the fringe effect 1.2 – 1.8<br />
ε r Dielectric constant of air 1<br />
ε 0 Permittivity of vacuum (F/m) 8.85e-12<br />
V d Applied voltage at the electrodes 5 V<br />
Pump charge capacitors C 1 , C 2<br />
w Comb finger width 5 μm<br />
t Comb finger thickness 10 μm<br />
L Comb finger length 80 μm<br />
h Overlapping height of electrodes 70 μm<br />
N Number of finger pairs 20<br />
d SS Stator to stator finger spacing gap 25 μm<br />
d RS1 C 1 rotor to stator finger spacing gap 8 μm<br />
d RS2 C 2 rotor to stator finger spacing gap 5 μm<br />
Table 2: MEMS device parameters.<br />
At resonance, the movable fingers oscillate around their initial<br />
positions with an amplitude of x peak . In the folowing we assume<br />
x min and x max being respectively the minimum and the maximum<br />
position of the fingers.<br />
(10)<br />
(11)<br />
Where d RS is the initial rotor to stator distance.<br />
For the present device – described in Figure 2 –, the left and<br />
right capacitance, C l and C r respectively, at x min and the total<br />
maximum capacitance C max (the sum of C l and C r ) can be<br />
calculated as follows [6]:<br />
.<br />
_ . <br />
(12)<br />
<br />
_ . <br />
.<br />
(13)<br />
<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
. . <br />
<br />
<br />
<br />
<br />
<br />
(14)<br />
In the same way the total minimum capacitance C min when the<br />
comb-finger is at x max can be expressed as:<br />
. . <br />
<br />
<br />
<br />
<br />
<br />
(15)<br />
The output voltage of the DC/DC converter can then be<br />
calculated as [7]:<br />
<br />
<br />
(16)<br />
We then find output voltages of 5.8V and 8.2V for the pump<br />
charge capacitors C 1 and C 2 respectively.<br />
C l<br />
Rotor<br />
C r<br />
d RS<br />
Stator<br />
a)<br />
h<br />
w<br />
L<br />
Figure 2: Interdigitated comb finger<br />
before and after x displacement.<br />
IV. MEMS+ Modelling<br />
Rotor<br />
Stator<br />
Figure 3 is a schematic of the whole system in Virtuoso. The<br />
output of the MEMS converter is connected to a 500 fF reservoir<br />
capacitor via a low leakage current diode to provide rectification<br />
with minimal loss during the pumping cycles. The device is<br />
driven at its resonant frequency.<br />
This model is used to study the influence of the mechanical and<br />
the electrical parameters of each block and their mutual<br />
interdependence on the performance of the whole system. Figure<br />
4 shows the influence of load resistance on the performance of<br />
the converter. The better performance is obtained for a very high<br />
resistive load (1TΩ) where an output voltage around 6.5V can<br />
be obtained at the first output C 1 of the DC/DC converter.<br />
However the output voltage drops to nil when the load resistivity<br />
is under a few tens of giga-ohms. This MEMS-electronics<br />
cosimulation allows us to understand that the system is suitable<br />
for purely capacitive loads only. This is due to the<br />
acknowledged inherent high impedance output of MEMS<br />
converters [8]. The modeling prediction of the converter<br />
behavior for low resistive loads is interesting and need some<br />
more study.<br />
x<br />
d SS<br />
b)<br />
255
MEMS DC/DC converter<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
Driving plate displacement<br />
ΔC<br />
Figure 3: Schematic view of the whole DC/DC converter system in Virtuoso.<br />
R load = 1 TΩ<br />
R load = 100 GΩ<br />
R load = 10 GΩ<br />
R load = 1 GΩ<br />
Figure 4: Transient response of the DC/DC converter for<br />
different resistive loads.<br />
This modeling approach also allows a geometrical analysis of<br />
the mechanical part regarding the output voltage of the whole<br />
system. Figure 5 shows the transient response of the converter<br />
for different rotor to stator distance d RS on the pump capacitor.<br />
The stationary output voltage increases when d RS decreases,<br />
which is explained by a higher maximum voltage multiplication<br />
factor (C max /C min ) when d RS is low. However a drastic increase in<br />
charging rate appears for very low rotor to stator distances. This<br />
is due to a the electrostatic spring which appears because of the<br />
electrostatic force generated at the capacitor plate when a<br />
voltage is present (pull-in effect) [9,10]. This added stiffness<br />
shifts the resonant frequency of the system during the charging<br />
cycles which increases the driving plate amplitude displacement.<br />
Figure 6 shows the FFT of the driving plate displacement for<br />
different values of d RS . We can see that for a small value of d RS<br />
another oscillating mode is superimposed on the driving mode<br />
frequency. This new oscillation mode corresponds to the new<br />
resonant frequency of the mechanical structure (around 8.1 kHz)<br />
C1_d RS = 8 µm<br />
C1_d RS = 5 µm<br />
C1_d RS = 5.5 µm<br />
C1_d RS = 6 µm<br />
C1_d RS = 5 µm<br />
C1_d RS = 8 µm<br />
Figure 5: Transient response of the DC/DC converter for<br />
different finger spacings.<br />
Figure 6: FFT of the x displacement signal of the driving plate<br />
for different rotor to stator distance d RS .<br />
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V. Characterization results and Conclusions <br />
References<br />
The fabricated device exhibits two output voltages of 9V<br />
and 6.8V, after an initial rise time of 50ms to reach its full<br />
output voltages when excited at its resonant frequency<br />
characterized at 8.32 kHz. There are some discrepancies<br />
between the model results and the characterization. In order to<br />
achieve output voltages of 6.8V and 9V, the distance d RS<br />
between the fingers for C1 and C2 needs to be 7µm and 6µm<br />
respectively. This can be explained, for the most part, by the fact<br />
that the high quality factor Q and the resonant frequency of the<br />
real device F res are slightly different from the ones from the<br />
model (process variation), and we know that the performance is<br />
highly dependent of the driving frequency through the effect this<br />
has on the amplitude. Therefore the “pull-in effect” modeled for<br />
the small rotor to stator spacing gap is not seen yet during<br />
characterization.<br />
Resonant Freq<br />
F res<br />
Output<br />
V out<br />
Analytical model 8.35 kHz 5.8 / 8.2 V<br />
FEM 8.24 kHz NA<br />
MEMS+ 8.26 kHz 5.8 / “pull-in effect”<br />
Characterisation 8.32 kHz 6.8 / 9V<br />
V. Conclusion<br />
Table 3: Results summary.<br />
This paper concerns the modeling and design of a<br />
MEMS single-input, multiple-output DC/DC converter. The<br />
proposed model is in good agreement with the characterization<br />
results. It can be used very rapidly to study the effects of<br />
geometrical and electrical parameters on the whole system<br />
performance. Because of its high output impedance, the present<br />
system is suitable only for purely capacitive loads. More works<br />
need to be done in designing pump charge capacitors with<br />
bigger C 0 values in order to achieve a generic DC-DC converter.<br />
The driving frequency of the MEMS controls the output voltage<br />
of the DC/DC converter. An adaptative and dynamic voltage<br />
scaling can therefore be considered with such MEMS apparatus.<br />
[1] L. Li et al.; “Single-input, dual-output MEMS DC/DC<br />
converter”; Electronics Letters, Vol. 43 No. 15; Jul 2007.<br />
[2] www.coventor.com/mems-ic/mems-product-designplatform.html<br />
[3] www.coventor.com/mems-ic/mems-product-designplatform.html<br />
[4] M.H. Bao; “Micromechanical transducers: Pressure<br />
sensors, accelerometers and gyroscopes”; Elsevier<br />
Sciences, 30 th October 2000; ISBN: 978-0444505583.<br />
[5] G.K. Fedder; “Simulation of Microelectromechanical<br />
Syatems”; Ph.D. dissertation; University of California,<br />
Berkley; 1994.<br />
[6] K. Sharma, “Design optimization of MEMS comb<br />
accelerometer”; American society for Engineering<br />
Education Zone 1 Conference, West Point, NY, 28 th -29 th<br />
March 2008<br />
[7] M. Hill et al.; “Modeling and performance evaluation of a<br />
MEMS DC-DC converter”; J. Micromech. Microeng. 2006,<br />
16, pp. S149-S155<br />
[8] C.H. Haas et al.; “Modelling and analysis of a MEMS<br />
approach to DC voltage step up conversion”; J.<br />
Micromech. Microeng. 2004.<br />
[9] D. Galayko et al.; “Coupled resonators micromechanical<br />
filters with voltage tunable bandpass characteristic in<br />
thick-film polysilicon technology”; Sensors and Actuators;<br />
Vol.126, 2006.<br />
[10] E.H. Francis et al.; “Electrostatic spring effect on the<br />
dynamic performance of micro resonators”; International<br />
Conference on Modeling and Simulation of Microsystems,<br />
San Diego, 27-26 March 2000, pp. 154-157<br />
257
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<br />
Design of the silicon membrane of high fidelity and<br />
high efficiency MEMS microspeaker<br />
Iman Shahosseini, Elie Lefeuvre, Emile Martincic,<br />
Marion Woytasik, Johan Moulin, Souhil Megherbi<br />
Univ. Paris Sud – CNRS<br />
Institut d'Electronique Fondamentale<br />
91405 Orsay Cedex, France<br />
Abstract- This study presents a novel approach to MEMS<br />
microspeakers design aiming to tackle two main drawbacks of<br />
conventional microspeakers: their poor sound quality and<br />
their weak efficiency. For this purpose, an acoustic emissive<br />
surface based on a very light but very stiff structured silicon<br />
membrane was designed. This architecture, for which the<br />
membrane undesirable vibration modes were reduced to only<br />
five within the microspeaker bandwidth, is promising to let the<br />
microspeaker produce high sound quality from 300 Hz to 20<br />
kHz. This silicon membrane is suspended by a whole set of<br />
silicon springs designed to enable out-of-plane displacements<br />
as large as 300 µm. Different geometries of springs were<br />
considered and the material maximum stress was analyzed in<br />
each case by finite element modeling. The proposed structure<br />
promises an efficiency of 10 -4 , that is to say ten times higher<br />
than that of conventional microspeakers.<br />
I. INTRODUCTION<br />
The broad development of mobile electronic devices<br />
embedding audio function is now strongly increasing the<br />
demand for higher sound level and better sound quality.<br />
From this point of view, the problem mainly comes from<br />
the poor quality of available microspeakers. Thus, more and<br />
more attention is being paid to acoustic performances of the<br />
microspeakers used for instance in mobile phones, which<br />
represent more than one billion units per year market. This<br />
explains why significant research efforts are currently<br />
focused on improvement of the performances of<br />
microspeakers [1, 2]. Until now, such microspeakers are not<br />
MEMS: they are manufactured using conventional<br />
"macroscopic" machining technologies. But limits of<br />
conventional technologies in terms of integration and sound<br />
quality are not far to be reached, and MEMS technologies<br />
present a very promising potential for overcoming these<br />
limitations, as shown by recent studies [3]. Indeed,<br />
microtechnologies bring outstanding dimensional precision<br />
and good reproducibility which are needed for<br />
manufacturing high quality sound transducers. Moreover,<br />
thanks to batch process the fabrication costs may be kept<br />
reasonably low.<br />
Another critical issue of mobile electronic devices is the<br />
autonomy of batteries. Taking again the example of cell<br />
Romain Ravaud and Guy Lemarquand<br />
Université du Maine – CNRS<br />
Laboratoire d'Acoustique de l'Université du Maine<br />
72085 Le Mans, France<br />
phones, nearly one quarter of the total power consumption<br />
is due to the audio system when used in free-hand mode.<br />
Analysis of the components usually used in the sound<br />
reproduction chain of mobile devices shows that D-A<br />
converters have very little consumption. Amplifiers have<br />
pretty good efficiencies, typically between 50% and 90%.<br />
From the efficiency point of view, the weakness is mainly<br />
due, again, to the microspeakers. Indeed, efficiency of the<br />
electrical-to-acoustic power conversion remains typically<br />
lower than 0.001%. So it is clear that improvement of the<br />
efficiency of microspeakers is the best approach to increase<br />
significantly the overall efficiency of the audio chain. For<br />
instance, improvement of the microspeaker efficiency by a<br />
factor of ten, that is to say reaching 0.01% efficiency, will<br />
roughly divide the consumption of the sound reproduction<br />
chain by the same factor ten. The total power consumption<br />
will thus be notably reduced, with significant gain in term<br />
of energy autonomy of mobile devices.<br />
The approach developed in this paper aims at improving<br />
both the efficiency and the sound quality of microspeakers.<br />
Until now, few works on MEMS microspeakers have been<br />
reported in literature. Transduction principles such as<br />
piezoelectric, electrostatic, electrostrictive, electrodynamic<br />
and thermoacoustic actuation, which are achievable using<br />
MEMS technologies, have been proposed [4]. But nonlinear<br />
response of piezoelectric, electrostrictive and<br />
thermoacoustic materials is a major drawback for high<br />
fidelity transduction. Electrostatic principle, which is<br />
broadly used for MEMS actuators because of its<br />
technological simplicity, has however low power density<br />
and requires relatively high driving voltages. So, although it<br />
requires magnets whose integration into MEMS is not very<br />
developed yet, electrodynamic actuation principle is the best<br />
way to meet the objectives in terms of linearity, power<br />
density and efficiency. This actuation principle lies on the<br />
Lorentz force which appears on a conductor, usually coil<br />
shaped, driven by an electrical current and surrounded by a<br />
magnetic field, usually created by a permanent magnet.<br />
Predictions developed in this paper, based on analytical and<br />
finite element method (FEM) modeling of the microspeaker<br />
show that the targeted efficiency of 0.01% is reachable<br />
using a planar copper coil and ring-shaped magnets with<br />
axial magnetization of 1.5 T.<br />
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11-13 <br />
May 2011, Aix-en-Provence, France<br />
Compared to the MEMS microspeakers formerly<br />
<br />
presented in literature which are mainly based on<br />
deformable membranes, the originality of the proposed<br />
structures lies on the use of quasi-undeformable silicon<br />
membrane suspended to the substrate by a whole set of Suspension beam<br />
silicon springs. Using such perfectly flat and rigid emissive<br />
surface is actually ideal for the sound quality. On the other<br />
Via<br />
hand, high efficiency requires very light membrane. A<br />
tradeoff between these two constraints was found using<br />
Insulator<br />
FEM modeling of a silicon membrane with stiffening ribs.<br />
This paper first presents the whole structure of the Conductor track<br />
MEMS microspeaker and gives the basic design relations<br />
between sound pressure level, efficiency, surface and<br />
displacement of the membrane, mobile masses of the coil<br />
and the membrane. Then, the FEM modeling of a light and<br />
rigid microstructured membrane is detailed in section III.<br />
Microcoil<br />
Afterwards, section IV highlights the importance of the<br />
springs design which was worked out to decrease as much<br />
as possible the stress concentration zones. Finally, section V<br />
outlines the perspectives of this work.<br />
Magnet<br />
Membrane<br />
II.<br />
MEMS STRUCTURE<br />
A loudspeaker is an electroacoustic transducer in which<br />
the sound is produced in response to an electrical signal.<br />
Basically, such kind of transducer generates acoustic wave<br />
by dynamic displacement of a volume of air. For this<br />
purpose, in classical loudspeakers the displacement of a<br />
solid, conic or dome-shaped diaphragm plays this role. In<br />
the case of the MEMS microspeakers presented in most of<br />
former works, the acoustic wave is produced using a<br />
deformable diaphragm, made in a materials such as<br />
parylene or polyimide [3],[5]. Usually, the diaphragm is<br />
clamped to the substrate on its peripheral edge, and<br />
consequently, the principal vibration mode of such<br />
microspeaker, besides many other uncontrolled vibration<br />
modes, is the drum mode. In this work, the sound wave is<br />
generated by a rigid membrane, resisting unwanted<br />
deformations and held by low stiffness suspension beams.<br />
The desired vibration mode is the piston mode, which<br />
means that the whole surface of the membrane runs always<br />
in parallel with its original position. Such operation<br />
principle is ideal from the acoustics point of view. It is<br />
actually a key factor for high quality sound reproduction.<br />
Fig. 1 represents a schematic view of the MEMS<br />
microspeaker structure.<br />
Here, the electrodynamics actuation is responsible for<br />
generation of the driving force, which produces the desired<br />
displacement of the membrane. A magnetic field and an<br />
electric current passing through a conductor are two<br />
essential elements to create the driving force, known as the<br />
Lorentz force. For the proposed structure, the magnetic field<br />
is created by a ring-shaped magnet surrounding the circular<br />
membrane. The conductor is a planar microcoil wound on<br />
top of the membrane. This microcoil is placed as close as<br />
possible to the magnet in order to use the maximum<br />
intensity of the magnetic field. Its electrical supply is<br />
achieved using two conductive tracks which are supported<br />
Fig. 1. Top view and cross-section of the schematic structure of the<br />
MEMS microspeaker<br />
by the suspension beams. Moreover, the coil is protected<br />
from short-circuits by an electrically insulating layer.<br />
Connections of the tracks to the coil ends are achieved<br />
through two via across the insulating layer.<br />
The first step to design a loudspeaker is to determine the<br />
size and the displacement of the emissive surface (the<br />
membrane). These parameters are linked to the sound<br />
pressure level (SPL) and the frequency bandwidth. As the<br />
new standard for high fidelity mobile audio systems stands,<br />
the target has been set to 80 dB SPL at 10 cm, within a<br />
bandwidth of 300 Hz to 20 kHz. With the help of equation<br />
(1), the acoustic power P acoustic that the microspeaker should<br />
produce is of 12.6 µW.<br />
L dB<br />
10 −12<br />
4 2<br />
acoustic<br />
10 10 π ×××=<br />
a (1)<br />
P<br />
In this equation, L dB is the SPL at the distance a.<br />
According to Eq. (2), for a loudspeaker with a circular<br />
membrane which moves in piston mode, the acoustic power<br />
is proportional to the displaced volume of air at a given<br />
frequency f. Both the membrane diameter d and the<br />
membrane out-plane displacement x peak determine the air<br />
volume moved by the membrane.<br />
244<br />
acoustic = 2 70 x peak<br />
f<br />
(2)<br />
It can be deduced that in order to compute the membrane<br />
maximum displacement, the minimum of the frequency<br />
band should be considered. Therefore, for a membrane<br />
whose diameter was fixed to 15 mm, the larger peak<br />
displacement x peak is of 300 µm. This value will be<br />
considered in section IV for the suspension springs design.<br />
In other words, the membrane must remain flat with no<br />
deformation, and it is all the task of the flexible suspensions<br />
to provide 600 µm stroke. As mentioned, acoustically the<br />
piston mode is the ideal vibration for the whole bandwidth.<br />
259
11-13 <br />
May 2011, Aix-en-Provence, France<br />
This means that the piston mode should take place before<br />
<br />
reaching 300 Hz frequency at which the membrane runs<br />
Central circle<br />
high displacements. While increasing the frequency, though<br />
the membrane vibration remains piston mode, the<br />
displacement amplitude reduces enormously. As for the<br />
emissive surface, it is indispensable to have a rigid and<br />
undeformable suspended membrane. However, its lightness<br />
is also an important factor as it plays a role in the<br />
microspeaker efficiency η. This point is highlighted by Eq.<br />
(3), which shows that the lighter it is, the higher the<br />
efficiency can be.<br />
=<br />
4<br />
.. rπρ 1 ⎛ f ⎞<br />
Force<br />
. .<br />
4<br />
⎜<br />
⎟<br />
Rc<br />
⎝ coil<br />
+ MM<br />
membrane ⎠<br />
η (3)<br />
In this equation, ρ is the air density (1.2 kg/m 3 at 20°C), r<br />
the membrane radius, c the sound speed (343 m/s at 20°C),<br />
R the coil resistance, M coil and M membrane the weight of the<br />
coil and that of the membrane. The force factor f Force which<br />
is determined as a result of the driving force per current unit<br />
meets 0.35 N/A. This value was attained through<br />
electromagnetic optimization of the coil and the magnet [6].<br />
III.<br />
MEMBRANE DESIGN<br />
The dynamic performances were first analyzed on a thin<br />
silicon disc structure using FEM simulations. Silicon was<br />
chosen deliberately because it fulfills both rigid and light<br />
criteria. Its Young modulus to density ratio of 71<br />
GPa.gr/cm 3 is actually three times higher than that of other<br />
common materials used in MEMS technology such as<br />
titanium or aluminum.<br />
The modal results showed that for a 20 µm thick disc,<br />
more than 40 different vibration modes exist in the<br />
microspeaker bandwidth. High sound reproduction quality<br />
asks for as little vibration modes as possible. Thickening the<br />
membrane can be considered as a solution for shifting most<br />
of the modes to frequencies higher than 20 kHz. For<br />
instance, FEM modal simulations of a 320 µm thick disc<br />
showed only two undesirable vibration modes, with the<br />
drum mode at 20 kHz. Unfortunately, such solution strongly<br />
increases the membrane weight, which reduces significantly<br />
the loudspeaker's efficiency. Indeed, the 320 µm thick<br />
membrane weights 132 mg, that is to say 16 times more<br />
than the 20 µm one. According to Eq. (3), the efficiency<br />
would be divided by a factor of 93 if considering an<br />
optimized coil of 6 mg.<br />
Several microstructures of the membrane were considered<br />
to prevent efficiency deterioration while keeping most of<br />
the vibration modes out of the frequency bandwidth. The<br />
idea was to dig up some areas in the membrane and to find a<br />
good trade-off between the membrane weight and its<br />
rigidity. Comparing different possible designs such as<br />
hexagonal shape or crossed beams, led us to conceive the<br />
ribbed structure shown in Fig. 2, which includes one 3 mm<br />
diameter central ring and one peripheral ring, each 200 µm<br />
wide, joined together by a series of radial ribs. In order to<br />
have results compatible with microfabrication process, the<br />
2<br />
Fig. 2. Structure of analyzed ribbed membrane for the microspeaker<br />
depth of the structured part was set to 300 µm. The<br />
thickness of the plain membrane was set to 20 µm. In fact,<br />
the micromachining process is based on a silicon-oninsulator<br />
(SOI) substrate for which the top side silicon layer<br />
and the substrate are respectively 20 µm and 300 µm thick.<br />
The effect of the number and the width of the radial ribs<br />
on the vibration modes were analyzed using FEM<br />
simulations. The results concerning the drum mode<br />
frequency are shown on Fig. 3 computed with a number of<br />
ribs between 10 and 40 and with four different widths of the<br />
ribs: 50 µm, 100 µm, 150 µm, and 200 µm. These<br />
simulation results show that the drum mode frequency is<br />
optimally shifted towards high frequencies for a ribs<br />
number between 14 and 15. The drum mode is the vibration<br />
mode which deteriorates mainly the sound quality. In<br />
particular, this vibration mode should not appear in the low<br />
and medium frequencies, but one can consider that its effect<br />
is not perceptible above 12 kHz.<br />
The membrane weight varies also with the number and<br />
the thickness of the radial ribs, as shown on Fig. 4. The<br />
maximum drum mode frequency and the corresponding<br />
membrane weight for each series of ribs thicknesses are<br />
summarized in Table I for each series of ribs width. The 50<br />
µm width seems theoretically promising to adopt, but<br />
microfabrication defects due to high aspect ratio may be a<br />
problem. Consequently, 100 or 150 µm width for the ribs<br />
leads to a good trade-off between the sound quality (related<br />
to the drum mode), the efficiency (related to the membrane<br />
weight) and the microfabrication yield (related to the aspect<br />
ratio of the ribs).<br />
Drum mode frequency (Hz)<br />
14000<br />
13500<br />
13000<br />
12500<br />
12000<br />
11500<br />
11000<br />
Rib<br />
Peripheral circle<br />
50 µm<br />
100 µm<br />
150 µm<br />
200 µm<br />
10 15 20 25 30 35 40<br />
Ribs number<br />
Fig. 3. Drum mode frequency of structured membrane as a function of<br />
ribs number, for four different ribs thicknesses, 50, 100, 150, and 200 µm<br />
260
Membrane weight (mg)<br />
Fig. 4. Structured membrane weight as a function of ribs number, for<br />
four different ribs thicknesses, 50, 100, 150, and 200 µm<br />
TABLE I<br />
Summary of maximum drum mode frequency for each ribs thickness<br />
Ribs width<br />
(µm)<br />
55<br />
45<br />
35<br />
25<br />
15<br />
50 µm<br />
100 µm<br />
150 µm<br />
200 µm<br />
10 15 20 25 30 35 40<br />
Ribs<br />
number<br />
Membrane weight<br />
(mg)<br />
50 15 18.1 12.3<br />
100 16 21.9 13<br />
150 14 24.2 13.4<br />
200 15 28.2 13.8<br />
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May 2011, Aix-en-Provence, France<br />
<br />
Drum mode frequency<br />
(kHz)<br />
Comparison of the two selected ribbed structures with a<br />
plain 20 µm thick silicon membrane shows that more than<br />
30 undesirables vibration modes have been eliminated from<br />
the microspeaker bandwidth and that the drum mode<br />
frequency takes place ten times higher. In addition, the only<br />
five vibration modes that the structured membranes show in<br />
the bandwidth are identical to the first five vibration modes<br />
of the 20 µm thick membrane, but with a difference that<br />
they do happen in much higher frequencies, as it is the case<br />
for the drum mode. Adding up the ribs increases the<br />
membrane weight and decreases the efficiency by a factor<br />
of four. Though 320 µm thick membrane practically<br />
eliminates the drum mode, its efficiency is at least 22 times<br />
lower than that of structured membranes. These results are<br />
summarized in Table II.<br />
IV.<br />
Ribs number<br />
SUSPENSION SPRINGS<br />
In order to obtain a piston movement, having a rigid<br />
membrane is not the only factor, but also possessing high<br />
flexible suspension beams which provide large out plane<br />
displacements for the membrane. Three different aspects<br />
must be considered for the suspensions springs design: good<br />
TABLE II<br />
Modes number, drum mode, and efficiency with different membranes<br />
Membrane<br />
Modes number in the Drum mode<br />
300 Hz - 20 kHz range (kHz)<br />
Efficiency *<br />
20 µm thick 40 1.3 5.3 10 -4<br />
320 µm thick 2 20 5.6 10 -6<br />
microstructured<br />
with16 ribs, 100<br />
5 13 1.4 10 -4<br />
µm width<br />
Microstructured<br />
with 14 ribs,<br />
5 13.8 1.2 10 -4<br />
150 µm width<br />
* with an optimized microcoil 6 mg in weight and 10 Ω in resistance<br />
mechanical linearity and piston mode frequency lower than<br />
300 Hz, which are required to ensure high sound<br />
reproduction quality, and also low stress concentration<br />
zones to ensure long lifetime of the device.<br />
FEM simulations enabled to determine the stress levels<br />
for 300 µm out-of-plane displacement for different shapes<br />
of the springs. It should be mentioned here that such an outof-plane<br />
displacement is very important compared to values<br />
usually found in MEMS literature. Indeed, some works have<br />
reported hundreds microns for in-plane displacements [7],<br />
but not many studies have been done for out-of-plane<br />
displacements of the same order of magnitude. Though<br />
solution studied in [8] has the potential of high<br />
displacements, because the suspensions are made of<br />
polymer they have linear response for small displacements<br />
only. For the first design we considered four suspension<br />
beam simply curved, clamped on one end into the substrate<br />
and on the other end into the membrane, as depicted on Fig.<br />
5. The curved shape was chosen for reducing the<br />
membrane-magnet distance, hence making intense magnetic<br />
field accessible to the coil. Due to the SOI-based fabrication<br />
process, the silicon suspension springs and the silicon<br />
membrane top layer have the same 20 µm thickness.<br />
Simulation results show that if the membrane is shifted to<br />
its peak displacement of 300 µm, 320 MPa maximal<br />
principal stress appears at the membrane-suspension<br />
anchorage zone. Despite this value is lower than theoretical<br />
elastic limit of silicon single crystal, which is often<br />
considered to be 1 GPa in literature, a safety factor of ten is<br />
chosen to raise the microspeaker resistance against<br />
unexpected operation or mechanical shocks.<br />
We experimented various shapes to lower the maximum<br />
stress, such as the "S" shaped and the stacked beam<br />
structures shown in Fig. 6-a and 6-b. Interestingly both<br />
geometries reduce the maximum principal stress to 60 MPa<br />
for the maximal out-of-plane displacement. Further analyses<br />
showed that for "S" type, the stress is a function of the beam<br />
length and the radius of the interior "U" turn. Increasing the<br />
beam width does not have an impact on the maximum value<br />
of the principal stress. However, wider the beams are,<br />
higher the beams stiffness will be. This interesting property<br />
enables to tune independently both parameters.<br />
Fig. 6-c shows that completely rounding the anchorage<br />
point in both ends helps the maximum principal stress to<br />
decrease down to 20 MPa only, which is more than 15 times<br />
lower than the initial structure shown in Fig.5.<br />
Clamped to<br />
substrate<br />
Suspension<br />
beam<br />
Membrane<br />
Fig. 5. Suspension beams clamped into the membrane and into the<br />
substrate, with a zoom box showing the anchorage maximum stress area<br />
(320 MPa)<br />
261
a) 60 MPa b) 60 MPa<br />
(c)<br />
d) 20 MPa e) 36 MPa<br />
Fig. 6. Principal stress distribution for different suspension beams<br />
designs, with the maximum principal stress value for a 300 µm<br />
displacement, "S" form (a), stacked beams (b), rounded anchorage point<br />
(c), membrane with 4-c type beams (d), membrane with 6-c type beams (e)<br />
By increasing the suspension beams from four (Fig.6-d)<br />
to six (Fig.6-e) to get better guided membrane, the<br />
maximum principal stress increases to 36 MPa because of<br />
the reduction of the springs length. The springs stiffness,<br />
which is the ratio of force to displacement, decreases from<br />
35 N/m for the first conceived structure to 1.3 N/m and 5.4<br />
N/m respectively for the 4-beam and 6-beam structures. It<br />
should be also highlighted that the stress values given on<br />
Fig. 6 correspond to the membrane maximum displacement.<br />
When working at frequencies higher than 300 Hz or SPL<br />
lower than 80 dB, less displacement is needed. As a result,<br />
the stress in anchorage points will be less critical.<br />
Modal analyses of the whole membrane and suspensions<br />
structure demonstrate that the piston mode frequency takes<br />
place at 35 Hz and 75 Hz respectively for 4 and 6<br />
suspension beams. Furthermore, contrary to the drum mode,<br />
the piston mode varies as function of both springs total<br />
stiffness, product of beams number and each beam stiffness,<br />
and the suspended mass, so the membrane weight.<br />
As stated before, the frequency of drum mode should be<br />
lower than 300 Hz. Nevertheless, in order to prevent the<br />
coupling phenomenon between the membrane vibration<br />
modes and the 50 Hz frequency radiated by many devices<br />
connected to the AC mains, the piston mode is better not to<br />
appear at lower frequencies than 50 Hz. Taking this factor<br />
into account, the design of Fig. 6-e is the most interesting.<br />
V. CONCLUSION AND PERSPECTIVES<br />
Today, conventional microspeakers are broadly used in<br />
different electronic mobile devices. However, their low<br />
efficiency and poor sound quality have stimulated this<br />
work. In the presented design based on MEMS technology,<br />
the emissive surface is a stiffened silicon membrane whose<br />
piston motions generate sound waves. The optimized<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
membrane in lightness and rigidity terms will allow<br />
obtaining better acoustic quality and higher efficiency than<br />
conventional microspeakers. This study has led to the<br />
design of suspension springs able to provide very important<br />
out plane displacements of 300 µm of the membrane. This<br />
essential step will enable to get the displacements desired<br />
for the membrane, while limiting the maximum principal<br />
stress value at 36 MPa. This is the guarantee of reliability<br />
and long lifetime for the microspeaker.<br />
Now that the structured membrane and its suspension<br />
have been designed, their micromachining is the next step<br />
of this work. In this way, a SOI substrate will be first<br />
patterned and etched in front side. After patterning and<br />
temporary bonding onto a silicon wafer, the back side will<br />
be structured using deep reactive ion etching (DRIE)<br />
process. Then, through an anisotropic etching of the buried<br />
silicon oxide layer, the membrane will be set free. Once the<br />
membrane is suspended, mechanical characterizations will<br />
be carried out to get the force-displacement characteristics<br />
and verify the vibration modes frequencies.<br />
ACKNOWLEDGMENT<br />
This work has been financially supported by the French<br />
Agence Nationale pour la Recherche (ANR).<br />
REFERENCES<br />
[1] W. Kim, G.-W. Jang and Y. Y. Kim, "Microspeaker diaphragm<br />
optimization for widening the operation frequency band and<br />
increasing sound pressure level", IEEE Trans. Mag., vol. 46, no. 1,<br />
January 2010, 59-66.<br />
[2] C.-M. Lee, J.-H. Kwon, K.-S. Kim, J.-H. Park, and S.-M. Hwang,<br />
"Design and analysis of microspeakers to improve sound<br />
characteristics in a low frequency range", IEEE Trans. Mag., vol.<br />
46, no. 6, June 2010, 2048-2051.<br />
[3] S.-S. Je, F. Rivas, R. E. Diaz, J. Kwon, J. Kim, B. Bakkaloglu, S.<br />
Kiaei, and J. Chae, “A compact and low-cost MEMS loudspeaker<br />
for digital hearing aids”, IEEE Trans. Biomed. Circ. and Syst., vol.<br />
3, n°5, 2009, 348-358.<br />
[4] I. Shahosseini, E. Lefeuvre, M. Woytasik, J. Moulin, X. Leroux, S.<br />
Edmond, et al. "Towards High Fidelity High Efficiency MEMS<br />
Microspeaker", 9 th annual IEEE Conference on Sensors 2010,<br />
November 1-4, Hawaii, 2010.<br />
[5] M. C. Cheng, W. S. Huang, and S. R. S. Huang, “A silicon<br />
microspeaker for hearing instruments” J. Micromech. Microeng. 14<br />
(2004) pp. 859-866.<br />
[6] I. Shahosseini1, E. Lefeuvre, J. Moulin, M. Woytasik, E.<br />
Martinicic, et al., "Efficiency optimization of a MEMS<br />
electrodynamics microspeaker", unpublished.<br />
[7] R. Liu, H. Wang, X. Li, J. Tang, S. Mao, and G. Ding, "Analysis,<br />
simulation and fabrication of MEMS springs for a micro-tensile<br />
system" J. Micromech. Microeng. 19, 2008, 015027 (10pp).<br />
[8] D. Bachmann, B. Schöberle, S. Kühne, Y. Leiner, C. Hierold,<br />
“Fabrication and characterization of folded SU-8 suspensions for<br />
MEMS applications”, Sensors and Actuators A 130–131, 2006, pp.<br />
379–386.<br />
262
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<br />
Modulation Instability in RF MEMS Devices<br />
Romolo Marcelli 1 , Giancarlo Bartolucci 1,2 , Giorgio De Angelis 1 , Andrea Lucibello 1 and Emanuela Proietti 1<br />
1 CNR-IMM Roma, Via Fosso del Cavaliere 100, 00133 Roma, Italy<br />
2 Dept. of Electronic Engineering, University of Roma “Tor Vergata”, Via del Politecnico 1, 00133 Roma, Italy<br />
Abstract- Modulation instability generated by mechanical<br />
frequencies in RF MEMS switches is predicted and its<br />
potential contribution to the RF signal degradation is<br />
discussed. In particular, evaluations have been performed for<br />
double clamped configurations in shunt capacitive devices. As<br />
a conclusion, it is evidenced the possibility for the excitation of<br />
satellites affecting as noise sources higher than -40 dB the<br />
spectral purity of microwave sources.<br />
I. INTRODUCTION<br />
High power effects in RF MEMS devices could be a<br />
limiting factor in their performances because of selfactuation<br />
of micro-switches [1][2][3] or non-linear response<br />
and excitation of satellite frequencies [4][5][6][7]. In the<br />
first case we have a failure of the device related to the unwanted<br />
actuation, eventually caused by micro-welding due<br />
to the increase of temperature during and after the bridge<br />
collapse. In the second case, the RF power carried by the<br />
signal flowing through the microwave transmission line can<br />
excite transversal and longitudinal mechanical modes,<br />
contributing to the degradation of the signal.<br />
In this paper the reconstruction of the spectrum due to the<br />
presence of mechanical resonances of the beam of a shunt<br />
connected RF MEMS switch is presented, with the aim to<br />
evaluate the contribution of inter-modulation products to the<br />
RF signal processed by the switch.<br />
II. RF SIGNAL PROCESSING IN RF MEMS<br />
Micro-electromechanical switches for Radio Frequency<br />
applications (RF MEMS switches) are movable microsystems<br />
which commute from an ON to an OFF state by<br />
means of the collapse of a metalized beam [8]. They can be<br />
actuated in several ways but, generally, the electrostatic<br />
actuation is preferred because no current is flowing in the<br />
device nor power absorption has to be involved in the<br />
process. In a coplanar waveguide (CPW) configuration, like<br />
that shown in Fig. 1 and Fig. 2, the bias DC voltage signal is<br />
usually separated with respect to the RF signal for<br />
application purposes. Anyway, in the simplest mechanical<br />
model, a voltage difference V is imposed between the metal<br />
bridge, connected to the ground plane of a coplanar<br />
waveguide (CPW) structure, and the central conductor of<br />
the CPW, which also carries the high frequency signal.<br />
Under these circumstances, the switch will experience an<br />
electrostatic force which is balanced by its mechanical<br />
stiffness, measured in terms of a spring constant k. The<br />
balance is theoretically obtained until the bridge is going<br />
down approximately (1/3) of its initial height. After that, the<br />
bridge is fully actuated, and it needs a value of V less than<br />
the initial one to remain in the OFF (actuated) position,<br />
because contact forces and induced charging effects help in<br />
maintaining it in the down position.<br />
Fig. 1. Typical capacitive RF MEMS shunt switch in CPW configuration<br />
Fig. 2. Cross section of the switch structure, where the metal bridge is<br />
suspended by means of dielectric anchors on a multilayer composed by: (i)<br />
the air gap g with respect to (ii) a metal thin layer at a floating potential<br />
(FM) to be used for improving the capacitance definition in the down<br />
position, (iii) a dielectric layer with thickness d deposited onto (iv) the<br />
metal M of the central conductor of the CPW, and finally (v) the SiO 2<br />
thermally grown layer onto the high resistivity silicon wafer.<br />
The RF power carried by the signal flowing in the central<br />
conductor of the CPW can be written as:<br />
RF<br />
out<br />
1<br />
2<br />
in<br />
2<br />
RF<br />
RF<br />
IVP<br />
RF<br />
Pin<br />
2 Z0<br />
(1)<br />
[ ]<br />
M<br />
1 V<br />
+−=<br />
PPPP<br />
r<br />
===<br />
Where: V RF and I RF are the effective RF voltage and<br />
current respectively, and Z 0 is the characteristic impedance<br />
of the line. P out is the output power, which is obtained from<br />
the input power P in decreased by the power coupled to the<br />
mechanical structure P M and the reflected power P r . The<br />
necessity to distinguish between the last two contributions<br />
depends on the different nature of the transferred power: the<br />
263
first one is due to the coupling of the EM field with the<br />
bridge, which senses a force induced by the RF voltage, and<br />
the second one is due to the electrical mismatch along the<br />
line, and it is almost independent of the presence of the<br />
bridge in a given location, for beams far at least 2 μm ca.<br />
from the central conductor of the CPW. On the other hand,<br />
for an almost perfectly matched line we can assume that the<br />
last contribution is negligible.<br />
The frequency of resonance for the bridge (or cantilever,<br />
or any other mechanical structure) is given by the well<br />
known equation:<br />
ω<br />
M<br />
k<br />
m<br />
1 k<br />
== π<br />
M<br />
;2<br />
ff<br />
M<br />
=<br />
(2)<br />
2π<br />
m<br />
i.e. the angular frequency ω M is defined by means of the<br />
spring constant k and the mass m, eventually modified in an<br />
effective value m eff with respect to the nominal one because<br />
of additional contributions (gas damping, holes, …) to be<br />
included and considered in the structure.<br />
Frequencies due to the longitudinal excitation modes have<br />
vlong<br />
to be also included. As well established flong<br />
= ,<br />
λlong<br />
where λ long is the wavelength of the longitudinal mode, and<br />
v long is the longitudinal velocity of the oscillating structure.<br />
For a double clamped configuration, we have λ long = 2L<br />
for the fundamental mode, where L is the full length of the<br />
T<br />
bridge, and v long = , where T is the tension and µ is<br />
μ<br />
the mass per unitary length. T is related to the constraints at<br />
the ends. For this reason, strain and residual stress on the<br />
beam due to the manufacturing process should play a<br />
dominant role. For µ, after some algebra, we can get<br />
μ eff = ρ eff LtA<br />
, being A eff the effective area of the bridge<br />
accounting for the presence of holes, t the thickness and ρ<br />
the density. Because of the above considerations we can<br />
write the following formula for f long :<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
power [7],[9]. The force sensed by the beam will be the<br />
result of the composition of voltage contributions coming<br />
from all of the above effects.<br />
By using the relation between the power and the energy<br />
P = ωE , and considering that the power processed by the<br />
MEMS has to be lower with respect to the threshold value<br />
needed for the actuation of the switch, we can also write:<br />
RF<br />
8 k<br />
2<br />
0 in<br />
=
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
⎛ ⎞<br />
⎜ −=<br />
M<br />
out PP 2P<br />
⎛ 2P<br />
⎞<br />
characteristic frequencies ω M and ω long .<br />
in 1 ⎟ ;<br />
⎝ P<br />
⎜ −=<br />
M<br />
IL Log 11 ⎟<br />
The 0<br />
approximated value for the actuation voltage for a<br />
in ⎠<br />
⎝ Pin<br />
⎠ (7)<br />
central actuation is obtained by using the definition of the<br />
spring constant for the entire structure. The spring constant<br />
The approach for calculating the absorbed power by<br />
k is a measure of the potential energy of the bridge<br />
longitudinal modes is the same given in Eq. (5), thus leading<br />
accumulated as a consequence of its mechanical response to<br />
to:<br />
the electrical force due to the applied voltage V. An<br />
approximated definition of it for central actuation can be<br />
given by [8],[10]:<br />
1<br />
2<br />
Plong<br />
= ωlongCVRF<br />
(8)<br />
2<br />
3<br />
k k k =+=<br />
K 32 Ewr<br />
i.e. the capacitance will be affected by both longitudinal<br />
and transversal modes, and, by using the same formalism<br />
introduced in the previous equations, the full power<br />
transferred to the mechanical system by the RF signal<br />
passing through the line will be:<br />
2<br />
( ) CV<br />
' 1<br />
PM<br />
M ωω+= long RF<br />
(9)<br />
2<br />
The above equation is the measure of the total power<br />
transferred to the beam because of the RF signal to both<br />
longitudinal fundamental mode and transversal mode.<br />
Higher order longitudinal modes will absorb a power<br />
fraction scaled by the order of the excited mode, with lower<br />
amount of power for the highest modes.<br />
It is worth noting that in the spectrum reconstruction the<br />
above contributions have to be separated, leading to<br />
different values of the peak power. In particular, we should<br />
have the following distribution, which accounts also for the<br />
excitation of the two satellites:<br />
I<br />
I<br />
I<br />
out<br />
out<br />
out<br />
( ω )<br />
RF<br />
( )<br />
RF<br />
long<br />
( ωω ) ω CZ<br />
RF<br />
M<br />
in<br />
long<br />
that the frequency of resonance and the capacitance<br />
associated to the beam can be calibrated to fully absorb the<br />
RF signal. Such a result is particularly interesting for<br />
resonating structures based on double clamped<br />
configurations, because the maximum absorption<br />
corresponds to a resonance condition. Actually, such a<br />
device is a notch filter and it could be used as a feedback<br />
element in a one port oscillator. Another conclusion coming<br />
out from Eq. (10) is that the power released to the<br />
mechanical structure does not depend on the frequency of<br />
the carrier, but just on the geometry of the beam and its<br />
+ K<br />
2<br />
where:<br />
K<br />
1<br />
r =<br />
=<br />
t<br />
L<br />
E<br />
σ<br />
[ ( − ) wr ]<br />
1<br />
( )<br />
18 νσ<br />
1<br />
; K<br />
2<br />
⎛ L ⎞⎛<br />
L ⎞<br />
⎜ 22<br />
−− ⎟⎜<br />
⎟<br />
⎝ L ⎠⎝<br />
L ⎠<br />
1<br />
=<br />
Lc<br />
2 −<br />
L<br />
2<br />
cc<br />
(11)<br />
(12)<br />
L is the bridge total length, L c is the switch length in the<br />
RF contact region (width of the central conductor of the<br />
CPW), w is the bridge width, t is the Au thickness of the<br />
bridge. The other parameters are the Young modulus E, the<br />
residual stress σ and the Poisson coefficient ν. As well<br />
established, the Young modulus is an intrinsic property of<br />
the material, and specifically it is a measure of its stiffness.<br />
Let’ use, as an example, the following structure for a RF<br />
MEMS switch in coplanar waveguide (CPW) configuration:<br />
L=600 μm as the bridge total length, L<br />
2( M<br />
PP<br />
long<br />
)<br />
c =300 μm as the<br />
+<br />
1−=<br />
Z021<br />
(<br />
M<br />
+− ωω<br />
long<br />
) C<br />
switch length<br />
=<br />
in the RF contact region (width of the central<br />
P<br />
conductor of the CPW), w=100 μm as the bridge width,<br />
in<br />
w S =100 μm for the switch width (transversal dimension of<br />
PM<br />
ωωω<br />
M<br />
==±<br />
0 MCZ<br />
(10) the switch, parallel with respect to the CPW direction),<br />
Pin<br />
d=thickness of the dielectric material=0.2 μm, with<br />
dielectric constant ε=3.94 (SiO<br />
P<br />
2 ), t=1.5 μm for the gold<br />
long<br />
==± bridge, ρ=19320 kg/m 3 for the gold density, E=Young<br />
0 long<br />
P<br />
modulus=80×10 9 Pa, ν=0.42 for the metal Poisson<br />
coefficient and σ=18 MPa as the residual stress of the metal<br />
From the first of Eq. (10) it is worth noting that the (measured on specific micromechanical test structures). A<br />
intensity of the central peak could vanish under the uniform distribution of holes with 5 µm radius and distant<br />
condition 21 Z 0 ( ω ω ) C =+−<br />
0 . This is an evidence 10 µm each other has been also considered, leading to<br />
effective values in terms of the beam area and spring<br />
constant.<br />
A recent experimental approach was also adopted for<br />
evaluating the contribution of the spring constant and for<br />
modeling it on the base of nano-indentation techniques[9].<br />
All the quantities previously introduced have to be redefined<br />
because of the presence of holes in the released<br />
beam. The holes need to be used for an easier removal of<br />
the sacrificial layer under the beam, and for mitigating the<br />
stiffness of the gold metal bridge, i.e. for better controlling<br />
the applied voltage necessary for collapsing it, to have not<br />
values too high because of the residual stress.<br />
265
11-13 <br />
May 2011, Aix-en-Provence, France<br />
In this framework, we have re-calibrated the material<br />
<br />
to ω M =1.38x10 5 s -1 , ω long =7.66x10 7 s -1 ). By using Eq. (10)<br />
properties accounting for the holes distribution on the metal we get the normalized values I(ω RF ±ω M )≈10 -6 and<br />
beam. Literature definitions [8],[11],[12], are generally I(ω RF ±ω long )≈5.7x10 -4 . The performed computation is valid<br />
accepted for analytically describing the effect of the holes for a bridge 100 µm wide. The characteristic frequencies are<br />
by means of the pitch, i.e. the center-to-center distance p not affected by the change in the area, but it is true for the<br />
between the holes and the edge-to-edge distance l. The capacitance C, which is proportional to the area. By using<br />
situation is explained in the Fig. 3. In this way, the ligament beams 50, 100 and 200 µm wide respectively, the<br />
efficiency will be given by the term (1-(l/p)) and such a term corresponding normalized intensities for the mechanically<br />
will be used in this paper for evaluating the effective coupled power will be I(ω RF ±ω M )≈(0.5, 1.0 and 2)×10 -6<br />
quantities which are decreased with respect to the original whereas I(ω RF ±ω long )≈(2.8, 5.7 and 11.4)×10 -4 . By<br />
one. Following this approach, σ eff =σ(1-(l/p)), while expressing these values in dB, we experience, for a matched<br />
E eff =E(1-(l/p)), ν eff =ν(1-(l/p)).<br />
CPW line, a negligible decrease in the output intensity. In<br />
For the effective mass, we preferred to use a definition particular, a -60 dB level is expected for transversal<br />
based on the ratio between the area with and without the contributions, and -40 dB for the longitudinal ones. In real<br />
holes, thus obtaining m eff =m(A/A 0 ), where A 0 is the situations, also the line can be not perfectly matched,<br />
geometrical area of the beam and A is the effective one because of the presence of the bridge, leading to some<br />
considering the presence of the holes. All the evaluations power loss due to the grounded metal beam surmounting the<br />
which will be shown in this paper are based on the central conductor of the CPW. In fact, from the<br />
previously defined quantities, calculated accounting for manufactured actual device shown in Fig. 4 we got the<br />
their effective contribution.<br />
transmission loss measured in Fig. 5, clearly higher than<br />
that expected just for the mechanical coupling, which<br />
should be 4×10 -3 dB in the worst case (longitudinal<br />
excitation).<br />
(a)<br />
(b)<br />
Fig 3. Typical shape of (a) a perforated beam used for RF MEMS<br />
double clamped switches and pitch definition, and (b) enlarged view of the<br />
actual device. Holes are realized for facilitating the sacrificial layer<br />
removal and their position, number and dimensions are properly tailored<br />
depending on the application.<br />
From the residual stress we can get a value of the tension<br />
by assuming that k σ in Eq. (11) is the longitudinal force per<br />
unitary length on the beam, and Eq. (3) can be transformed<br />
into:<br />
f<br />
long<br />
1<br />
=<br />
2L<br />
v<br />
λ<br />
long<br />
long<br />
T<br />
eff<br />
1 T<br />
==<br />
2L<br />
μ<br />
eff<br />
1 kσ<br />
=<br />
2LLm<br />
m<br />
eff<br />
(13)<br />
From the above equations, and considering that Z 0 =50<br />
ohm, C=0.15 pF, it turns out V threshold =12 V ca for central<br />
actuation, f M =20 kHz, and f long =13 MHz (which corresponds<br />
Fig. 4. Test-fixture structure of the manufactured RF MEMS switch. The<br />
input and output ports are connected to a vector network analyzer by means<br />
of coplanar probes for on-wafer characterization.<br />
S 21<br />
[dB]<br />
0.0<br />
-0.2<br />
-0.4<br />
-0.6<br />
-0.8<br />
Simple Line<br />
W = 50 μm<br />
W = 100 μm<br />
-1.0<br />
0 5 10 15<br />
Frequency [GHz]<br />
Fig. 5. Measured Insertion Loss for the RF MEMS switch having the bridge in<br />
up position (ON state). A simple CPW line is compared with the response of<br />
the 50 and 100 µm wide beams.<br />
266
So far, the expected major contribution in low power<br />
regime of the satellite excitation will be not in having an<br />
increase in the losses, but in those applications where signal<br />
routing is essential for re-directing the RF output by means<br />
of filtering stages needing a high rejection ratio, low ripple<br />
level and a very high spectral purity. Moreover, we think<br />
that the simplified approach given in Eq. (10), based on low<br />
power computation, has to be improved for medium and<br />
high power applications, where a non-linear treatment of the<br />
coupling, which will be power dependent, is more correct.<br />
As a matter of fact, to have noise sources higher than -40 dB<br />
will have contra-indications in the power spectral purity of<br />
Digital Synthesizers, because the contributions will be close<br />
to the source carrier and not far like the high order<br />
harmonics, which can be easily canceled by means of bandpass<br />
filters. Experimental analysis in the frequency domain<br />
is in progress for the exact determination of the satellites<br />
contribution to the spectral purity of the RF signal.<br />
III. CONCLUSION<br />
In this paper, the contribution of the input power to the<br />
modulation instability of RF MEMS devices associated to<br />
transversal and longitudinal oscillation modes of a double<br />
beam structure has been studied.<br />
As a result, the transversal mode will be the main<br />
responsible for signal degradation, because of its proximity<br />
with respect to the RF carrier frequency, but the longitudinal<br />
one could also contribute for losses and degradation<br />
depending on the system requirements.<br />
REFERENCES<br />
[1] Karl M. Strohm et al., “RF-MEMS Switching Concepts for High<br />
Power Applications”, Proceed. of 2001 IMS, pp.42-46 (2001).<br />
[2] B. Pillans, J. Kleber, C. Goldsmitht, M. Eberly, Proceedings of the<br />
2002 IEEE MTT-Symposium 329 (2002).<br />
[3] E.P. McErlean, J.-S. Hong, S.G. Tan, L. Wang, Z. Cui, R.B. Greed<br />
and D.C. Voyce, IEE Proceedings on Microwave Antennas Propagation,<br />
Vol.152, 449 (2005).<br />
[4] J. B. Muldavin and G. M. Rebeiz, “Nonlinear Electro-Mechanical<br />
Modeling of MEMS Switches”, Proceed. of IEEE MTT Symposium,<br />
pp.21119-2122 (2001).<br />
[5] Conor O’Mahony Russell Duane Martin Hilland Alan Mathewson,<br />
“Electromechanical Modelling of Low-Voltage RF MEMS Switches”,<br />
Proceed. of DTIP Montreux, Switzerland, 12-14 May 2004.<br />
[6] E.K. Chan, E.C. Kan and R.W. Dutton: “Nonlinear Dynamic<br />
Modeling of Micromachined Switches”, Proceed. of IEEE MTT-Symposium,<br />
pp.1511-1514 (1997).<br />
[7] E. Brusa and M. Gh. Munteanu, “Role of Nonlinearity and Chaos<br />
on RF-MEMS Structural Dynamics”, Proceed. of DTIP 2009 Conference,<br />
Roma, 1-3 April 2009.<br />
[8] G. M. Rebeiz, “RF MEMS, Theory, Design and Technology”, John<br />
Wiley and Sons, Hoboken, 2003.<br />
[9] Robert W. Stark, “Bistability, higher harmonics, and chaos in<br />
AFM”, Materials Today, Vol. 13, Issue 9, pp.24-32 (2010).<br />
[10] Balaji Lakshminarayanan, Denis Mercier, and Gabriel M. Rebeiz<br />
“High Reliability Miniature RF-MEMS Switched Capacitors”, IEEE Trans.<br />
on Microwave Theory and Tech., Vol.56, No. 4, pp.971-981 (2008).<br />
[11] V. L. Rabinov, R. J. Gupta and S. D. Senturia, “The effect of<br />
release etch-holes on the electromechanical behaviour of MEMS structures”,<br />
in IEEE Int. Conf. on Solid-State Actuators, Chicago, pp. 1125-1128, 1997<br />
[12] G. De Pasquale, T. Veijola, and A. Somà, “Gas Damping Effects<br />
on Thin Vibrating Gold Plates: Experiment and Modeling”, Proceed. of DTIP<br />
2009 Conference, Roma, 1-3 April 2009.<br />
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May 2011, Aix-en-Provence, France<br />
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267
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Study of Screen-printing Microlens Array Using<br />
Electroforming Molds<br />
Ming-Je Lin 1 , Hsiharng Yang 1 , Feng-Tsai Weng 2<br />
1 Institute of Precision Engineering, National Chung Hsing University, Taichung, Taiwan 402<br />
2 Institute of Mechanical and Electro-Mechanical Engineering, National Formosa University, Yunlin, Taiwan 632<br />
Abstract- This paper aims to produce micro-lens array by<br />
using MEMS technology. It combines LIGA-like technology for<br />
micro-lens array fabrication to design and integration of<br />
screen-printing process. In the experiment, we found that the<br />
working temperature played an important role in shape transfer<br />
process. The microlens density and arrangement also affect the<br />
mean height of photoresist after thermal reflow. This study uses<br />
LIGA-like technologies at the early stage to develop<br />
screen-printing mold. Then, the electroforming screen mold can<br />
be re-used, that synthetic development of high-mirror high (high<br />
sag) micro-lens array, this is an innovative and unique method.<br />
The experiment success and make the simple micro-lens<br />
manufacturing process, with forming fast, easy to melt, reduce<br />
costs and micro-structure, appearance, easy to retain the<br />
advantages.<br />
I. INTRODUCTION<br />
Microlens array is the key component to the miniaturization<br />
of conventional optical devices. The field of micro-optics<br />
plays an important role in visual display products such as,<br />
liquid crystal displays (LCDs), mobile phone panels and<br />
personal digital accessories (PDAs). One major benefit of<br />
using microlenses is that they enhance the illumination<br />
brightness and simplify light-guide module construction. In a<br />
laptop display, a 25% increase in light output has been<br />
reported when using the microlens technology [1]. There are<br />
other potential benefits too, such as focal plane optical<br />
concentration, optical efficiency enhancements, color<br />
separation, beam shaping and miniature optical scanning.<br />
Micromanufacturing technology allows compact, and<br />
mini-features to be fabricated. Micro-electro-mechanical<br />
system (MEMS) technology has a growing number of<br />
applications in military, industrial, and consumer markets.<br />
For this reason, many academic and research institutions are<br />
currently involved in MEMS technology product research and<br />
development. Component miniaturization is a common<br />
objective in electro-optical systems. Miniaturizing devices<br />
using micro-optics has revolutionized many electro-optical<br />
systems - including video cameras, video phones, compact<br />
disk data storage, robotic vision, optical scanners, and high<br />
definition projection displays. Higher accuracy and lower<br />
microlens fabrication costs are needed to meet the rapid<br />
growth in demand for these devices. Micro-scale refractive<br />
lenses offer several important features: significantly reduced<br />
wavelength sensitivity compared to diffractive optics<br />
(necessary for broadband applications), the possibility of very<br />
large numerical apertures and high light efficiency.<br />
Several fabrication techniques have been applied to the<br />
refractive microlens fabrication processes. One method of<br />
fabricating refractive microlenses is by melting cylindrical<br />
photoresist posts. This is known as microlens reflow<br />
processing [2]. Photoresist cylinders are formed using a<br />
lithographic process and then heated above the photoresist<br />
glass transition temperature. Surface tension causes the<br />
photoresist cylinders to assume a spherical shape. Surface<br />
tension also leads to relatively short focal lengths in the<br />
resulting microlenses (i.e., high numerical apertures). The<br />
reflow process produces large microlens arrays. This process<br />
is extraordinary compared with conventional macro-optic<br />
fabrication methods. In very large scale integration (VLSI)<br />
based processing techniques, coherent refractive microlens<br />
arrays are made on a silicon surface using a combination of<br />
lithography and reactive ion etching (RIE) techniques.<br />
Multi-level photoresist mask patterning and sequential RIE<br />
are used to form binary optic microlens arrays. A laser writing<br />
system for continuous-relief microoptical element fabrication<br />
in photoresist was described by Gale et al. [3]. The<br />
photoresist-coated substrate was exposed using x-y raster<br />
scanning under a focused HeCd laser beam (λ=442 nm),<br />
synchronously programmable controlled in intensity to write<br />
two-dimensional (2-D) exposed patterns. Further<br />
development of 3-D microstructures with analogous topology<br />
using excimer laser ablation (λ=248 nm) produced versatile<br />
micro-optic applications [4]. Microlens arrays with lateral<br />
dimensions from 10 to 1000μm and profile heights of up to<br />
10μm were fabricated using this technique. An optimal gray<br />
scale mask is required to produce fine roughness.<br />
Micro-optics printing technology prints a number of droplets<br />
onto a substrate to form circular microlens arrays [5].<br />
Microlenses ranging in diameter from 20μm to5mmhave<br />
been fabricated in this way. The piezoelectric actuator-based<br />
and drop-on-demand ink-jet printing method was developed<br />
to control different fluid volumes. Liquid droplets were<br />
dispensed onto a substrate to form refractive microlens arrays.<br />
The use of deep x-ray lithography to fabricate microoptical<br />
components shows great potential for mass production. Lee et<br />
268
al. used a modified LIGA (German acronym for LIthografie,<br />
Galvanoformung, and Abformung) process to fabricate<br />
microlenses by melting the deep x-ray irradiated pattern onto<br />
a PMMA (poly-methyl methacrylate) substrate. Using this<br />
technique, microoptical components of any desired shape can<br />
be fabricated [6, 7]. The resulting components have smooth<br />
and vertical sidewalls, lateral dimensions in the micrometer<br />
range, and sag heights of several hundred micrometers. A<br />
molding process (either injection molding or hot embossing)<br />
is required before mass production can be achieved. The<br />
microlens array mold or mold inserts play an important role in<br />
the mass molding production process. This replication<br />
process promises the desired profile as final products.<br />
A new method for producing microlens array with large<br />
sag heights was investigated for integrated fluorescence<br />
microfluidic detection systems [8]. Three steps in that<br />
production technique were included for concave microlens<br />
array formations to be integrated into microfluidic systems.<br />
The micro concave lens molds were then finished and ready<br />
to produce convex microlens in PDMS material. Using a<br />
LIGA-like process to fabricate microlens arrays is<br />
considerably less expensive using a UV exposure tool instead<br />
of deep x-ray lithography. A new microlens array fabrication<br />
method using a UV proximity printing method has been<br />
invented [9]. It uses a slice to control the gap size, resulting in<br />
microlens array formation in the resist. However, this method<br />
was limited to round microlens arrays with low sag heights.<br />
They produced microstructures with smooth surfaces, high<br />
yield rates, and good reliability.<br />
The LIGA-like process provides microlens array<br />
fabricators with high optical quality at low cost. By using the<br />
vacuum pressure to form a microlens array was investigated<br />
[10]. This vacuum suction technique is feasible for certain<br />
microlens array fabrication sizes. Based on the LIGA-like<br />
technology development, this paper will present the<br />
promising technique using the LIGA-like process to pressing<br />
microlens array and investigate the processing parameters for<br />
making microlens array.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Fig. 1. Illustration of the screen-printing process for microlens array<br />
fabrication.<br />
2.1 Contact angle measurement<br />
The liquid contacts on a solid surface, there is a contact<br />
angle between the solid and liquid drop surfaces. The surface<br />
hydrophobicity of the substrate can determine the microlens<br />
profile. It is necessary to find the contact angles between the<br />
photoresist and different substrates. A surface tension<br />
examiner (FTA200) was used to measure the contact angle.<br />
An example to measure the contact angle between water and<br />
copper coating substrate, the contact angle is 78.35°. The<br />
contact angle between water and silver coating substrate is<br />
60.98°. The contact angle between water and stainless steel<br />
304 is 56.67°. A low contact angle between water and glass<br />
substrate is 22.24° as shown in Fig. 2. The further<br />
experiments to measure the contact angle between photoresist<br />
AZ4620 and stainless steel 304, sopper coating substrate,<br />
glass substrate, the resulted contact angle are 26.43°, 39.33°<br />
and 42.42°. It means that the larger contact angle can result in<br />
a high sag mirolens. From the above experiments, the glass<br />
substrate is chosen for screen printing mirolens array in<br />
photoresist.<br />
II EXPERIEMNTS<br />
The fabrication process mainly applies the LIGA-like<br />
technology. In the conventional microlens array fabrication,<br />
photoresist patterns are formed by lithography process, it<br />
includes mask pattern design, photoresist coating, UV<br />
exposure and development, and thermal reflow. Microlens<br />
array in photoresist is formed by the above steps. The further<br />
mass production will apply electroforming to replicate the<br />
microlens array mold. A different approach is to pattern an<br />
electroforming mold, then directly screen printing photoresist<br />
patterns and thermal reflow formicrolens array fabrication. It<br />
will be suitable for mass production of Microlens array by<br />
using the same mold. The fabrication process is illustrated in<br />
Fig. 1.<br />
Fig. 2 Contact angle measurement of water and glass substrate.<br />
2.2 Lithography process<br />
The lithography process used a PET mask with pattern<br />
layout design is illustrated in Fig. 3. Eight patterns with four<br />
diameters 30, 45, 60, and 80μm as well as two different<br />
spacings 20 and 40 μm are included. Since negative<br />
photoresist JSR THB-126N was used, the resulted patterns<br />
were micro-post array. The opening area is exposed to UV<br />
lithography, monomers in negative photoresist are<br />
cross-linking by photons. Micro-post array in photoresist<br />
remains after development. Micro-post array height is<br />
controlled by spin coating thickness. The relationship is<br />
269
Development 3.5min<br />
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May 2011, Aix-en-Provence, France<br />
depicted in Fig. 4. Theoretically, the JSR-THB-126N<br />
<br />
600mJ/cm 2<br />
photresist thickness can be more than 100μm. However, the Hot process<br />
160℃ , 30 min<br />
experimental result was achieved only 23μmthick as shown<br />
in Fig. 5. The related parameters in lithography process is<br />
listed in Table 1. After complete micro-post array, Ni<br />
electroforming was applied to make the screen mesh mold.<br />
Fig. 3. Illustration of PET mask design.<br />
Fig. 4. Relationship between spin speed and photoresist thickness when spin<br />
coating JSR-THB126N.<br />
(a)<br />
Fig. 5. Micro-post thickness measurement.<br />
Table 1 Experimental parameters in lithography process.<br />
Base plate clean H 2SO 4:H 2O 2=3:1 wash<br />
Acetone:60 min<br />
DI Water wash,N 2 dry<br />
120℃bake 20 min dry<br />
Seed layer deposition Sputtering Ag 200 nm<br />
Spin coating Spread : 500 rpm 10 sec<br />
Spin : 2000-800 rpm 30 sec<br />
Soft bake 90℃ 3min<br />
Hold 5 min<br />
Exposure<br />
350W, Near UV<br />
2.3 Ni electroforming screen mesh mold<br />
The electrolyte composition of Ni electroforming<br />
include nickel sulfamate 450 g/L, 40 boric acid 40 g/L, and<br />
wetting agent 3 mL/L. The wetting agent is also called<br />
interface surfactant, it reduces the electrolyte surface tension<br />
to help hydrogen bubbles away from the substrate surface.<br />
The electrolyte pH value was controlled between 3.7 and 4.0.<br />
Since the pH value trends to slightly increase during<br />
electroforming, the initial pH value is set to 3.7. The<br />
operating temperature was controlled at 45℃. The starting<br />
current density was 1 ASD (A/dm 2 ). A high current density<br />
(larger than 7 ASD) may result in pin holes and large grains, a<br />
low current density also results in low growth rate and<br />
impurities. Once complete Ni electroforming, the mold was<br />
immersed into DI water with ultrasonic agitation to release<br />
from the substrate. The finished screen printing mold in<br />
nickel with 30mm×20mm size is shown in Fig. 6(a). The<br />
optical microscopic photograph showing the spacing between<br />
two openings is 25μm.<br />
(b)<br />
Fig. 6. Ni electroforming screen printing mold; (a) outlook of the screen<br />
printing mold, (b) OM photograph of the mold.<br />
2.4 Screen printing process<br />
The finished electroforming mold is a thin sheet. A frame<br />
to support the sheet is machined by a CNC machine as shown<br />
in Fig. 7(a). Then the sheet mold is attached to the supporting<br />
frame as shown in Fig. 7(b). The screen printing process was<br />
performed as described in Fig. 1. The screen mold was to<br />
define photoresist patterns after scraping. The achieved<br />
patterns by using different mold diameters are shown in Fig. 8.<br />
After defining patterns, the specimen was put in the oven at<br />
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<br />
temperature 120 ℃ for 2 minutes. The OM photographs of<br />
patterns before and after thermal reflow are shown in Fig. 9.<br />
(a)<br />
(b)<br />
Fig. 7. The frame (a) and screen mold (b) for screen printing.<br />
(a)<br />
(b)<br />
Fig. 9. OM photographs of patterns before (a) and after (b) thermal reflow.<br />
III. RESULTS<br />
The finish screen mesh mold has the average thickness<br />
16μm. Eight microlens arrays with different patterns were<br />
completed by using this screen printing process. The large<br />
lens diameter trends to have a low sag height. An example is<br />
shown in Fig. 10. Microlens with diameter 108.5μm, its sag<br />
height is 3.87μm. Comparing a smaller microlens with<br />
diameter 85µm, its sag height is 7.6µm. Since the screen<br />
mold has the same thickness, the large lens diameter resulted<br />
in a large radius of curvature. The sag height is getting<br />
smaller. The experiment is successful to make the microlens<br />
array in a simple manufacturing process with fast formation,<br />
cost reduction and controllable microstructure appearance<br />
advantages. The further related process technology is benefit<br />
to optical lens industry.<br />
(b)<br />
Fig. 8. Pattern profiles of differenet mesh mold diameters; (a) 60μm, (b)<br />
100μm.<br />
(a)<br />
(a)<br />
(b)<br />
Fig. 10. Microlens array fabricated by screen printing; (a) OM photograph,<br />
(b) 3D profile.<br />
IV CONCLUSION<br />
The experiment was successful to use the screen printing<br />
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mold for fabricating Microlens array. The thermal reflow<br />
<br />
temperature is 120 ℃ enable to melt micro-post patterns.<br />
Microlens arrays with different sags ranged from 3.87 to 7.6<br />
µm are completed. This technology is scalable for large area<br />
microlens array replication.<br />
ACKNOWLEDGMENT<br />
This work was supported by the National Science Council<br />
(series no. NSC 99-2221-E-150-047) of Taiwan<br />
REFERENCES<br />
[1] B. Ezell,“Making microlens backlights grow up,”Inf. Disp.,17,<br />
pp.42–45, 2001.<br />
[2] N.F.Borrelli,D.L.Morse,R.H.Bellman,andW.L.Morgon,<br />
“Photolytic technique for producing microlenses in photosensitive<br />
glas,” Applied Optics, 24, 2520, 1985.<br />
[3] M.T. Gale, M. Rossi, J. Pedersen and H. Schutz H.,“Fabrication of<br />
continuous-relief micro-optical elements by direct laser writing in<br />
photoresists,”Optical Engineering, vol. 22, no. 11: 3556-3566,<br />
1994.<br />
[4] K. Zimmer, D. Hirsch and F. Bigl,“Excimer laser machining for the<br />
fabrication of analogous microstructures,”Applied Surface Science,<br />
vol. 96-98: 425-429, 1996.<br />
[5] W.R. Cox, T. Chen and D. Hayes,“Micro-Optics fabrication by<br />
ink-jet printing,”Optics & Photonics News, vol. 12, no. 6: 32-35,<br />
2001.<br />
[6] J. Goettert and J. Mohr,“Characterization of micro-optical components<br />
fabricated by deep-etch x-ray lithography”SPIE: Micro-Optics II,<br />
vol. 1506: 170-178, 2002.<br />
[7] S.K. Lee et al.,“A simple method for microlens fabrication by the<br />
modified LIGA process,” Journal of Micromechanics and<br />
Microengineeing, vol. 12: 334-340, 2002.<br />
[8] H. Yang, R.F.Shyu,J.-W. Huang,“New production method of<br />
convex microlens arrays for integrated fluorescence microfluidic<br />
detection systems,”Microsystem Technologies, vol. 12, pp. 907-912,<br />
2006.<br />
[9] C.-P. Lin, H. Yang, C.-K. Chao,“A new microlens array fabrication<br />
method using UV proximity printing,”Journal of Micromechanics<br />
and Microengineering, vol. 13: 748-757, 2003.<br />
[10] R. F. Shyu and H. Yang,“A promising thermal presing used in<br />
fabricating microlens array,”International Journal of Advanced<br />
Manufacturing Technology, vol. 36, pp. 53-59, 2008.<br />
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<br />
Polymer-based Fabrication Techniques for Enclosed<br />
Microchannels in Biomedical Applications<br />
Annabel Krebs, Thorsten Knoll, Dominic Nussbaum, Thomas Velten<br />
Fraunhofer Institute for Biomedical Engineering<br />
Ensheimer Str. 48, 66386 St. Ingbert, Germany<br />
Abstract- Investigations and analyses of body fluids like<br />
serum or whole blood are essential tasks in biomedical<br />
research in order to understand and diagnose diseases, to<br />
conduct pharmacological tests or to culture cells. Therefore,<br />
microfluidic systems provide a favorable tool for processing<br />
fluid samples as they allow downscaling of sample volumes and<br />
handling of single fluid components such as cells or proteins.<br />
For this reason, we present simple fabrication techniques for<br />
microchannel systems using polymer materials only. On the<br />
one hand, these materials are low-priced compared to<br />
conventional silicon or glass. On the other hand, they do not<br />
show any interaction with biological fluids. Furthermore, their<br />
transparency guarantees an easy observability of all processes<br />
within the system. Depending on the channel dimensions,<br />
different adhesion bonding techniques for closing of the<br />
systems are applied, whereas the fluidic interfaces are<br />
included. Summing up, we provide complete fabrication<br />
processes for fluidic systems which are simpler and more costeffective<br />
than conventional methods and yet cope with all<br />
essential requirements for microfluidic applications.<br />
I. INTRODUCTION<br />
Microfluidic systems have become a common tool in<br />
research fields like analytical chemistry or biomedicine as<br />
miniaturized devices require only small material and sample<br />
volumes. Additionally, effects can be exploited which are<br />
only dominant in the micro range. For example,<br />
investigations on cells or other components in blood<br />
samples can be carried out in lab-on-chip systems as<br />
dimensions of the cells are in the same scale as the<br />
microchannels. In biomedicine, lab-on-chip systems have<br />
emerged to indispensable devices for the accomplishment of<br />
medical tests with body fluids or extractions from fluids.<br />
Besides, they also serve for cell handling and culturing,<br />
mixing of liquids, detection and analysis of diseases or for<br />
measuring quantitative amounts of components like glucose<br />
or hormones in blood [1-6].<br />
Depending on the applications of micro systems, certain<br />
requirements to the systems need to be met. In case of<br />
biomedical applications, the transparency of materials is<br />
often a vital aspect in order to be able to follow processes<br />
within the system channels or chambers. Moreover, the<br />
employed materials should not interact with the biological<br />
sample and channels should hold defined dimensions.<br />
Other general demands come along such as leak-proof<br />
closure of the channels and implementation of fluidic<br />
interfaces. Importantly, the whole fabrication procedure has<br />
to be affordable at the same time.<br />
Therefore, we present fabrication techniques for<br />
microfluidic systems with focus on biomedical applications,<br />
meeting the requirements named above. Based on acrylic<br />
glass substrates and the epoxy resists SU-8 and<br />
PerMX3020, reasonable cost of the materials is assured.<br />
These polymer materials are pellucid and do not show<br />
interactions with whole blood or blood components. As<br />
biocompatibility tests with SU-8 have not given cause for<br />
concern [7, 8], SU-8 is currently used in divers biological<br />
research fields [9, 10], although further biocompatibility<br />
studies might be advisable [8]. In Ref. [5], we introduced a<br />
manufacturing technique which also bases upon these<br />
polymer materials. Yet, the here presented, modified<br />
processes enable a wider choice of material combinations as<br />
well as enhanced fabrication reliability and yield.<br />
Additionally, we achieved to double the feasible aspect<br />
ratio.<br />
Hence, a simple fabrication technique for microchannels<br />
is presented which rests upon photolithography and polymer<br />
adhesion bonding. In doing so, very small channels with<br />
aspect ratios higher than 10:1 can be created. In contrast to<br />
conventional hybrid techniques for sealing of channels, we<br />
use different full wafer bonding methods depending on the<br />
channel dimensions. Unlike other working groups that have<br />
already presented high aspect ratios and full wafer adhesion<br />
bonding using silicon or glass substrates [11-13], we<br />
accomplish these processes on polymer substrates. Very<br />
cost-effective and simple structuring techniques can be<br />
applied to these substrates, e.g. fluidic interfaces can be<br />
implemented by mechanical drilling. More complicated,<br />
costly procedures like laser drilling, sand blasting, glass<br />
etching and substrate removals can be obviated. Altogether,<br />
we obtain a complete manufacturing procedure preferable<br />
for biomedical applications which outplays common silicon<br />
or glass techniques in terms of material and total costs. In<br />
comparison with other polymers like polydimethylsiloxane<br />
(PDMS), the presented techniques are adaptive for smaller<br />
channels and higher aspect ratios, thus they qualify for a<br />
wider range of applications.<br />
II.<br />
MATERIALS AND METHODS<br />
A. Materials<br />
1 mm thick 10 cm x 10 cm acrylic glass (polymethyl<br />
methaacrylate, PMMA) plates were used as substrate<br />
materials. Two different epoxy-based photoresists, SU-8<br />
273
and the dry film resist PerMX3020 (DFR), were used as<br />
channel layers. These polymers hold similar chemical<br />
compositions and can be structured via photolithography.<br />
Either SU-8 or DFR also served as adhesive layer for<br />
closing of the channel systems.<br />
B. Channel Layer Fabrication<br />
The PMMA substrate plate was rinsed with isopropyl<br />
alcohol (IPA) and nitrogen. After that, all channel structures<br />
were defined using photolithography. Therefore, two<br />
optional methods and materials were used, either spin<br />
coating of SU-8 or lamination of DFR.<br />
SU-8 was chosen as channel layer material when the<br />
fabrication of very narrow structures with aspect ratios<br />
higher than 2:1 was intended. As it was found out earlier<br />
that structured SU-8 features a stronger adhesion to another<br />
SU-8 layer than to PMMA [5], a thin layer of SU-8 served<br />
as an adhesion layer for the channel SU-8 layer. For this<br />
purpose, a thin layer of SU-8 was spun onto the substrate.<br />
After a pre-exposure bake, the SU-8 was fully exposed to<br />
UV-light and post-baked to achieve an entire<br />
polymerization. Then, the second SU-8 layer was spincoated<br />
on top of the first one and pre-baked. The fluidic<br />
systems were defined during exposure to UV-light using a<br />
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<br />
1. Spin-coating of SU-8 or<br />
lamination of DFR (adhesive<br />
layer) on PMMA substrate<br />
2. UV- Exposure<br />
photo mask which contains the channel structures.<br />
Thereafter, a post-exposure bake took place followed by a<br />
development in a PGMEA developer solution.<br />
For channel dimensions with aspect ratios lower than 2:1,<br />
DFR was opted for the active structure. By means of a<br />
desktop laminator, DFR was laminated on the substrate and<br />
pre-baked. Like in the case of SU-8, the first DFR was fully<br />
exposed for complete cross-linking to form an adhesive<br />
layer. The second layer was exposed using a mask which<br />
contains the channel systems. This layer was post-baked<br />
and also developed in PGMEA. The process flow for the<br />
channel layer fabrication including the fluidic interfaces is<br />
pictured in Fig. 1.<br />
C. Fluidic Interfaces<br />
In contrast to Ref. [5], fluidic interfaces were realized by<br />
CNC-assisted mechanical drilling of the structured PMMAepoxy-stack.<br />
In order to protect the fluidic structures from<br />
contaminations and mechanical damage, they were covered<br />
by a protective foil (V-8-T, Nitto) which can easily be<br />
drawn off after drilling (see Fig. 1).<br />
D. Closing of Channels<br />
Three different techniques of adhesive bonding were<br />
applied and tested for the closure of channels with different<br />
dimensions. The bonding options are shown in Fig. 2.<br />
The first option was to use an additional DFR layer that<br />
was laminated on top of the channel layer, like it was<br />
presented before for moderately large channels with 220 µm<br />
width [6]. This last DFR was fully exposed to serve as the<br />
lid or bottom, respectively.<br />
3. Spin coating of SU-8 or<br />
lamination of DFR<br />
(channel layer)<br />
First bonding approach<br />
1. Lamination of DFR on<br />
top of the channel layer<br />
4. UV-Exposure through<br />
photo mask<br />
2. UV-Exposure of<br />
DFR<br />
5. Development of SU-8/<br />
DFR in devmr600<br />
6. Covering with<br />
protective foil<br />
7. Mechanical drilling of<br />
fluidic interfaces<br />
Second and third bonding approaches<br />
1. Spin-coating of SU-8<br />
or lamination of DFR on<br />
second PMMA substrate<br />
2. Bonding partners are<br />
pressed together<br />
8. Removal of protective<br />
foil<br />
Fig. 1. Process flow of the channel layer fabrication including the fluidic<br />
interfaces.<br />
3. UV-Exposure of<br />
SU-8 / DFR through<br />
PMMA lid<br />
Fig. 2. Process flows of the bonding techniques.<br />
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May 2011, Aix-en-Provence, France<br />
In a second approach, we used another PMMA plate as a<br />
<br />
lid substrate. DFR was laminated on this plate and brought a)<br />
into contact with the channel layer. Defined pressure and<br />
temperature were applied by passing this stack through the<br />
desktop laminator again, so that bonding took place. Final<br />
polymerization and bond stabilization was attained by<br />
exposing the DFR to UV-light through the transparent<br />
PMMA lid plate after bonding.<br />
Akin to this, we also used a PMMA lid substrate for the<br />
third bonding technique. Instead of DFR, a 10 µm SU-8<br />
layer was deposited on this lid plate. This adhesive layer<br />
was laid on the channel structure and bonded under certain<br />
pressure and temperature conditions. Again, exposure was<br />
carried out through the PMMA lid substrate after bonding.<br />
A related approach was attempted in [5], but using a predrilled<br />
lid substrate, which led to deposition and bonding<br />
problems as will be described below.<br />
III.<br />
RESULTS AND DISCUSSION<br />
A. Channel Layer Fabrication<br />
Both SU-8 and DFR were found to be suitable as channel<br />
layer materials as clearly defined channels with straight<br />
walls were fabricated. However, DFR allows aspect ratios<br />
up to approximately 2:1, whereas SU-8 can be used for any<br />
channel dimensions up to more than 10:1. Fig. 3 a) shows a<br />
confocal microscope picture of channels with widths<br />
between 1.34 µm and 12 µm and a height of 20 µm. Due to<br />
the weaker optical properties and surface quality of PMMA<br />
compared to glass, scattering of the exposure light was<br />
expected, which would lead to difficulties in producing high<br />
aspect ratios. However this problem held off and even<br />
channels with an aspect ratio higher than 10:1 were<br />
fabricated. Furthermore, the use of PMMA as substrate<br />
material is advantageous compared with the commonly used<br />
silicon or glass due to its coefficient of thermal expansion<br />
(CTE). As mentioned in [5], the CTEs of PMMA and SU-8<br />
(85 ppm/K and 52 ppm/K, respectively) are closer to each<br />
other than SU-8 and silicon (2 ppm/K) or borosilicate glass<br />
(3,25 ppm/K). Similar thermal expansion coefficients<br />
prevent distortion of resist during the baking steps.<br />
Moreover, no flaking of resist from PMMA or from the<br />
adhesive resist layer was visible, not even near the drilled<br />
holes (Fig. 3 b)). Thus, good adhesion of the layers was<br />
provided, although there were no elaborate cleaning steps<br />
necessary prior to deposition but only flushing of the<br />
PMMA plate with IPA and nitrogen.<br />
B. Fluidic Interfaces<br />
Mechanical drilling of the fluidic inlets and outlets<br />
proofed of value as a very simple and efficient technique.<br />
By means of the protective foil, absolutely no damaging of<br />
the channel layer occurred and the foil was removed<br />
without difficulties. The drilled holes were well defined<br />
without chipping and no cracks or other damages were<br />
induced (see Fig. 3 b)).<br />
In Ref. [5], fluidic inlets and outlets were drilled into the<br />
PMMA lid plate instead of the PMMA substrate which<br />
contains the channel structures. SU-8 was not suitable as an<br />
b)<br />
Fig. 3. a) Confocal microscope image of structured SU-8 on PMMA<br />
(height: 20 µm). The black line marks the position of the profile shown<br />
below. b) Confocal microscope image of a channel with a fluidic inlet<br />
drilled into PMMA.<br />
adhesive layer as it formed cords because of the holes.<br />
Thus, DFR was deposited on top of this pre-drilled PMMA<br />
lid to serve as the adhesive bonding layer. When the release<br />
liner of the DFR was removed, most inlets and outlets were<br />
open as the DFR layer could not stick to a surface in these<br />
areas. In doing so, some pieces of DFR fell onto the<br />
deposited adhesive layer where they formed artifacts and<br />
led to bond defects. This problem, that also involves low<br />
yield and reliability, is completely evaded using the new<br />
fabrication process.<br />
C. Closing of Channels<br />
Comparing the adhesion bonding techniques, all three of<br />
them turned out to be successful methods for closing of<br />
channels. On account of exposing the resist after bonding,<br />
good adhesion and bond strengths were assessed. Yet each<br />
275
one of the bonding methods is best adequate for different<br />
ranges of channel dimensions.<br />
Laminating a DFR layer on top of the channel layer was<br />
the most comfortable way of closing since it is the simplest.<br />
As shown in Fig. 4 a), channels were tightly sealed by this<br />
method at an optimum lamination temperature of 75 °C [6].<br />
However, for channels exceeding widths of 250 µm, the lid<br />
DFR layer sagged into the broad chambers and stuck to the<br />
bottom (Fig. 4b)). Therefore, this bonding method is limited<br />
to smaller channel structures. Though, for channels smaller<br />
than 20 µm, unbonded spots occurred at the channel edges.<br />
With respect to biomedical applications, it is desirable for<br />
the biological fluid to be in contact with as few materials as<br />
possible in order to avoid interaction of biological<br />
substances with other materials. On this score, combination<br />
of DFR as the channel layer with lamination of DFR as a lid<br />
forms the simplest way of fabricating a complete<br />
microfluidic system in which the biological fluid is in<br />
contact only with DFR and no other material. Still, as<br />
described above, this proceeding is best applicable for<br />
moderate channel dimensions between 20 µm and 250 µm.<br />
Based on the second bonding approach, the application<br />
range was extended to smaller and larger channel<br />
dimensions. As DFR was laminated onto a PMMA lid plate<br />
at first, sagging of DFR into broad channels or chambers<br />
was obviated entirely. This bonding technique did not<br />
involve any constraints for dimensions of the channels to be<br />
covered. However, if the whole system is supposed to<br />
consist of DFR, channel widths are restrained to aspect<br />
ratios lower than 2:1. Alternatively, smaller channels<br />
fabricated from SU-8 can also be sealed by DFR resulting in<br />
an equally stable bond. Although two materials (SU-8 and<br />
DFR) will be in direct contact with the biological<br />
substances, these materials are chemically very similar.<br />
Yet, when SU-8 is chosen as the channel layer, SU-8 can<br />
also be employed as adhesive bonding layer using the third<br />
bonding approach. This method also revealed stable and<br />
homogenous bonds for any channel dimensions. Fig. 5<br />
depicts a microfluidic system covered by this technique.<br />
Obviously, the channels are open and the bond is<br />
homogenous without any air entrapments or other defects.<br />
Rhodamine was used as a test liquid for checking the leaktightness<br />
of the systems. No bonding defects could be<br />
observed as the liquid filled the channels completely but did<br />
not flow between the resist layers. Admittedly, this process<br />
turned out more sensitive against process parameter<br />
deviations than DFR bonding. For example, marginally<br />
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May 2011, Aix-en-Provence, France<br />
<br />
exceeding the optimal temperature for bonding (69 °C)<br />
about 1-2 °C already led to flowing of resist into the<br />
channels, which resulted in clogging. At slightly lower<br />
temperatures, bond defects were found occasionally,<br />
especially in regions of small bonding areas. Totaling, this<br />
fabrication technique is suitable for any channel designs<br />
when process parameters are set accurately.<br />
Generally, this bonding method can be accomplished<br />
with any adhesive material. In comparison with the<br />
technique of Ref. [5], for which only dry film is applicable,<br />
the new fabrication process turns out more flexible with<br />
regard to material choice and combinations. For this reason,<br />
it is adaptive to a broader range of applications.<br />
IV. CONCLUSION<br />
All things considered, the presented manufacturing<br />
options base upon a combination of the polymer materials<br />
PMMA, SU-8 and DFR. The fabrication methods stand out<br />
due to inexpensive materials and manufacturing techniques<br />
compared to conventional silicon and glass assemblies.<br />
Besides the low costs, an eminent benefit is also given by<br />
the transparency of the materials as observability of<br />
processes is a crucial requirement for many biomedical<br />
applications.<br />
Although polymers have poor temperature stability, this<br />
fact is not of disadvantage in the biological field where low<br />
temperature processes are needed to prevent denaturing or<br />
alike damages of biological substances. In comparison with<br />
PDMS systems, the presented techniques are suitable for a<br />
wider range of applications as they allow the fabrication of<br />
smaller channels with higher aspect ratios.<br />
The bonding techniques could also be adapted for CMOScompatible<br />
encapsulation of micromechanical devices such<br />
as switches. For these applications, high temperature<br />
bonding techniques like anodic bonding are often<br />
inappropriate as they can cause thermal distortion<br />
discharging of structural elements.<br />
Summing up we highlighted complete fabrication<br />
techniques for microfluidic systems suitable for a large<br />
variety of biomedical applications. A great advantage is<br />
given by implementing the fluidic interfaces simply by<br />
CNC-assisted mechanical drilling. Very high aspect ratios<br />
(>10:1) were achieved and three different adhesion bonding<br />
techniques for closing of the channels were applied and<br />
compared. These bonding methods cover all dimension<br />
ranges of channels or chambers to be sealed.<br />
a) b)<br />
Fig. 4. Channels on PMMA closed by lamination of DFR. a) Cross-section<br />
of a 220 µm wide channel [6], b) Top view of a 1000 µm wide channel.<br />
Fig. 5. SU-8-channel on PMMA closed by bonding to a SU-8-PMMA lid.<br />
For testing the leak-tightness of channel, rhodamine was flown through.<br />
276
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[1] G. Mehta, J. Lee, W. Cha, Y.-C. Tung, J.J. Linderman and S.<br />
Takayama, “Hard top soft bottom microfluidic devices for cell<br />
culture and chemical analysis”, Analytical Chemistry, 81(10),<br />
2009, 3714-3722.<br />
[2] H. Lee, E. Sun, D. Ham and R. Weissleder, “Chip-NMR biosensor<br />
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[3] H. Sugino, K. Ozaki, Y. Shirasaki, T. Arakawa, S. Shoji and T.<br />
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[4] C.-Y. Lee, G.-B. Lee, J.-L. Lin, F.-C. Huang and C.-S. Liao,<br />
“Integrated microfluidic systems for cell lysis, mixing/pumping<br />
and DNA amplification”, Journal of Micromechanics and<br />
Microengineering, 15, 2005, 1215-1223.<br />
[5] A. Krebs, T. Knoll, D. Nußbaum and T. Velten, “Fabrication of<br />
enclosed SU-8 microchannels for cell handling applications”, Proc.<br />
10th Int. Conf. on Management of Innovative Technologies, Fiesa,<br />
Slovenia, 2009.<br />
[6] D. Nußbaum, D. Herrmann, T. Knoll and T. Velten, “Micromixing<br />
Structures for Lab-on Chip Applications: Fabrication and<br />
Simulation of 90° Zigzag Microchannels in Dry Film Resist”,<br />
Proc. 4M/ICOMM Conference, Karlsruhe, Germany, 2009.<br />
[7] G. Voskerician, M. S. Shive, R. S. Shawgo, H. von Recum, J. M.<br />
Anderson, M. J. Cima and R. Langer, “Biocompatibility and<br />
biofouling of MEMS drug delivery devices”, Biomaterials, 24,<br />
2003, 1959-1967.<br />
[8] G. Kotzar, M. Freas, P. Abel, A. Fleischman, S. Roy, C. Zorman, J.<br />
M. Moran and J. Melzak, “Evaluation of MEMS materials of<br />
construction for implantable medical devices”, Biomaterials, 23,<br />
2002, 2737-2750.<br />
[9] A. Altuna, G. Gabriel, L. M. de la Prida, M. Tijero, A. Guimerá, J.<br />
Berganzo, R. Salido, R. Villa and L. J. Fernández, “SU-8-based<br />
microneedles for in vitro neural applications”, Journal of<br />
Micromechanics and Microengineering, 20, 2010, 064014 (6pp).<br />
[10] L. Bogunovic, D. Anselmetti and J. Regtmeier, “Photolithographic<br />
fabrication of arbitrarily shaped SU-8 microparticles without<br />
sacrificial release layers”, Journal of Micromechanics and<br />
Microengineering, 21, 2011, 027003 (5pp).<br />
[11] F.J. Blanco, M. Agirregabiria, J. Garcia, J. Berganzo, M. Tijero,<br />
M.T. Arroyo, J.M. Ruano, I. Aramburu and K. Mayora, “Novel<br />
three-dimensional embedded SU-8 microchannels fabricated using<br />
a low temperature full wafer adhesive bonding”, Journal of<br />
Micromechanics and Microengineering, 14(7), 2004, 1047-1056.<br />
[12] S. Tuomikoski and S. Franssila, “Free-standing SU-8 microfluidic<br />
chips by adhesive bonding and release etching”, Sensors and<br />
Actuators A, 120, 2005, 408-415.<br />
[13] P. Vulto, N. Glade, L. Altomare, J. Bablet, L. Del Tin, G. Medoro,<br />
I. Chartier, N. Manaresi, M. Tartagni and R. Guerrierri,<br />
“Microfluidic channel fabrication in dry film resist for production<br />
and prototyping of hybrid chips”, Lab on a Chip, 5, 2005, 158-162.<br />
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277
11-13 May, Aix-en-Provence, France<br />
<br />
Hot embossing of biodegradable polymers<br />
Matthias WORGULL, Alexander KOLEW, Heilig MARKUS, Marc SCHNEIDER, Heinz DINGLREITER<br />
Karlsruher Institute of Technology, Eggenstein-Leopoldshafen, Germany<br />
Text unavailable at the time of printing.<br />
278
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May 2011, Aix-en-Provence, France<br />
<br />
SU-8-based rapid tooling for thermal roll<br />
embossing<br />
Khaled Metwally, Laurent Robert, Roland Salut and Chantal Khan Malek<br />
FEMTO-ST Institute - UMR CNRS 6174 / Dpt. MN2S<br />
32 Avenue de l’Observatoire, 25044 Besançon cedex, France<br />
Abstract- Rapid, flexible and low-cost tooling is particularly<br />
required in replication processes especially with small and<br />
medium volume as in research labs or startups. Epoxy stamps<br />
have been used in hot embossing and injection moulding since<br />
a few years. In this work, SU-8 epoxy–based patterns were<br />
generated on silicon wafers, which were employed as stamps<br />
in hot roll embossing of COC and PMMA foils using a<br />
commercial laminator. This method combines the accuracy of<br />
lithographic patterning of SU-8 resist with the mass<br />
production capability of roll embossing. The stamp fabrication<br />
process can be performed in less than a few hours using<br />
photolithography for tens to hundreds micrometer features<br />
and electron beam lithography for the sub-micronic range.<br />
I. INTRODUCTION<br />
Emerging non-silicon manufacturing technologies<br />
involve large sheets of materials and continuous processes.<br />
This is particularly true for low-cost disposable devices<br />
based on polymers, such as microfluidic systems. One of<br />
the manufacturing options for fabrication of such devices is<br />
using roll-based processes, where rollers are used to apply<br />
heat and pressure between stamp and substrate, while<br />
rotating with certain angular velocity to achieve continuous<br />
or semi-continuous embossing/ imprinting. Roll embossing<br />
becomes a viable fabrication technology as it provides<br />
advantages such as increased patterning speed and large<br />
area processability [1-2].<br />
Two possible configurations are available in roll-based<br />
embossing set-ups [3], roll-on-flat where structures are<br />
patterned on a planar surface used as a stamping tool and<br />
roll-to-roll where features are either directly patterned on<br />
the roller or on a foil which will be wrapped around one of<br />
the rollers to form an exchangeable tool on the embossing<br />
cylinder.<br />
Another upcoming trend concerns developing solutions<br />
based on thin films or foils. The manufacture of lab-on-chip<br />
based on thin film technologies (lab-on-a-foil) [4-5] opens<br />
the door for different applications, particularly those<br />
requiring better heat transfer efficiency, mechanical<br />
flexibility and lower material consumption.<br />
The usage of clean room technologies allows<br />
manufacturing of highly accurate silicon-based stamps that<br />
can be used in hot embossing either planar [6] or roll-based<br />
[7]. The resultant replicas are almost perfect as stamps.<br />
However, due to the different thermal expansion<br />
coefficients of tool and substrate induced thermal stresses<br />
replication errors might occur. As typical values for the<br />
thermal expansion coefficient of polymers are in between<br />
50 and 90 ppm K -1 , while silicon masters have a thermal<br />
expansion coefficient of about 2.6 ppm K -1 .<br />
Use of epoxy-based stamps with thermal expansion<br />
coefficient values much closer to those of the polymer<br />
substrates will overcome these problems. As master and<br />
substrate will shrink at a similar rate and result in less stress<br />
during the cooling phase of a silicon /substrate stack.<br />
Usually, epoxy-based stamps are fabricated by casting for<br />
either one or two-component epoxies. Such stamps can be<br />
used in planar hot embossing or injection moulding [8-10].<br />
They can be also extended for roll embossing as reported by<br />
Velten et al., where a one-component thermo-curable resin<br />
(Hysol 9509) was double casted, firstly from a UVpatterned<br />
SU-8 resist template to a silicone mould then<br />
followed by second casting from the silicone mould to an<br />
epoxy master [11]. This epoxy grade was chosen as it<br />
provided high adherence to metals subjected to high-service<br />
temperatures.<br />
SU-8 is an epoxy-based negative tone photosensitive<br />
resist optimized for ultraviolet photolithography and it is the<br />
most popular resist in microsystem technologies. It is also<br />
known to be highly sensitive to electron beam lithography<br />
(EBL). In particular, patterns down to 250 nm in 150 nm<br />
thick SU-8 were written with 50 keV electron beam<br />
lithography (EBL), at a dose less than 2µC/cm 2 [12]. More<br />
recently, Bilenberg et al. reported line widths down to 24<br />
nm with a pitch of 300 nm in a 99 nm thick SU-8 layer<br />
patterned by 100 keV EBL [13]. SU-8 also allows hybrid<br />
patterning, using a combination of photolithography for<br />
larger feature sizes and EBL for finer ones [14].<br />
SU-8 has been used as a stamp for direct imprinting for<br />
micro-features [15], and for sub-micronic hill-like features<br />
down to 650 nm in thermoplastic polymer [16]. In addition,<br />
SU-8 has a thermal expansion coefficient of 52 ppm K -1 ,<br />
which can minimize thermally induced stresses in the<br />
embossed material as well as replication errors [17].<br />
In this work, multi-scale SU-8–based patterns from<br />
micronic down to sub-micronic were generated and<br />
employed as stamps in hot roll embossing of COC and<br />
PMMA foils using a commercial laminator. Roll-on-flat<br />
configuration with processing parameters such as roller<br />
temperature, applied pressure, roller speed and number of<br />
passes has been investigated.<br />
279
II.<br />
EXPERIMENTAL WORK<br />
A. Micronic SU-8 mould manufacturing<br />
A layer of 70 µm thick SU-8 2075 (MicroChem Corp.)<br />
negative photoresist was spin-coated either on a silicon<br />
wafer or on an aluminum layer sputtered on a silicon wafer,<br />
then soft-baked on a hotplate at 95 °C for 7 min.<br />
Microstructures were formed by UV contact photolithography<br />
using an EVG620 mask aligner (SUSS<br />
MicroTec Corp, Germany) with a power 300 mJ/cm 2 . After<br />
exposure the SU-8 was post-exposure baked (PEB) at 95 °C<br />
for 4 min on a hotplate and subsequently developed for 10<br />
min in the SU-8 developer, that is propylene glycol<br />
monomethyl ether acetate (PGMEA), rinsed in ethanol for 1<br />
min, and dried in nitrogen gas.<br />
B. Submicronic SU-8 mould manufacturing<br />
Submicronic ridges of nominal lateral dimensions ranging<br />
from 300 nm to 5µm, with a fixed pitch of 5 µm were<br />
generated in thinned SU-8 NANO TM SU-8 2002 resist<br />
(MicroChem Corp.) diluted (2:3) with NANO TM SU-8<br />
thinner (cyclopentanone, MicroChem Corp.) to achieve a 1<br />
µm thick film resist. Before the spin-coating of the thinned<br />
SU-8 on silicon wafers, an adhesive promoter (Omnicoat TM ,<br />
MicroChem Corp.) was dispensed, followed by a 100 rpm/s<br />
ramped spread cycle to 500 rpm, and a 300 rpm/s ramp to<br />
reach a final 2000 rpm in 30 s long spin cycle. The<br />
Omnicoat TM was baked for 1 min at 200 °C, followed by<br />
dispensation of thinned SU-8 at a 1000 rpm/s ramped cycle<br />
to reach a final 3000 rpm in 30 s long spin cycle.<br />
After spinning of resist on silicon substrate, the SU-8 was<br />
pre-baked at 95 °C for 40 s on a hotplate, prior to exposure<br />
by 30 keV electron beam lithography. Exposures were<br />
carried out using a Raith E-line pattern generator with doses<br />
ranging from 1 µC/cm 2 to 4 µC/cm 2 with an increment step<br />
of 0.2 µC/cm². The exposed SU-8 was post-baked at 95 °C<br />
for 2 min on a hotplate and subsequently developed in<br />
PGMEA for 10 min, rinsed in ethanol for 1 min, and dried<br />
in nitrogen gas.<br />
Reactive ion etching (RIE) under oxygen plasma with<br />
process parameters of 20 sccm/ 20 µbar/ 50 watt/ 90 sec<br />
was used to remove the Omnicoat TM layer and SU-8<br />
residues caused by proximity effect in EBL. The RIE step<br />
was performed at an etching rate of 60 nm/min.<br />
Finally, stamps with both micronic and submicronic SU-8<br />
structures were ramp hard baked at 200 °C on a hot plate for<br />
1 hr to further cross-link the patterned resist and increase its<br />
adherence to the silicon substrate.<br />
C. Roll embossing process<br />
Roll embossing was conducted using a Rohm & Haas<br />
350HR laminator. It consists of two rollers with metallic<br />
cylinders covered with a rubbery material. The rotating<br />
motor is attached to the lower roller, the speed of rotation<br />
being an adjustable parameter. The upper roller has a degree<br />
of freedom in vertical translation allowing two positions,<br />
separation or contact (with the lower roller). The contact<br />
pressure is adjustable via a relief valve pressurized by air<br />
supply. The upper roller is rotated by transmission during<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
contact with the lower roller. The roller operates within a<br />
temperature range from room temperature to 180 °C.<br />
Software limit of 166 °C was set for the temperature<br />
controllers to avoid heater damage.<br />
Thin films of two different thermoplastic materials were<br />
fed as flexible polymer substrates, cyclic-olefin-copolymer<br />
(COC) (Topas 8007 from Ticona GmbH) and poly-methylmethacrylate<br />
(PMMA) (Goodfellow) of respective nominal<br />
thickness of 130 and 125 µm. Their glass transition<br />
temperature (Tg) is equal to 78 and 105 °C, respectively.<br />
A structured SU-8/Si was used as a stamping tool. It was<br />
placed with its features upwards on a supporting metallic<br />
plate. The polymer/SU-8 stamp assembly was then forced to<br />
pass between both embossing rollers under given embossing<br />
pressure, roller temperatures, and rolling speed.<br />
A multiple-pass process of the polymer foil/SU-8 stamp<br />
assembly through the roller offers several advantages<br />
compared with the single pass process as it allows<br />
embossing at lower temperature, hence decreasing the<br />
thermally induced stress in the polymer film during the<br />
embossing process. However at lower temperature, the<br />
polymer requires more holding time to fill the microcavities<br />
on the stamp. Moreover, the polymer viscoelastic behaviour<br />
resists forming of polymer and exhibits time dependent<br />
strain. Each pass results in not only preheating for next pass<br />
but also in improving the transferred depth in the polymer<br />
foil as the applied pressure and holding time are repeated.<br />
The first pass is a critical pass as it is used to mark the first<br />
imprint on the polymer substrate that will guide the stamp<br />
features during the following passes of embossing, which<br />
avoid the occurrence of predicted misalignment between<br />
different passes. Moreover, multiple passes could be applied<br />
in multi-roller embossing system that will allow full control<br />
of thermal cycle without affecting the production rate.<br />
Several experiments were performed to optimize the<br />
filling of microcavities of the stamp. Pattern transfer in<br />
polymer foils was investigated as a function of roller<br />
temperature, applied pressure, feeding rate, and number of<br />
passes. Both mould and replicas were characterized using<br />
optical microscopy (Leica), profilometry (Alpha-Step 200)<br />
and scanning electron microscopy (SEM) (Leica S-440).<br />
III.<br />
RESULTS AND DISCUSSION<br />
A. SU-8 mould for micronic features and its replicas<br />
Microfeatures of 100 µm width and 80 µm depth were<br />
successfully manufactured in SU-8 with straight sidewalls<br />
as shown in Fig.1 and its inset, on both the silicon wafer and<br />
the sputtered aluminum (Al) layer over silicon wafer. For<br />
the second one, it was planned to check the durability of<br />
SU-8 adherence with Al in order to extend the work to a<br />
roll-to-roll configuration with an aluminum foil wrapped<br />
around the roller.<br />
Aluminum metal was chosen according to a previous<br />
study of Nordström group which compared the bond<br />
strength between SU-8 and four different materials by pulltest<br />
experiments [18], without any adhesion promoter<br />
between the SU-8 and the respective material.<br />
The bond between SU-8 and gold (Au) was the weakest,<br />
280
followed by titanium (Ti) which showed a slight increase in<br />
the bond strength with SU-8. The bond of SU-8 to Al was<br />
even stronger, with a value of 12.1 (±2.8) MPa. The<br />
measured bond strength between SU-8 and Si was 18.5(±<br />
4.6) MPa. However, after less than five replicas the SU-8<br />
was delaminated from the aluminum.<br />
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The PMMA replica shown in Fig.3 was produced in six<br />
passes at a temperature of 166 °C for both upper and lower<br />
rollers while applying a pressure of 6 bars and a feed rate<br />
0.1 m/min. In this particular case, only the number of passes<br />
varies in the optimization process as the other parameters<br />
were physically or software limited.<br />
Fig. 1. SU-8/Si mould with micro-features of 100 µm wide and 80 µm height.<br />
Inset: Magnification view of the ridge section 250X<br />
This low number of replication is not acceptable for a roll<br />
embossing process and SU-8/Al is not durable to be used as<br />
stamp. Delamination is explained by reaching a relatively<br />
high PEB temperature (95 °C, 2 min) without ramping on<br />
two different materials with thermal expansion coefficient<br />
mismatch.<br />
On the another side the SU-8/Si stamp has been used in<br />
roll-to-flat configuration and microfeatures of 100 µm width<br />
and 80 µm depth were successfully replicated in Topas<br />
8007 COC and PMMA. The COC replica shown in Fig.2<br />
was produced in a two-pass process at a temperature of 110<br />
and 145 °C for upper and lower roll respectively, using a<br />
pressure of 6 bars and a feed rate of 0.3 m/min. These<br />
processing parameters show an increase in embossing<br />
temperature and reduction in speed compared with the<br />
parameters used with the silicon moulds [7]. This increase<br />
was expected as the thermal conductivity of SU-8 is less<br />
than that of silicon (on the order of 0.3 vs. 149 W/m.°K,<br />
respectively).<br />
Fig. 3. PMMA replica embossed from SU-8/Si mould with micro-features<br />
of 100 µm wide and 80 µm deep. Inset: Magnification view of the channel<br />
section 250X<br />
B. SU-8 e-beam dose evaluation and patterning of submicronic<br />
features<br />
The e-beam exposure dose variation was used to measure<br />
the resist height versus dose characteristic curve. The<br />
exposure dose was obtained by measuring the current using<br />
a Faraday cage. The remaining SU-8 resist thickness was<br />
measured by a profilometry at the 5µm features after<br />
development and PEB. Fig. 4 shows the remaining SU-8<br />
resist thickness versus exposure.<br />
Normally, SU-8 contrast and sensitivity are determined<br />
for a resist thickness which is half the initial film thickness<br />
(D H=0.5 , which also depends on PEB temperature and<br />
development conditions).<br />
The optimum dose depends on pattern size and period so<br />
that doses were selected for each feature size after SEM<br />
observation to avoid proximity effect and to maintain a<br />
given profile.<br />
1400<br />
Remaining resist thickness vs dose<br />
1200<br />
Remaining resist thickness (nm)<br />
1000<br />
800<br />
600<br />
400<br />
200<br />
Fig. 2. Topas COC 8007 replica embossed from SU-8/Si mould with microfeatures<br />
of 100 µm wide and 80 µm deep. Inset: Magnification view of the<br />
channel section 250X<br />
0<br />
0 0.5 1 1.5 2 2.5 3 3.5 4<br />
Dose (µC/cm2)<br />
Fig. 4. Remaining SU-8 2002 resist thickness (after a 2 min PEB at 95 °C &<br />
10 min development in SU-8 developer) versus e-beam exposure dose at 30<br />
keV, measured by profilometry, for pads of nominal width 5 µm.<br />
281
Sub-micronic features with a final dimension of 400 nm<br />
wide, 570 nm height and 50 µm long were produced in SU-<br />
8 on silicon after exposure, development, PEB and hard<br />
bake as shown in Fig.5. For the 400 nm feature array,<br />
exposure dose of 4 µC/cm² was selected. However, with<br />
larger feature size less exposure dose is required.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
m/min as shown in Fig.7. Reduction of feed rate in PMMA<br />
case shows difference as it is inversely proportional to<br />
holding time of applied pressure and allows polymer filling.<br />
Also, its value directly affects the production rate. Good<br />
pattern transfer is obtained by changing only the number of<br />
passes compared with the micronic features replication.<br />
Fig.5 SU-8/Si mould with submicronic features of 400 nm wide and 570<br />
nm height. Inset: Magnification view of SU-8 feature cross-section 40KX<br />
First set of trials for roll embossing with 400 nm wide<br />
features was conducted in Topas 8007 COC. Embossed fine<br />
features of 400 nm width and 570 nm depth were produced<br />
in three passes at a temperature of 133–137 °C for upper<br />
and lower rolls, using a pressure of 4 bars and a feed rate of<br />
0.4 m/min as shown in Fig.6. However, an incomplete<br />
filling problem at sharp edges appears (see inset of Fig.6).<br />
Fig.7. PMMA replica embossed from SU-8/Si mould with submicronic<br />
features of 400 nm wide and 570 nm deep. Inset: Magnification view of the<br />
channel section Mag. 40KX<br />
C. Mould defects after replication<br />
The SU-8/Si mould has been used for roll embossing in<br />
COC foils without facing any problem. However, after<br />
using the SU-8/Si mould to emboss in PMMA, a complete<br />
de-lamination at the 300 nm features and partial delamination<br />
at the 400 nm features were observed on the<br />
stamp as shown in Fig.8 and its inset.<br />
Fig.6. Topas COC 8007 replica embossed from SU-8/Si mould with<br />
submicronic-features of 400 nm wide and 570 nm deep. Inset:<br />
Magnification view of the channel section Mag. 40KX<br />
It was expected that by increasing the embossing<br />
temperature and pressure, as well as decreasing feed rate,<br />
this problem could be resolved. But comparing these results<br />
with another COC embossed replica produced in three<br />
passes at the higher temperature of 166 °C for both rollers<br />
and a higher pressure of 6 bars and a feed rate of 0.3 m/min,<br />
shows no big difference at the sharp edges.<br />
Submicronic features replicated in PMMA in four passes<br />
at a temperature of 166 °C for both upper and lower rollers<br />
while applying a pressure of 6 bars and a feed rate of 0.1<br />
Fig.8. SU-8/Si mould with submicron features of 400 nm wide and 570 nm<br />
height after embossing. Inset: Magnification view of feature cross-section<br />
Mag. 40KX<br />
This de-lamination could be explained by the relatively<br />
high adhesive forces between the master patterns in SU-8<br />
and the PMMA foil to be embossed. These adhesive forces<br />
could be reduced if an anti-sticking layer is deposited before<br />
starting the embossing process. This anti-sticking layer<br />
could be a deposited layer of Teflon-like octo-fluoro-butane<br />
(C 4 F 8 ) gas.<br />
Also, the adherence between patterned SU-8 resist and<br />
the silicon wafer could be increased by optimizing the<br />
282
exposure dose, post exposure bake temperature and time for<br />
each set of feature size. Indeed Hong et al. reported in their<br />
study of the significant fabrication parameters associated<br />
with the de-lamination of 100µm thick SU-8 film from a<br />
silicon wafer substrate and their effect using a neural<br />
network model. The SU-8 layer had been blank-exposed by<br />
photolithography and they showed that a higher exposure<br />
dose lowers the temperature at which de-lamination starts to<br />
occur, increasing de-lamination [19]. It is consistent with<br />
our observation that the smaller feature sizes on the SU-8<br />
stamp, that is 300 nm, which require a higher electron beam<br />
exposure dose, are also the first ones to be delaminated.<br />
With larger feature size obtained by photolithography, no<br />
delamination was visible on COC films, but it occurs with<br />
PMMA films.<br />
IV. CONCLUSION AND PERSPECTIVE<br />
The feasibility of using SU-8 epoxy stamps has been<br />
demonstrated, as well as a high-speed fabrication process of<br />
large area micronic and sub-micronic array of lines suitable<br />
for small and medium mass production.<br />
Micro-features of 100 µm width and 80 µm depth were<br />
successfully replicated in 130µm thick Topas 8007 COC<br />
foils in a two-pass process at a temperature of 110 and 145<br />
°C for upper and lower roll respectively, using a pressure of<br />
6 bars and a feed rate of 0.3 m/min. The replica was<br />
produced in 125µm thick PMMA foils in six passes at a<br />
temperature of 166 °C for both upper and lower rollers<br />
while applying a pressure of 6 bars and a feed rate 0.1<br />
m/min.<br />
Sub-micronic features of 400 nm width and 570 nm depth<br />
were produced in COC in three passes at a temperature of<br />
133–137 °C for upper and lower rolls, using a pressure of 4<br />
bars and a feed rate of 0.4 m/min, and in PMMA in four<br />
passes using the same parameters for the embossing process<br />
as those for embossing 100 µm wide features.<br />
The SU-8 stamp fabrication process was performed in<br />
less than a few hours using photolithography for tens to<br />
hundreds micrometer features and electron beam<br />
lithography for the sub-micronic range down to 300nm. It<br />
also gives the possibility of hybrid manufacturing of both<br />
micronic and submicronic features on the same stamp.<br />
The high resolution capability of SU-8 combined with the<br />
high production rate of roll embossing technique are highly<br />
promising for a wide range of applications.<br />
ACKNOWLEDGMENT<br />
This work was performed within the framework of the<br />
Carnot-Fraunhofer French-German project “3μP: Microfluidic<br />
platform for multiple samples with multiple analytics<br />
to run diagnostic analysis” and the French FUI CONPROMI<br />
project.<br />
REFERENCES<br />
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[4] S. Miserere, J. Weber, B. De Lambert, JL. Viovy, L. Malaquin, “A<br />
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[5] M. Focke, D. Kosse, C. Müller, H. Reinecke, R. Zengerle and F.<br />
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[6] H. Becker and U. Heim, “Hot embossing as a method for the<br />
fabrication of polymer high aspect ratio structures", Sensors and<br />
Actuators, 83, pp.130–135, 2000.<br />
[7] K. Metwally, S. Queste, L. Robert, R; Salut, C. Khan-Malek, “Hot<br />
roll embossing in thermoplastic foils using dry-etched silicon<br />
stamp and multiple passes", Microelectronic Engineering, in press<br />
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[8] T. Koerner, L. Brown, R. Xie, R. D. Oleschuk, “Epoxy resins as<br />
stamps for hot embossing of microstructures and microfluidic<br />
channels", Sensors and Actuators B 107, pp. 632–639, 2005.<br />
[9] J. Steigert, S. Haeberle, T. Brenner, C. Müller, C. P. Steinert, P.<br />
Koltay, N. Gottschlich, H. Reinecke, J. Rühe, R. Zengerle and J.<br />
Ducrée, “Rapid prototyping of microfluidic chips in COC", J.<br />
Micromech. Microeng. 17, pp.333–341, 2007.<br />
[10] M. Svoboda, W. Schrott, Z. Slouka, M. Pribyl, D. Snita, “Plastic<br />
microfluidic systems made by imprinting against an epoxy stamp",<br />
Microelectronic Engineering 87, pp.1527–1530, 2010.<br />
[11] T. Velten, F. Bauerfeld, H. Schuck, and T. Knoll , “Low cost<br />
master fabrication for roll-to-roll hot embossing based on eboxy<br />
resin”, 4M2010 Proc. 7th international conference on<br />
Multimaterial micro manufacture- Oyonnax, pp. 135–138, 2010.<br />
[12] A.L. Bogdanov, “Use of SU-8 Negative Photoresist for Optical<br />
Mask Manufacturing", Proc. SPIE -Advances in resist technology<br />
and processing XVII (Santa Clara CA), 3999, pp.1215-1225, 2000.<br />
[13] B. Bilenberg, S. Jacobsen, M.S. Schmidt, L.H.D. Skjolding, P. Shi,<br />
P. Bøggild, J.O. Tegenfeldt, A. Kristensen, “High resolution 100<br />
kV electron beam lithography in SU-8”, Microelectron. Eng. 83,<br />
pp.1609–1612, 2006.<br />
[14] M. Gersborg-Hansen, L.H. Thamdrup, A. Mironov, A. Kristensen,<br />
“Combined electron beam and UV lithography in SU-8”.<br />
Microelectronic Engineering 84, pp. 1058–1061, 2007.<br />
[15] J. Greener, W. Li, J. Ren, D. Voicu, V. Pakharenko, T. Tang and E.<br />
Kumacheva, “Rapid, cost-efficient fabrication of microfluidic<br />
reactors in thermoplastic polymers by combining photolithography<br />
and hot embossing”. Lab Chip, 10, pp 522–524, 2010.<br />
[16] J.K. Chen, F.H. Ko, C.H. Chan, C.F. Huang and F.C. Chang,<br />
“Using imprinting technology to fabricate three-dimensional<br />
devices from moulds of thermosetting polymer patterns”.<br />
Semicond. Sci. Technol. 21, pp.1213–1220, 2006.<br />
[17] M. B. Esch, S. Kapur, G. Irizarry and V. Genova, “Influence of<br />
master fabrication techniques on the characteristics of embossed<br />
microfluidic channels”. Lab Chip. 3, pp.123–127, 2003.<br />
[18] M. Nordström, A. Johansson, E. S. Nogueron, B. Clausen, M.<br />
Calleja , A. Boisen , “Investigation of the bond strength between<br />
the photo-sensitive polymer SU-8 and gold”. Microelectronic<br />
Engineering 78–79, pp.152–157, 2005.<br />
[19] S. J. Hong, S. Choi, Y. Choi, M. Allen, G.S. May,<br />
“Characterization of low-temperature SU-8 photoresist processing<br />
for MEMS applications”. IEEE conference and workshop<br />
ASMC’04, pp.404–408, 2004.<br />
<br />
283
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Success in MEMS, "From DRIE Technology<br />
to Social Innovation"<br />
Susumu Kaminaga,<br />
Sumitomo Precision Products Co., Ltd<br />
I. INTRODUCTION<br />
Over the past 10 years MEMS devices have become more<br />
established in a number of commercial, high volume<br />
applications such as digital projectors, ink jet printers, and<br />
automotive motion sensors. The invention of the mobile<br />
phone has had an enormous effect on society in terms of<br />
interpersonal communications and working life. The demand<br />
for “smarter” consumer products (e.g. smart phones, handheld<br />
tablets, and gaming consoles) in the past 2-3 years has<br />
stimulated significant growth in manufacturing MEMS devices<br />
for mobile applications [1] .<br />
Meanwhile, more MEMS devices are emerging with the<br />
potential for beneficial applications in other areas such as life<br />
sciences and energy conservation. For example, Yole<br />
Développement forecasts, “microsystem technologies market<br />
for healthcare applications will grow from $1.2B in 2009 to<br />
$4.5B in 2015, representing over 1B units per year in 2015” [2] .<br />
II. DEEP REACTIVE ION ETCH<br />
Deep reactive ion etching (DRIE) is a key enabling process<br />
which has been adopted by the MEMS manufacturers for<br />
etching deep, high aspect ratio features into silicon, the most<br />
common material used for MEMS manufacturing. Feature<br />
sizes range from sub-micron to many hundreds of microns.<br />
SPP Process Technology Systems (SPTS) was the first<br />
equipment manufacturer to commercialise DRIE, also named<br />
the “Bosch Process”, over 15 years ago, when the MEMS<br />
industry was in its infancy.<br />
Fig 1 – Schematic diagram illustrating the process steps of the DRIE<br />
process<br />
III. EMERGING MEMS APPLICATIONS<br />
A. Wireless networking<br />
A Wireless Sensor Network (WSN) consists of spatially<br />
distributed autonomous sensors to monitor physical or<br />
environmental conditions, such as temperature, sound,<br />
vibration, pressure, motion or pollutants, and to cooperatively<br />
pass their data through the network to a main location. Some<br />
networks are bi-directional, allowing the system to control the<br />
activity of the sensors. Although the development of wireless<br />
sensor networks was motivated by military applications such<br />
as battlefield surveillance; today such networks are used in<br />
many industrial and consumer application, such as industrial<br />
process monitoring and control, machine health monitoring,<br />
environment and habitat monitoring, home automation, and<br />
traffic control. MEMS sensors can be designed to measure a<br />
wide range of conditions at each node and are small, reliable<br />
and easily integrated into an electronic system. MEMS may<br />
also be used to scavenge energy to re-charge batteries which<br />
maintain power at the individual nodes.<br />
Fig 2 Multi-hopping Ad-hoc Wireless Network concept<br />
(Courtesy of Crossbow/SPP)<br />
288
B. Energy Applications<br />
Society is faced with many issues associated with energy<br />
use and generation. The semiconductor industry is working<br />
hard to find solutions such as photovoltaic/piezoelectric<br />
generation, ensuring power consumption of every device is<br />
minimised, and scavenging energy which would normally be<br />
wasted within a system.<br />
MEMS are used in a number of different ways to either<br />
replace traditional devices with higher energy consumption,<br />
generate power or harvest energy from the surroundings<br />
Micro heat engines for<br />
power generation &<br />
propulsion<br />
Piezo electric and<br />
electromagnetic power<br />
generators<br />
Thermoelectric and<br />
thermophotovoltaic systems<br />
Micro fuel cells and<br />
micro reactors for fuel<br />
processing and power<br />
generation<br />
Micro coolers and other<br />
thermal management<br />
technologies<br />
Energy scavenging for<br />
embedded microsystems<br />
Fig 3 Some examples of applications for MEMS in energy generation,<br />
conservation and scavenging.<br />
A micro fuel cell is a portable power source for low power<br />
electronic devices that converts chemical energy into useable<br />
electrical energy, like a battery. It generates power through the<br />
electrochemical reaction of a fuel in the presence of a catalyst.<br />
Hydrocarbon based fuels have very high energy densities<br />
compared to batteries, and should in theory offer improved<br />
performance.<br />
devices can easily be retrofitted or placed in inaccessible<br />
places.<br />
C. Life Sciences<br />
As society’s average lifespan increases, the market for life<br />
science products is ever-increasing. In 2010, the medical<br />
segment was expected to represent almost 10 percent of the<br />
$29.7 billion global industrial semiconductor market, or $2.9<br />
billion. Medical electronics is the fastest growing segment in<br />
the industrial semiconductor market, with an average growth<br />
rate of 10 percent per year [3] .<br />
MEMS devices are being developed to improve medical<br />
monitoring, diagnosis and patient care. While stringent safety<br />
testing can lengthen time-to-market, this emerging market is<br />
forecast to become a significant contribution to revenue for<br />
MEMS companies over the coming years.<br />
Many researchers are investigating a variety of methods to<br />
fabricate microneedles and actuator/dosing units for drug<br />
delivery applications. The advantages of transdermal (across<br />
skin) drug delivery include the absence of degradation in the<br />
gastrointestinal tract and liver associated with oral delivery,<br />
and the elimination of pain and inconvenience of an<br />
intravenous injection. The small size of the microneedles<br />
means that the outer epidermis layer of the skin can be<br />
penetrated and the drug delivered without reaching the nerves<br />
situated deeper in the dermis layer. This method of drug<br />
delivery could also reduce the amount of biohazardous waste.<br />
Microneedles may also be used in taking samples from the<br />
body for disease detection or monitoring levels of substances<br />
such as glucose. Implantable devices, controlled by wireless<br />
communication, are also being developed which can offer long<br />
term drug delivery.<br />
Fig 4 Fuel cell flow field formed in silicon substrate using SPTS DRIE<br />
(Courtesy of Lawrence Livermore National Laboratory)<br />
Energy harvesting is not a single technology but a broad<br />
spectrum that can be classified by the type of energy used, e.g.<br />
temperature differences, light radiation, electromagnetic fields,<br />
and kinetic energy.<br />
MEMS can be used to harvest the energy from equipment<br />
vibrations which would otherwise be wasted. This<br />
“scavenged” kinetic energy can be transferred into electrical<br />
energy, stored and used to power small wireless devices, for<br />
applications such as monitoring the equipment’s own structural<br />
health and performance. The advantages of these “selfpowered”<br />
systems are lower running costs and, by eliminating<br />
the need for cabling or replacing batteries, the monitoring<br />
Fig5 Array of micro-needles fabricated by QinetiQ, using a combination<br />
of isotropic and anisotropic etching in SPTS DRIE system, plus wet etching<br />
and electrochemical anodization.<br />
In general, the driving factors which promote the use of<br />
micro and nano-scale diagnosis technologies are:<br />
Improved Ease of Use<br />
Smaller, portable system<br />
Improved Performance<br />
Reducing the sensor element to the scale of<br />
the target species provides a higher sensitivity to a<br />
single entity/molecule.<br />
Reduce Costs<br />
Reduced reagent volumes<br />
Reduced time to result due to small volumes<br />
289
Point-of-care diagnostic<br />
Multi-agent detection capability<br />
In the future there is the potential for use in-vitro and invivo<br />
sensors. In vitro types include the analysis of samples<br />
extracted from the body for determining electrochemistry,<br />
blood pressure and temperature, blood glucose, genetics,<br />
immunology, and toxicology. In vivo sensors measure<br />
biological information inside the body and include catheterbased<br />
biosensor arrays, internal imaging systems, online blood<br />
assays, and neural recording arrays.<br />
[2] http://www.imicronews.com/upload/Rapports/Yole_BioMEMS_Report_Oc<br />
tober_2010_flyer_Web.pdf<br />
[3] Databeans 2010 Medical Semiconductors Report<br />
[4] J.Fernando Alfaro et al, Proc IEEE Engineering in<br />
Medicine & Biology 27 th Conf, 2005<br />
Fig 6 Silicon MEMS sensor to measure biomechanical stresses<br />
(Courtesy of Carnegie Mellon University)<br />
Fig 6 shows an implantable CMOS MEMS sensor<br />
fabricated by researchers at Carnegie Mellon University, using<br />
SPTS’ DRIE processing [4] . Such a sensor could be used to<br />
measure biomechanical stresses in situ, monitoring the strength<br />
of regenerating bone or the interfaces between bone and<br />
prosthetic implants.<br />
IV. CONCLUSIONS<br />
Many MEMS devices like air-bag sensors and ink jet heads<br />
now well-established and are being manufactured in high<br />
volumes. DRIE is a key enabling process for manufacturing<br />
most silicon-based MEMS. Over the past 15 years DRIE<br />
process capability has been continuously improved and<br />
developed, to meet the needs of MEMS manufacturers.<br />
MEMS devices are now finding new, emerging<br />
applications which promise to improve our lives in many other<br />
ways such as environmental monitoring and control, energy<br />
conservation and harvesting, life sciences and security.<br />
V. ACKNOWLEDGEMENTS<br />
The author would like to thank Carolyn Short, Evelyn Tay<br />
and David Butler of SPP Process Technology Systems (SPTS)<br />
for their assistance in the writing of this paper.<br />
REFERENCES<br />
[1] http://www.isuppli.com/MEMS-and-<br />
Sensors/MarketWatch/Pages/New-Consumer-and-Mobile-<br />
MEMS-to-Post-Spectacular-157-Percent-Growth-in-2011.aspx<br />
290
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Brightness Enhancement of OLEDs by Using<br />
Microlens Array Film with Silicon Oil and Ag<br />
Particles<br />
Shan-Shan Hsu 1 , Tung-Yu Chang 1 , Hsiharng Yang 1 , Jen-Sung Hsu 2<br />
1 Institute of Precision Engineering, National Chung Hsing University, Taichung, Taiwan 402<br />
2 Chemical Systems Research Division, Chung-Shan Institute of Science & Technology, Tao-Yuan, Taiwan 325<br />
Abstract- This paper introduces a new method to<br />
improve the external quantum efficiency of organic<br />
light emitted diode (OLED) by adding Ag particle into<br />
the silicon oil layer in OLED. By measuring the<br />
brightness of the front luminance, when the lens is 11μm<br />
in height and 30μm in diameter, the brightnes is<br />
increased 45% by adding the silicon oil between OLED<br />
module and brightness enhancement film, and the<br />
brightness could be further improved to 61% by adding<br />
the Ag particle into the silicon oil. In this work, the<br />
optical waveguide was disturbed by adding Ag particle<br />
into silicon oil and let more lights emit from the OLED<br />
component, it results in higher external quantum<br />
efficiency and lower energy consumption.<br />
I. INTRODUCTION<br />
Various light sources play important roles in recent<br />
displays. Especially for those consumers’electronic devices,<br />
low power consumption and light weight are required.<br />
Conventional CRT (cathode ray tube) television has been<br />
replaced by TFT LCD (thin film technology liquid crystal<br />
display). Obviously, backlighting modules using LEDs<br />
(light emitted diodes) have replaced CRT as the main stream.<br />
However, LED is a point light source, it requires a light<br />
guide plate and other optical films to achieve a lighting plane<br />
for the display. The plane lighting technology may need to<br />
develop for the displays. In 1987, Tang and VanSlyke using<br />
vacuum deposition method to produce the organic light<br />
emitted diode (OLED), the research of OLED grow up<br />
rapidly since that [1]. Due to OLED fit the requirement of<br />
energy saving, high uniformity, flatness, and large size<br />
capability, also it has the benefits such as simple<br />
manufacture process, self -illumination, short response time,<br />
no perspective limit, high contrast ,and low driving voltage,<br />
OLED draw a lot of attention for the next generation display<br />
device. The most important reason people interest in OLED<br />
is the capability of using flexible base plate. Therefore,<br />
OLED not only have the potential to be the illuminate<br />
component for next generation, it also become the major<br />
competitor of the flat display technique in the future [2-4].<br />
The external quantum efficiency is the ratio of the light<br />
generated by the illuminate component and the external light.<br />
The quantum efficiency is usually improved by disturbing<br />
the optical waveguide. Kwon et al., 2008 [5], successfully<br />
fabricates a high-sag microlens array film with a full fill<br />
factor by a simple micromachining process including trench<br />
formation and the conformal vapor phase deposition of a<br />
polymer. According to the fabrication design process, the<br />
lens diameter increased from 12 μm to 17.32 μm, the<br />
required initial radius of curvature of the reflowed<br />
photoresist pattern is 6 μm, and both the final radius<br />
curvature and the sag height after the conformal gap-filling<br />
proces are 8.66 μm. The normal brightnes of the ful white<br />
light emitted from the test panel was measured by PR-650<br />
spectra colorimeter, the result shows the luminance was<br />
increased by up to 48%.<br />
Wei et al., 2006 [6], study the influences of the edge<br />
length and the gap of the microlens array on the luminance<br />
efficiency on OLED. They use the photolithography and hot<br />
melt process to transform these base plates into the shape of<br />
the microlens array on substrate. The duplicated microlens<br />
array adhering to the PMMA film was formed after<br />
separating the mold from the film. The luminance efficiency<br />
of OLEDs has been found to increase linearly with the<br />
increasing area ratio of the microlens base area to the device<br />
area and microlens number density. The luminance<br />
efficiency measured by CS-100 spectra colorimeter was<br />
increase by up to 55%.<br />
Wei et al., 2006 [7], analyze the influences of the fill<br />
factor and the sag of hexagon-based microlens on the optical<br />
characteristics of OLED device. Compared to OLED, the<br />
luminous current and power efficiency of the device can be<br />
enhanced by 35% and 40%, respectively, by attaching a<br />
microlens array having a fill factor of 0.90 and a height ratio<br />
of 0.56. The result shows the efficiency increased as the lens<br />
size decreased and the height ratio increased.<br />
Möller and Forrest [8] demonstrate that ordered<br />
microlens arays with 10μm diameter siloxane lenses<br />
attached to glass substrates increase the light output of<br />
OLED by a factor of 1.5 over unlensed substrates. Peng et.<br />
al., 2004 [9], employed a simple soft-lithography approach<br />
to fabricate the microlens array on glass substrates. With the<br />
use of an optimized lens pattern, an increase of 70%<br />
efficiency in the coupling efficiency is achieved. Lee et al.,<br />
2003 [10], introduced a photonic crystal pattern into the<br />
glass substrate of an OLED. The finite-difference<br />
time-domain method was used to optimize the structural<br />
parameters of the photonic crystal pattern. With the use of an<br />
optimized photonic crystal pattern, an increase in the<br />
extraction efficiency of over 50% was achieved<br />
experimentally.<br />
294
Yamasaki et al., 2000 [11], consisting hexagonally<br />
closed-packed arrays of silica microspheres with the<br />
diameter of 550 nm, were incorporated into OLED with a<br />
conventional two-layer structure to enhance the external<br />
quantum efficiency. Tsutsui et al., 2000 [12], incorporated<br />
closed-packed arrays of silica microspheres with the<br />
diameter of 550 nm into organic light-emitting devices,<br />
increase the external quantum efficiency of 80%. Hobson et<br />
al. 2002 [13], apply wavy component surface cause surface<br />
plasmon resonance let light deliver out of the component.<br />
From these references, the external quantum efficiency of<br />
OLED could be increased by disturbing optical waveguide<br />
such as adding high packing ratio array lens brightness<br />
enhancement film. In this work, a new method of improve<br />
the external quantum efficiency of OLED by adding Ag<br />
particle into the silicon oil layer were applied.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
electron injection layer (EIL) Alq3 (700 Å), electron<br />
transport layer (ETL) LiF (12 Å), and Cathode Al (1500 Å).<br />
Pack under nitrogen atmosphere to finish green light OLED<br />
(Fig 3).<br />
II. EXPERIEMNTS<br />
The optical lithography of like-LIGA technique is used<br />
to manufacture the micro-scale array lens brightness<br />
enhancement film, combined with Ag particle adding silicon<br />
oil, form a low optical waveguide structure. The processes<br />
include optical lithography, the fabrication of OLED with<br />
Ag particle adding silicon oil, micro-structure measurement<br />
and optical brightness measurement.<br />
2.1 Device fabrication<br />
The upper and lower rows of apertures were arranged<br />
equidistantly. Round patterns laid out in an ortho-triangle on<br />
the PET-based mask, to provide the cylinder of photo<br />
resistor for further proces. The pore size is 30μm, gap is<br />
50μm (Fig 1). The experiment parameter shows in Table 1.<br />
The base plate was washed by acetone for cleaning the oil<br />
and dust on the surface, further washed by DI water then dry<br />
by nitrogen gas. The optical lithography processes include<br />
photo resistor coating, soft bake, exposure, and hot process.<br />
The AZ4620 photo resistor is spread on base plate through<br />
two stage spin coating. The purpose of first stage (300 rpm,<br />
10 sec) coating is spraying the photo resistor on the base<br />
plate. The thickness of photo resistor is controlled by the<br />
second stage spin coating (1500 rpm, 25 sec). After soft bake<br />
(90 ℃ , 3min) and exposure (15sec), the base plate was put<br />
into AZ 300MIF developer for several minutes, then the<br />
cylinder structure was formed on the plate surface. In hot<br />
process (160 ℃ , 10min), the kinetic energy of photo resistor<br />
molecular increased due to the temperature raise, with the<br />
effect of surface tension, the photo resistor will form the<br />
shape similar to spherical surface.<br />
Mixing PDMS and hardener at the ratio 10:1 then spray it<br />
on the plate surface, hard baking at 60C for 4 hr. Detach<br />
PDMS from the base plate, then get PDMS plate with indent<br />
array lens structure on it. The UV-curing PMMA was coated<br />
between PDMS plate and PET plate. Exposure this<br />
combined plate under UV light for curing the PMMA.<br />
Detach PDMS mold from this combined plate, attach this<br />
plate on the OLED device surface by the Ag particle adding<br />
silicon oil. The concentration of Ag particle is 0.5%, the<br />
particle size is up to 0.5 μm. The fabrication of OLED is first<br />
etching desired figure on ITO glass plate, then evaporation<br />
coating in order of hole transport layer (HTL) NPb (500 Å),<br />
Fig. 1 Design pattern layout on the PET mask. (unit:μm).<br />
Table 1 Experimental parameters in lithography process.<br />
Base plate clean H 2 SO 4 :H 2 O 2 =3:1 wash<br />
Acetone:60 min<br />
DI Water wash,N 2 dry<br />
120℃bake 20 min dry<br />
Spin coating Spread : 500 rpm 10 sec<br />
Spin : 2000-800 rpm 30 sec<br />
Soft bake 90℃ 3min<br />
Hold 5 min<br />
Exposure 350W, Near UV<br />
600mJ/cm 2<br />
Develop<br />
3.5min<br />
Hot process 160℃ , 30 min<br />
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<br />
Fig. 4. Optical performance measurement setup for OLED brightness.<br />
Fig. 2. The process flow to enhance the OLEDs brightness.<br />
Fig. 3. The structure of green light OLED.<br />
2.2 Optical performance measurement<br />
The micro-structure size of photo resistor was measured<br />
by optical microscope NanoFocus 3D confocal surface<br />
measurement system. The measurement of optical<br />
brightness shows in Fig 4 and Fig 5. The brightness<br />
enhancement film was apply on the OLED modulus and<br />
fixed by a frame. The PR-650 was used to measure the<br />
brightness of OLED with nothing, with silicon oil, and with<br />
silicon oil and Ag particles between OLED and brightness<br />
enhancement film.<br />
Fig. 5. Photograph of the OLED brightness measurement.<br />
III. RESULTS AND DISCUSSION<br />
3.1 Fabrication results<br />
The measurement of photo resistor structure by<br />
NanoFocus 3D confocal surface measurement system was<br />
showsinFig6andFig7.The result shows the lens is 30μm<br />
in diameter and 11 μm in height, the gap between lenses is<br />
50μm.<br />
Fig. 6. Photograph of the fabricated microlens array in photoresist.<br />
296
Fig. 7. Structural profile measurement by using NanoFocus 3D confocal<br />
microscope.<br />
3.2 Optical performance measurement<br />
The performance of OLED with varies fill-factor shows in<br />
Table 2 and Fig 8. Borrelli [14] have mentioned that the lens<br />
efficiency affect by the lens shape and fill-factor, the higher<br />
fill-factor lead to higher lens efficiency. The fill-factor could<br />
be defined as function 3-1.<br />
Lens Area<br />
fill - factor (%)= 100%<br />
(3-1)<br />
Base Area<br />
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<br />
and Fig 9. Without any medium between OLED and<br />
brightness enhancement film, the brightness increased by up to<br />
35% when the lens is 11μm in height. When use silicon oil as<br />
the medium, the brightness increased by up to 45% while the<br />
lens height is 11μm. When add sliver particle into silicon oil,<br />
the brightness increased by up to 61% while the lens height is<br />
11μm. From the results listed in Table 3, the optimize lens size<br />
is 30μm in diameter, 50μm in gap, and 11μm in height.<br />
Fig 10 (a) and (b) shows the brightness measurement of<br />
OLED with and without brightness film, from the result, the<br />
brightness significantly increased when add the enhancement<br />
film. The effect of lens shape also been studied in this work.<br />
When the lens shape is square, the length is 30μm and the gap<br />
is 50μm, the results of brightness measurement are shows in<br />
Table 4 to 6, and Fig 11 to 13. From the results, the square lens<br />
with no medium increase the OLED brightness by 31% when<br />
lens height is 11μm. With silicon oil medium, the brightness<br />
increased by up to 42%; with Ag silicon oil, the brightness<br />
increased by up to 59%. These results shows optimize square<br />
lens size is 30μm in length, 11μm in height and 50μm in gap.<br />
The square lens increase OLED brightness by up to 59%, but<br />
the overall performance of square lens is not as good as<br />
circular lens.<br />
In this work, the circular lenses are arranged in hexagonal<br />
array. From the results shows in Table 2 and Fig 8, when the<br />
lens gap decreased from 100μm to 50μm, the fill-factor<br />
increased from 7.1% to 28.3%, the brightness enhancement<br />
ratio increased from 22% to 40%. From the results above,<br />
under the same lens height, higher fill-factor leads to higher<br />
optical efficiency.<br />
Lens gap<br />
(μm)<br />
Table 2 The OLED brightness with varies fill-factor.<br />
Fill-fac<br />
tor<br />
(%)<br />
Original<br />
brightness<br />
(cd/m 2 )<br />
Brightness<br />
with lens<br />
cd/m 2 )<br />
Increase<br />
amount<br />
(cd/m 2 )<br />
Increase<br />
ratio<br />
(%)<br />
50 28.3 2348 2935 587 25<br />
70 14.4 2299 2713 414 18<br />
100 7.1 2249 2406 157 7<br />
Lens<br />
height<br />
(μm)<br />
Table 3 The OLED brightness with varies lens height.<br />
Medium<br />
Original<br />
brightnes<br />
s(cd/m 2 )<br />
Brightness<br />
with lens<br />
cd/m 2 )<br />
Increase<br />
amount<br />
(cd/m 2 )<br />
Increase<br />
ratio<br />
(%)<br />
5.6 None 2345 2626 281 12<br />
10.5 None 2288 3088 800 35<br />
15.4 None 2248 2720 472 21<br />
5.6 Silicon oil 2376 2803 427 18<br />
10.5 Silicon oil 2300 3335 1035 45<br />
15.4 Silicon oil 2279 3145 866 38<br />
5.6 Silicon oil<br />
with Ag 2321 31180 789 34<br />
particle<br />
10.5 Silicon oil<br />
with Ag 2309 3717 1408 61<br />
particle<br />
15.4 Silicon oil<br />
with Ag<br />
particle<br />
2349 3547<br />
1198<br />
51<br />
Fig. 8. The OLED brightness with varies fill-factor.<br />
The sliver particle and silicon oil was added into OLED<br />
modulus to improve the OLED efficiency. When the lens is<br />
30μm in diameter and 50μm in gap, measuring the front<br />
luminance with varies lens height, the result shows in Table 3<br />
Fig. 9. The OLED brightness with varies lens height.<br />
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15.4 Square 2296 3008 712 31<br />
10.5 Circular 2279 3669 1390 61<br />
15.4 Circular 2343 3538 1195 51<br />
(a)<br />
Fig. 11. The brightness of OLED square/circular lens and no medium.<br />
(b)<br />
Fig. 10. Comparison of the brightness of the same OLED panel (a) with and<br />
(b) without the MLA film.<br />
Table 4 The brightness of OLED with square/circular lens and no medium.<br />
Fig. 12. The brightness of OLED with square/circular lens and silicon oil<br />
medium.<br />
Lens<br />
height<br />
(μm)<br />
Shape<br />
Original<br />
brightness<br />
(cd/m 2 )<br />
Brightness<br />
with lens<br />
(cd/m 2 )<br />
Increase<br />
amount<br />
(cd/m 2 )<br />
Increase<br />
ratio<br />
(%)<br />
10.5 Square 2336 3060 724 31<br />
15.4 Square 2311 2588 277 12<br />
10.5 Circular 2284 3106 822 36<br />
15.4 Circular 2309 2956 647 28<br />
Table 5 The brightness of OLED with square/circular lens and silicon oil<br />
medium.<br />
Lens<br />
height<br />
(μm)<br />
Shape<br />
Original<br />
brightness<br />
(cd/m 2 )<br />
Brightness<br />
with lens<br />
(cd/m 2 )<br />
Increase<br />
amount<br />
(cd/m 2 )<br />
Increase<br />
ratio<br />
(%)<br />
10.5 Square 2299 3265 966 42<br />
15.4 Square 2311 2773 462 20<br />
10.5 Circular 2303 3339 1036 45<br />
15.4 Circular 2323 3206 883 38<br />
Table 6 The brightness of OLED with square/circular lens and Ag silicon oil<br />
medium.<br />
Lens<br />
height<br />
(μm)<br />
Shape<br />
Original<br />
brightness<br />
(cd/m 2 )<br />
Brightness<br />
with lens<br />
(cd/m 2 )<br />
Increase<br />
amount<br />
(cd/m 2 )<br />
Increase<br />
ratio<br />
(%)<br />
10.5 Square 2332 3707 1375 59<br />
Fig. 13. The brightness of OLED with square/circular lens and Ag silicon oil<br />
medium.<br />
IV. CONCLUSION<br />
A new method to improve the external quantum efficiency<br />
of OLED has been developed. The result shows the brightness<br />
affect by the medium in OLED. When sliver particle add into<br />
silicon oil in OLED, the brightness increase significantly.<br />
When the lens is 11μm in height, 30μm in diameter, and<br />
50μm in gap, add sliver particle into silicon oil increase<br />
OLED brightness by up to 61%.<br />
298
ACKNOWLEDGMENT<br />
This work was supported by the National Science Council<br />
(series no. NSC98-2221-E-005-058-MY3) and Chung-Shan<br />
Institute of Science & Technology in Taiwan. The OLED<br />
device fabricated by Professor Fuh-Shyang Juang of National<br />
Formosa University is acknowledged.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
REFERENCES<br />
[1] C.-W. Tang and S.-A. VanSlyke, “Organic electroluminescent<br />
diodes”,Appl. Phys. Lett., Vol. 51, No. 12, pp. 913-915, 1987.<br />
[2] N.-C. Greenham, R.-H. Friend and D.-D.-CBradley, “Angular<br />
dependence of the emission from a conjugated polymer<br />
light-emiting diode: implications for eficiency calculations”, Adv.<br />
Mater., Vol. 6, No. 6, pp. 491-494, 1994.<br />
[3] C.-F. Madigan, M.-H. Lu and J.-C.Sturm, “Improvement of output<br />
coupling efficiency of organic light-emitting diodes by backside<br />
substrate modification”, Appl. Phys. Lett., Vol. 76, No. 13, pp.<br />
1650-1652, 2000.<br />
[4] M.-H. Lu and J.-C.Sturm, “External coupling eficiency in planar<br />
organic light-emiting devices”, Appl. Phys. Lett., Vol. 78, No. 13,<br />
pp. 1927-1929, 2001.<br />
[5] H. Kwon, Y. Yee, C.-H. Jeong, H.-J. Nam and J.-U.Bu, “A high-sag<br />
microlens array film with a full fill factor and its application to<br />
organic light emiting diodes”, J. Micromech. Microeng., Vol. 18,<br />
pp. 1-6, 2008.<br />
[6] M.-K. Wei, I.-L. Su, Y.-J. Chen, M. Chang, H.-Y. Lin, H.-Y. and<br />
T.-C.Wu, “The influence of a microlens aray on planar organic<br />
light-emiting devices,”J. Micromech. Microeng., Vol. 16, pp.<br />
368-374, 2006.<br />
[7] M.-K. Wei, J.-H. Lee, H.-Y. Lin, Y.-H. Ho, K.-Y. Chen, C.-C. Lin,<br />
C.-F. Wu, J.-H. Tsai and T.-C.Wu, “Eficiency improvement and<br />
spectral shift of an organic light-emitting device by attaching a<br />
hexagon-based microlens aray,”J. Opt., Vol. 10, pp. 1-9, 2008.<br />
[8] S. Möller and S.-R.Forest, “Improved light out-coupling in<br />
organic light emiting diodes employing ordered microlens arays”,<br />
J. Appl. Phys., Vol. 91, No. 5, pp. 3324-3327, 2002.<br />
[9] H. Peng, Y.-L. Ho, X.-J. Yu and M.-W.Wong, “Coupling efficiency<br />
enhancement in organic light-emitting devices using microlens<br />
array-theory and experiment” J. Display Technol., Vol. 1, No. 2,<br />
pp.278~282, 2005.<br />
[10] Y.-J. Lee, S.-H. Kim, H. Joon, G.-H. Kim, Y.-H. Lee, S.-H. Cho,<br />
Y.-C. Kim and Y.-R. Do, “A high-extraction-efficiency<br />
nanopatterned organic light-emiting diode”, Appl. Phys. Lett., Vol.<br />
82, No. 21, pp.3779-3781, 2003.<br />
[11] T. Yamasaki, K. Sumioka and T.Tsutsui, “Organic light-emitting<br />
device with an ordered monolayer of silica microspheres as a<br />
scatering medium”,Appl. Phys. Lett., Vol. 76, No. 21,<br />
pp.1243-1245, 2000.<br />
[12] T. Tsutsui, M. Yahiro, H. Yokogawa, K. Kawano and M. Yokoyama,<br />
“Doubling coupling-out efficiency in organic light-emitting<br />
devices using a thin silica aerogel layer”, Adv. Mater., Vol. 13, No.<br />
15, pp.1149-1152, 2001.<br />
[13] P.-A. Hobson, S. Wedge, J.-A.-E Wasey, I. Sage and W.-L.B Barnes,<br />
“Surface plasmon mediated emission from organic light-emitting<br />
diodes,” Adv. Mater., Vol. 14, No. 19, pp.1393-1396, 2002.<br />
[14] N.-F. Boreli, “Eficiency of microlens aray for projection LCD,”<br />
Electronic Components and Technology Conference, pp. 338-345,<br />
1994.<br />
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<br />
Integration of hybrid optical filter with buried quad<br />
pn-junction photodetector for multi-labeling<br />
fluorescence detection applications<br />
Charles Richard a , Patrick Pittet b,c , Stéphane Martel d , Luc Ouellet d , Guo-Neng Lu b,c ,<br />
Vincent Aimez a , Paul G. Charette a<br />
a Laboratoire de Biophotonique et d’Optoélectronique, Centre de recherche en nanofabrication et nanocaractérisation, Université<br />
de Sherbrooke, 2500 boul. de l’Université, Sherbrooke, Québec, J1K 2R1, Canada<br />
b Université de Lyon, F-69622, Lyon, France<br />
c CNRS, UMR5270, Institut des Nanotechnologies de Lyon, Université Lyon1, Villeurbanne<br />
d Teledyne DALSA, 18 boul. de l'Aéroport, Bromont, Québec, J2L 1S7, Canada<br />
Abstract - We present a hybrid optical filter combining<br />
interference and absorbing filtering approaches for enhancedperformance<br />
fluorescence detection systems. The filter is<br />
fabricated in a CMOS compatible process for monolithic<br />
integration with a CMOS photodetector. The rejection of<br />
fluorescence emission at 532 nm compared to 650 nm is over<br />
43 dB. The CMOS photodetector can be a BMJ (Buried<br />
Multiple-pn-Junction) structure for both photodetection and<br />
spectral discrimination. Employing a CMOS Buried Quad pn-<br />
Junction (BQJ) detector, multi-labeling spectral contributions<br />
of fluorescence emission can be determined (for cases up to 4<br />
tags).<br />
I. INTRODUCTION<br />
The emergence of technologies for biomedical<br />
applications has created new miniaturized devices that can<br />
perform the same biological analysis (for cells, bacteria,<br />
viruses, genes) as that in traditional medical laboratories [1].<br />
Also, integrated microfluidic systems with their portability,<br />
low cost and rapidity of analysis promise frontline<br />
analytical applications in food, environmental or industrial<br />
areas. Such systems can integrate detection methods for<br />
biosensors that can be mechanical, electrical, and optical.<br />
The use of fluorescent labels is an indirect optical detection<br />
method with good specificity and sensitivity [2]. The<br />
challenge with this method is to detect the relatively weak<br />
fluorescence signal in the presence of strong excitation<br />
light. The quality of this discrimination can be maximized<br />
by using high-performance optical filters and by optimizing<br />
the spectral selectivity of the photodetector.<br />
A lab-on-a-chip implementing fluorescence detection<br />
makes use of integrated optical filters and photodetector,<br />
where the optical filter is a critical component for the<br />
sensitivity of detection of the system [3]. If electronics are<br />
to be included in the chip implementation, the technological<br />
solution for filter integration must be compatible with the<br />
CMOS process. In a chip implementation, one particular<br />
challenge is the very short distance between the site of<br />
fluorescence emission and the photodetector, which requires<br />
that the intervening optical filter remain efficient over a<br />
large range of incidence angles. In addition, the<br />
photodetector may be a CMOS BMJ (Buried Multiple pn-<br />
Junction) structure, which enables both photo- and<br />
wavelength-sensitive detection [4]. The latter characteristic<br />
of the BMJ can be exploited for multiple-wavelength<br />
discrimination in multi-labelling biochemical assays.<br />
II.<br />
DEVICE AND PROCESS INTEGRATION<br />
A. Hybrid optical filter<br />
The role of the optical filter is to block the excitation<br />
signal and to allow transmission of (relatively) weak<br />
fluorescence signal. Such a filter can be either band-stop or<br />
long-pass filters – here, we consider only stop-band filters.<br />
The two main types of filter technologies used in<br />
fluorescence analysis are absorption filters (ex: colored<br />
glass or polymer) and interference filters (ex: thin-film<br />
dielectric stacks). Interference filters have the advantage of<br />
simple stop band tuning with sharp transitions [5].<br />
However, they have poor performance at off-axis<br />
illumination and high rejection in the stop-band can only be<br />
achieved at the price of many layers (up to 70). Absorbing<br />
filters are simple to integrate by spin-coating. However,<br />
autofluorescence of the filter material can be significant and<br />
can limit the sensitivity of detection. Moreover, a thick<br />
layer is required to obtain attenuations comparable to that of<br />
interference filters where the unwanted attenuation in the<br />
pass-band may become significant.<br />
To overcome the drawbacks of these conventional filters,<br />
we have developed a hybrid optical filter [6]. This filter<br />
consists of superimposed interference and absorbing filters.<br />
The interference filter is a stack of 9 alternating thin-film<br />
layers of TiO 2 /SiO 2 , deposited using E-Beam evaporation<br />
and RF sputtering. The absorbing filter is fabricated using a<br />
dilution of red dye (Orasol Red, Ciba-Geigy, USA) in<br />
KMPR negative photoresist (MicroChem Corp., USA).<br />
Orasol Red is a metal complex (Chrome complex) dye [7]<br />
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<br />
and where the Cr particles act as quenchers of fluorescence<br />
V4 V3 V2 V1<br />
similarly to TiO 2 [8].<br />
It should be mentioned that the absorbing filter must be<br />
implemented with low-autofluorescence materials (see<br />
autofluorescence measurements of different polymeric<br />
p+<br />
materials in Fig. 1). Our measurements show that the<br />
I1<br />
autofluorescence of our chosen material combination,<br />
n-base<br />
I2<br />
KMPR/Orasol, is low, comparable to that of glass substrate.<br />
P-Well<br />
I3 Deep N-Well<br />
B. CMOS BMJ (Buried Multiple pn-Junction) structure<br />
I4 P+ substrate<br />
Integrated BMJ photodetectors such as BDJ and BTJ<br />
Fig. 2 BQJ a) structure b) Chip micrograph<br />
(Buried Double/triple pn-Junction) have been reported [9-<br />
10]. These devices feature two or three photodiodes in<br />
stacked form and have an aggregate spectral response<br />
covering the visible and near-IR ranges [11]. These BMJ<br />
can thus be operated for photodetection like a photodiode,<br />
but with a more sensitive response owing to signal<br />
contributions from the multiple photodiodes. In addition,<br />
the wavelength-sensitive characteristics of BMJ detectors<br />
are particularly interesting for spectral discrimination.<br />
When measuring a ratio between two signals from the C. Process Integration<br />
stacked photodiodes, it is possible to distinguish different<br />
fluorescence spectra, and in cases of biochemical analysis,<br />
to lead to molecular identification. Such detection does not<br />
require dispersive optical devices such as gratings, and is<br />
particularly suitable for low-level fluorescence [4].<br />
With an increased number of stacked photodiodes, the<br />
spectral sensitivity as well as the capability of spectral<br />
discrimination of the BMJ detector can be improved. Thus<br />
we propose a BQJ (Buried Quad pn-Junction) detector that<br />
has been designed and fabricated using the Teledyne-<br />
DALSA (Bromont, Canada) C08G 0.8 μm multi-high<br />
voltage CMOS/DMOS process.<br />
The proposed BQJ photodetector structure (Fig. 2)<br />
consists of four stacked buried junctions consisting of a<br />
I3 + I4<br />
I2 + I3<br />
I1 + I2<br />
I1<br />
shallow p+-diffusion/n-base well junction (J1), a deeper P-<br />
Well/n-base junction (J2), a P-Well/Deep N-Well junction<br />
(J3) and P+ substrate/Deep N-Well junction (J4),<br />
respectively. It has 4 outputs allowing bias setting of the<br />
buried junctions and signal readout. Simple processing of<br />
these output signals determines the 4 photodiodes currents<br />
(I 1 , I 2 , I 3 and I 4 ) as well as their sum.<br />
For monolithic integration of a hybrid optical filter with<br />
the BQJ photodetector, we have experimented with two<br />
CMOS post-processing approaches. The first one consists in<br />
depositing the hybrid filter directly onto the packaged<br />
CMOS die (see Fig. 3). Though this technique is not<br />
compatible with industrial processes, it is nevertheless a<br />
convenient lab approach to validate the performance of our<br />
integrated hybrid filters on BMJ photodetector.<br />
The interference filter (thickness: ~ 1.2 µm) can be<br />
deposited onto a packaged CMOS BMJ photodetector as<br />
shown on Fig. 5. The interference filter was designed using<br />
the Essential Macleod software (Thin Film Center Inc.).<br />
The second step of this integration is the deposition of the<br />
absorbing component (thickness: ~ 1.6 µm) by spin-coating.<br />
An alternative process integration approach is to deposit<br />
the hybrid filter directly on a CMOS wafer, while protecting<br />
the PAD interconnections for use of the device under a<br />
probing station or for interconnects in a standard package.<br />
Fig. 1 Autofluorescence measurements of the various candidate materials for<br />
use as an absorbing filter.<br />
Fig. 3 Hybrid filter (integration of the interference filter shown into the<br />
inset) on BDJ photodetector packaged in DIP-28 package.<br />
301
Presently, the microfabrication processes we use are the liftoff<br />
process (LOR 3B + Shipley 1813) for the deposition of<br />
the interference filter and the utilization of a hard-mask for<br />
delimitation of the photodetector active region, combined<br />
with plasma O 2 for stripping the absorbance component.<br />
III.<br />
TESTS AND RESULTS<br />
A. Optical filter tests<br />
For the optical filter measurements, the stop-band<br />
attenuation is evaluated at the excitation wavelength of<br />
532 nm while the pass-band attenuation is evaluated at<br />
emission wavelengths above 650 nm.<br />
With the E-Beam evaporation manufacturing process, the<br />
interference filter has an attenuation of -12.6 dB at 532 nm<br />
and -0.76 dB at 650 nm. Absorbing filter attenuation is<br />
-32.6 dB at 532 nm and -1.28 dB at 650 nm. The<br />
combination of these two filter types gives a rejection of<br />
43 dB between the excitation and the emission bands. This<br />
spectral rejection remains practically constant over angular<br />
incidences ranging from 0 to 60 degrees (Fig. 4).<br />
To improve the fabrication reproducibility of the<br />
interference filter via a better control over film thickness,<br />
we moved from E-Beam evaporation to RF-sputtering.<br />
With this technique, the interference filters have a measured<br />
rejection of 16 dB – nearly identical to the previous<br />
fabrication process with E-Beam evaporation. Combined<br />
with absorbing component, the total rejection is 47dB.<br />
B. BQJ detector and signal processing<br />
At the BQJ detector’s outputs, reverse bias setting and<br />
photocurrent measurements were performed using a probing<br />
station. The photocurrent of each junction were calculated<br />
from I ph,i = I i – I dark,i , where I dark,i was measured in dark<br />
conditions (I dark,i = 1.04 pA, 0.86 pA, 2.22 pA, 9.29 pA, for i<br />
= 1 to 4, at room temperature).<br />
Fig. 5a shows the measured spectral responses of the BQJ<br />
detector prior to filter deposition. To eliminate fluctuations<br />
in the response due to mutual interference from the<br />
reflections at the dielectric interfaces and power variations<br />
of the optical source, the normalization ( N<br />
~ i ( ) ) of BQJ<br />
responses are used according to the following definition:<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
~ S<br />
, ( )<br />
i ( )<br />
I phi <br />
Ni<br />
( ) <br />
with i = 1, ... 4 (1)<br />
4<br />
4<br />
S j <br />
I ph,<br />
j <br />
<br />
j1<br />
j1<br />
where I ph is the photocurrent of each junction.<br />
Fig. 5b shows the normalized responses of the BQJ<br />
detector for a wavelength range between 450 nm and<br />
700 nm, showing the desired smooth curves. For each<br />
specific wavelength, the graphs on Fig. 5b show the<br />
contribution of each junction relative to the sum of the<br />
photocurrents.<br />
Using the BQJ detector for multi-tag fluorescence<br />
applications, up to four-tag contributions can be determined.<br />
To evaluate this capacity of spectral discrimination, a<br />
multiple-LED source with peak emission wavelengths<br />
respectively at λ 1 = 463nm, λ 2 = 575nm, λ 3 = 625nm,<br />
λ 4 = 655nm) was used to simulate a multi-tag fluorescencelike<br />
mixture: λ 1 , λ 1 + λ 2 , λ 1 + λ 3 , …. Using a vectorial<br />
analysis (calculation of an error vector in 4-dimensional<br />
space), multi-wavelength test optical inputs were correctly<br />
identified in twelve situations over the fifteen possibilities.<br />
Identification errors are due to the fact that two LEDs have<br />
very close emission wavelengths (λ 3 and λ 4 ). This spectral<br />
proximity caused an ill-conditioned problem in the vectorial<br />
analysis, to be avoided.<br />
(a)<br />
Fig. 4 Performance of the hybrid filter as a function of illumination angle.<br />
(b)<br />
Fig. 5 (a) Spectral response of BQJ detector (b) Normalization of BQJ<br />
outputs (Reverse bias: 1.5 V for each junction).<br />
302
The transmittance of the filter deposited onto a BDJ<br />
detector can be obtained by calculating the logarithmic ratio<br />
of the measured photocurrents before and after filter<br />
deposition (Fig. 6). We validated this measurement method<br />
by comparing results from an on-chip integrated<br />
interference filter with another filter deposited on a<br />
reference glass substrate (Eagle XG) in the same fabrication<br />
process. The measured attenuation of the interference filters<br />
was -12.0 dB and -12.6 dB, respectively. The transmittance<br />
was evaluated by monitoring the photocurrents from the two<br />
photodiodes.<br />
The transmittance of the hybrid filter compared to a<br />
reference glass substrate shows a potential of total rejection<br />
near 50 dB from 532 nm (reject band) to 650 nm (passband).<br />
However, measurements for the same filter<br />
deposited on the detector were close to the noise floor of<br />
our SMU (HP4142B, Agilent, USA). For this reason, the<br />
photocurrent at 532 nm (illumination with a frequencydoubled<br />
YAG laser) was obtained using the on-chip<br />
integrated charge amplifier coupled with the integration<br />
time method [12].<br />
The attenuation of the hybrid filter deposited on a BDJ<br />
detector was evaluated to be -60 dB at 532 nm, compared to<br />
-50 dB for the reference case with filter deposition on a<br />
glass substrate. The difference between the two cases may<br />
be due to a mismatch of deposited absorbing filter<br />
thickness. In the packaged die, the thickness is probably<br />
greater larger than that of the reference due to the non-ideal<br />
spin-coating and the presence of more pronounced edgebead.<br />
The main characteristics of the BQJ detector and the<br />
on-chip integrated filters are summarized in Table I.<br />
IV. CONCLUSION<br />
We have proposed a hybrid interference-absorbing filter<br />
with optimized performance (for normal incidence and offaxis<br />
illumination) for fluorescence detection requiring high<br />
excitation rejection and minimal autofluorescence. The<br />
rejection reaches -60 dB between 532nm to 650nm.<br />
Fig. 6 Attenuation of the interference filter over the BDJ photodetector<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Photodetector<br />
Interference<br />
filter<br />
Absorbing filter<br />
Hybrid filter<br />
TABLE I<br />
MAIN RESULTS<br />
0.8μm Multi-HighVoltage CMOS/DMOS DALSA<br />
Type : Buried Quad pn-Junction (BQJ)<br />
Active area: 200 x 200 µm² (shallow-junction)<br />
Total area: 420 x 420 µm 2<br />
Dark currents @ room temperature<br />
I 1=1.04 pA, I 2=0.86 pA, I 3=2.22 pA, I 4=9.29 pA<br />
9 thin-film layers (TiO 2/SiO 2)<br />
Thickness: ~ 1.2 µm<br />
Manufacturing process : E-Beam evaporation<br />
-12.6 dB @ 532 nm (stop-band)<br />
-0.76 dB @ 650 nm (pass-band)<br />
Manufacturing process : RF sputtering<br />
-16.6 dB @ 532 nm (stop-band)<br />
-0.5 dB @ 650 nm (pass-band)<br />
KMPR photoresist + Orasol Red<br />
Thickness: ~ 1.6 µm<br />
Manufacturing process : spin-coating<br />
-32.6 dB @ 532 nm (stop-band)<br />
-1.28 dB @ 650 nm (pass-band)<br />
Rejection between the excitation and the emission<br />
bands:<br />
43 dB for E-Beam evaporation process<br />
47 dB for RF sputtering process<br />
The developed fabrication process is compatible with<br />
CMOS integration. Its on-chip integration with a CMOS<br />
BQJ detector allows determination of multi-labeling<br />
spectral contributions (up to 4 tags).<br />
ACKNOWLEDGMENTS<br />
This work was supported in part by grants from the<br />
Natural Sciences and Engineering Research Council of<br />
Canada (NSERC), Nano-Québec, and Teledyne-DALSA.<br />
The collaborative work was supported by the Laboratoire<br />
International Associé en Nanotechnologies et Nanosystèmes<br />
(LIA-LN2).<br />
REFERENCES<br />
[1] M. Dandin, P. Abshire, and E. Smela, "Optical filtering<br />
technologies for integrated fluorescence sensors," Lab on a Chip,<br />
vol. 7, pp. 955-77, 2007/08/ 2007.<br />
[2] G. T. Roman and R. T. Kennedy, "Fully integrated microfluidic<br />
separations systems for biochemical analysis," Journal of<br />
Chromatography A, vol. 1168, pp. 170-188, Oct 19 2007.<br />
[3] R. Bashir, "BioMEMS: State-of-the-art in detection, opportunities<br />
and prospects," Advanced Drug Delivery Reviews, vol. 56, pp.<br />
1565-1586, 2004.<br />
[4] P. Pittet, J. M. Galvan, G. N. Lu, L. J. Blum, and B. D. Leca-<br />
Bouvier, "CMOS LIF detection system for capillary analysis,"<br />
Sensors and Actuators B (Chemical), vol. B97, pp. 355-61, 2004.<br />
[5] K.-S. Shin, Y.-H. Kim, J.-A. Min, S.-M. Kwak, S. K. Kim, E. G.<br />
Yang, J.-H. Park, B.-K. Ju, T.-S. Kim, and J. Y. Kang,<br />
"Miniaturized fluorescence detection chip for capillary<br />
electrophoresis immunoassay of agricultural herbicide atrazine,"<br />
Instrumental Methods of Analysis - IMA 2005, Analytica Chimica<br />
Acta, vol. 573-574, pp. 164-171, 2006.<br />
303
[6] C. Richard, A. Renaudin, V. Aimez, and P. G. Charette, "An<br />
integrated hybrid interference and absorption filter for fluorescence<br />
detection in lab-on-a-chip devices," Lab on a Chip, vol. 9, pp.<br />
1371-1376, 2009.<br />
[7] K. Wojtach, et al., "Characteristics of colored inorganic-organic<br />
hybrid materials," Journal of Non-Crystalline Solids, vol. 353, pp.<br />
2099-2103, 2007.<br />
[8] M. Yamazaki, et al., "Non-emissive colour filters for fluorescence<br />
detection," Lab on a Chip, 2011.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
[9] K. Liang, W. Li, H.R. Ren, X.L. Liu, W.J. Wang, R. Yang, D.J.<br />
Han“Color measurement for RGB white LEDs in solid-state<br />
lighting using a BDJ photodetector”, Displays, vol. 30 pp. 107–<br />
113, 2009.<br />
[10] M. B. Chouikha, G. N. Lu, M. Sedjil, and G. Sou, "Colour<br />
detection using buried triple pn junction structure implemented in<br />
BiCMOS process," Electronics Letters, vol. 34, pp. 120-122, 1998.<br />
[11] G. N. Lu, "A dual-wavelength method using the BDJ detector and<br />
its application to iron concentration measurement," Measurement<br />
Science & Technology, vol. 10, pp. 312-315, 1999.<br />
[12] P. Pittet, et al., "Variable time synchronous detection method for<br />
sensitive optical detection," Electronics Letters, vol. 39, pp. 860-<br />
862, 2003.<br />
304
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May 2011, Aix-en-Provence, France<br />
<br />
Fabrication and Characteristics of a Fused<br />
Silica-Based Optical Waveguide with Femtosecond<br />
Fiber Laser Pulses<br />
Ting-Chou Chang 1 , Chien-Hsing Chen 2 , Wei-Hung Shih 3 , Jian-Neng Wang 4 , Chai-Yu Lee 1 , Jaw-Luen Tang 2 , Shau-Chun<br />
Wang 1 , Lai-Kwan Chau 1 , Wei-Te Wu 5*<br />
1 Department of Chemistry and Biochemistry, National Chung Cheng University<br />
168 University Road, Minhsiung, Chiayi 621, Taiwan<br />
2 Department of Physics, National Chung Cheng University<br />
168 University Road, Minhsiung, Chiayi 621, Taiwan<br />
3 Department of Mechanical Engineering, National Chung Cheng University<br />
168 University Road, Minhsiung, Chiayi 621, Taiwan<br />
4 Department of Construction Engineering, National Yunlin University of Science and Technology,<br />
123 University Road, Section 3, Douliou, Yunlin 640, Taiwan<br />
5* Department of Biomechatronics Engineering, National Pingtung University of Science and Technology<br />
1, Shuefu Road, Neipu, Pingtung 912, Taiwan<br />
Tel: +886-8-770-3202 Ext. 7599; Fax: + 886-8-774-0420; weite@mail.npust.edu.tw<br />
Abstract<br />
This study investigates the fabrication characteristics<br />
of a femtosecond fiber laser on a fused-silica-based optical<br />
waveguide. The wavelength and repetition rate of the<br />
femtosecond fiber laser are 532 nm and 1 MHz,<br />
respectively. We selected three main fabrication<br />
parameters for systematic adjustment: laser power (E),<br />
scanning speed ( v s<br />
) and focus depth (d = 0 at the surface<br />
of substrate). We succeeded in fabricating a waveguide<br />
layer inside the silica subtracts. By analyzing the light<br />
translation path and the net fluence in the waveguide, the<br />
range of fabrication energy of the waveguide on the fused<br />
silica was kept within 0.973 - 1.438 KJ/cm 2 .<br />
I. Introduction<br />
Recently, developments in nanotechnology have led to<br />
a proliferation of electro-optic system applications. To<br />
minimize system size, industries including communications,<br />
construction and biomedical detection have widely applied<br />
optical waveguides such as photonic crystal fibers [1], fiber<br />
interferometers [2], surface plasma resonance (SPR) sensors<br />
[3], localized plasma resonance (LPR) sensors [4] and<br />
guided-mode resonance (GMR) sensors [5].<br />
Waveguide device are fabricated through techniques<br />
including ion bombardment, laser machining,<br />
photolithography, and mechanical stamping [6], commonly<br />
using fused silica as a substrate. Laser machining is a low cost,<br />
high speed and high yield method for the localized heat<br />
treatment of fused silica. However the linear absorption of<br />
fused silica depends on the laser source. Using an ultraviolet<br />
laser requires a process to bind oxygen to the fused silica to<br />
increase light sensitivity [7]. Using CO 2 laser [8] results in a<br />
greatly increased linear absorption of the fused silica which<br />
makes precise machining more difficult and can cause<br />
damage around the machining area. High-power density<br />
femtosecond fiber lasers with a pulse of 10 -15 seconds are an<br />
appropriate tool for the fabrication of optical waveguides due<br />
to their independence in the linear absorbing effect of fused<br />
silica.<br />
This study investigates the fabrication characteristics of<br />
femtosecond fiber lasers on fused-silica-based optical<br />
waveguides. We selected three main fabrication parameters,<br />
laser power (E), scanning speed ( v s<br />
) and focus depth (d = 0 at<br />
the surface of substrate) which are systematically adjusted to<br />
investigate the differences of post-machining light waveguide<br />
characteristics, transmission loss rate and the relation<br />
between the net influence and light waveguide.<br />
II. Experimental section<br />
1. Waveguide principles<br />
As shown in Fig. 1, the light waveguide is composed of<br />
a layer of Media 1 (i.e. a media different from the substrate)<br />
sandwiched between two layers of Media 2 (i.e. the<br />
substrate).<br />
One of two application phenomena of light waveguides<br />
is the refraction within these media with different refraction<br />
indices. Based on the Snell’s law, the refraction angle, φ , is<br />
smaller than the incident angle, θ , as light is incident into<br />
Media 1. The other application phenomenon is total reflection<br />
for keeping and transmitting all laser energy within the Media<br />
1 layer. This means that Snell’s law requires the refraction<br />
index of Media 1, n 1 , to be larger than that of Media 2.<br />
The numerical aperture (NA), (i.e., the maximum<br />
acceptable energy of light wave guide), is defined as.<br />
305
NA sinθ c<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
2 2<br />
<br />
1<br />
−=≡<br />
nn (1) rotating the half-wavelength and polarization slides. The laser<br />
2<br />
is focused by an objective lens (Mitutoyo-10X,NA=0.28).<br />
The charge-coupled device (CCD) is used to help aim the laser<br />
on the machining area to ensure beam quality.<br />
The machining path and machining rate are controlled by<br />
programming the X-Y micro-positioning platform. An optical<br />
microscope is used to inspect the machined products.<br />
where θ c is the maximum acceptance angle.<br />
The light among the incident light increases with NA. If<br />
the incident angle is larger than θ c , some light is refracted<br />
into Media 2. Therefore, the incident angle must be smaller<br />
thanθ c to satisfy the total reflection and forming guide mode.<br />
Air, n air<br />
n 2<br />
media 2<br />
n 1<br />
media 1<br />
n 2<br />
III. Results and Discussion<br />
1. Waveguide fabrication<br />
In this study we selected three main fabrication<br />
parameters, laser power (E), scanning speed ( v s<br />
) and focus<br />
depth (d = 0 at the surface of substrate). By fixing the laser<br />
power at 170 mW and the focus depth at 0 μm, the fused silica<br />
was modified at scanning speeds.<br />
media 2<br />
Fig. 1 Waveguide translation principle<br />
In general, the fused silica is homogeneous with the<br />
constant refraction index. However, the refraction index of<br />
fused silica increases with the annealing rate [9]. The<br />
femtosecond laser’s pulse characteristic makes it appropriate<br />
for decreasing the annealing rate. The pulse energy does not<br />
integrated easily in the working area and results in a lower<br />
annealing rate.<br />
5.1μm<br />
(a)<br />
1mm/s<br />
4.1μm<br />
3.0μm<br />
2.7μm<br />
(b) (c) (d)<br />
2mm/s 3mm/s 4mm/s<br />
2. Experimental Setup<br />
The specifics of the apparatus used in this study are<br />
shown in Table 1 and Fig. 2. The central wavelengths of the<br />
laser are 532 and 1064 nm, the pulse duration is less than 400<br />
fs and the repetition rate is 1 Hz – 1 MHz. The laser beam is a<br />
Gaussian beam.<br />
Fig. 2 Femtosecond fiber laser machining system schematic<br />
Table 1 Femtosecond fiber laser machining system<br />
specification<br />
Wavelength 1064 nm & 532 nm<br />
Repetition rate<br />
1 Hz~1 MHz<br />
Pulse duration<br />
increased.<br />
11-13 May 2011, Aix-en-Provence, France<br />
Power meter<br />
LD@1553nm<br />
collimator<br />
3.6μm<br />
2.5μm<br />
MMF-fiber<br />
(a)<br />
1mm/s<br />
(b)<br />
2mm/s<br />
(c)<br />
3mm/s<br />
Fig. 4 Fabrication with various scanning speeds at E = 170 mW and<br />
d = 10 μm<br />
3.6μm<br />
(a)<br />
5mm/s<br />
3.7μm<br />
(d)<br />
8mm/s<br />
3.3μm<br />
(b)<br />
6mm/s<br />
3.3μm<br />
(e)<br />
9mm/s<br />
2.3μm<br />
(c)<br />
7mm/s<br />
2.5μm<br />
(f)<br />
10mm/s<br />
Fig. 5 Fabrication with various scanning speeds at E = 170<br />
mW and d = 0 μm<br />
The laser power was increased to 230 mW to modify fused<br />
silica 10 μm in depth. Scanning speed should be increased to<br />
avoid surface ablation. The results in Fig. 5 show that the<br />
fused silica is ablated given scanning speeds between 5 mm/s<br />
and 7 mm/s; and modified widths of3.7μm, 3.3μm and 2.5μm<br />
correspond to scanning speeds of 8 mm/s, 9 mm/s and 10<br />
mm/s, proving that the fused silica can be modified at different<br />
depths through focusing and tuning the laser power and<br />
scanning speed.<br />
2. Waveguide propagating loss measurement<br />
Fig. 6 shows the system for measuring waveguide<br />
propagating loss. The system conducts the laser diode (LD,<br />
center wavelength = 1553 nm) to the waveguide layer on the<br />
XYZ-rotation stage. In the end of waveguide layer, the<br />
collimator couples the multi-mode optic fiber to the power<br />
meter for acquiring and analyzing signal. The results show<br />
that the propagating loss are 4.6 dB/cm、4.8 dB/cm、6.2<br />
dB/cm as the scanning velocities are 8 mm/s, 9 mm/s and 10<br />
mm/s, respectively, with 230 mW and 10 μm of depth. It<br />
indicates that the increased scanning velocity causes the<br />
larger energy loss due to the low absorbing energy of fused<br />
silica.<br />
XYZ-rotation<br />
stage<br />
Transmission (dBm)<br />
0<br />
-5<br />
-10<br />
-15<br />
-20<br />
-25<br />
-30<br />
-35<br />
-40<br />
LD<br />
-45<br />
1553.2 1553.4 1553.6 1553.8 1554.0 1554.2<br />
Wavelength (nm)<br />
Fig. 6 The system for measuring waveguide propagating loss<br />
Table 2 Fabrication parameters of waveguide using<br />
femtosecond laser<br />
Laser<br />
power<br />
E<br />
(mW)<br />
170<br />
230<br />
Scanning focusing<br />
velocity depth<br />
NF<br />
results<br />
v<br />
s d<br />
(KJ/cm 2 )<br />
(mm/s) (μm)<br />
1<br />
ablation 7.191<br />
2 ablation 3.596<br />
3 ablation 2.397<br />
4 0 ablation 1.798<br />
5 waveguide 1.438<br />
6 waveguide 1.198<br />
7 waveguide 1.027<br />
5<br />
ablation 1.946<br />
6 ablation 1.621<br />
7 ablation 1.390<br />
10<br />
8 waveguide 1.216<br />
9 waveguide 1.081<br />
10 waveguide 0.973<br />
3. Waveguide discussion<br />
The laser power, the diameter of the laser beam, the<br />
scanning speed and the rate of repetition are all very<br />
influential factors in laser machining. This study analyzes the<br />
machining performance with NF factor [10], as shown below.<br />
2ω0<br />
PRF<br />
NF =<br />
(5)<br />
vs<br />
where ω is the minimal radius of the laser beam, R = 1<br />
0<br />
MHz is the repetition rate, and E<br />
F p<br />
= is the average<br />
2<br />
Rπω 0<br />
fluence per lasing. In this study, the laser wavelength, λ , is<br />
532 nm. The focus distance of the lens, f , is 20 mm. The<br />
diameter of the incident laser, D, is 5 mm. The ω is 1.5 μm,<br />
0<br />
estimated by the 1.05 of the measured laser beam quality<br />
factor ( D<br />
M<br />
2 πω0<br />
= ). E is the laser power. Substituting these<br />
2λf<br />
parameters into Eq.(5), the range of the NF value is 1.438 –<br />
0.973 KJ/cm 2 , indicating the range of fabrication energy of<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
307
the waveguide on the fused silica, as shown in Table 2 and Fig.<br />
7. Furthermore, based on Table 2, the laser power should be<br />
increased with the scanning speed, thus increasing machining<br />
speed.<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
silica is kept within 1.438 – 0.973 KJ/cm 2 . Future research<br />
could study parameters such as NA, NF and transmission loss<br />
in optimal methods to develop new design applications for<br />
biochemistry sensors and micro-optic systems.<br />
Modification type<br />
Waveguide<br />
No change<br />
Damage<br />
d=0μm<br />
d=10μm<br />
1 2 3 4 5 6 7<br />
NF (kJ/cm 2 )<br />
Fig. 7 The modification type with NF variation<br />
In this study, a waveguide fabricated with 170 mW laser<br />
power, 5mm/s scanning speed and 0 μm focus depth<br />
successfully conducted light as shown in Fig. 8, proving that<br />
femtosecond lasers can be used to fabricate waveguides on<br />
fused silica.<br />
Fig. 8 Waveguide fabricated with 170 mW laser power,<br />
5mm/s scanning speed and 0 μm focus depth<br />
IV. Conclusions<br />
This study describes the successful fabrication of a<br />
fused-silica-based optical waveguide using a femtosecond<br />
fiber laser, and an investigation of the fabrication<br />
characteristics of femtosecond fiber lasers on<br />
fused-silica-based optical waveguides. The results show that<br />
the modified width decreases with increasing scanning speed,<br />
regardless of machining depth. By analyzing the light<br />
translation path and the net fluence in the waveguide, the<br />
range of fabrication energy of the waveguide on the fused<br />
References<br />
1. C. H. Chen, S. C. Chen, Y. C. Chen, H. T. Hu, T. H.<br />
Wei, W. T. Wu, J. N. Wang and J. L. Tang, Research<br />
on laser-induced long-period fiber grating sensor<br />
modified with gold nano-rods, The 8th Pacific Rim<br />
Conference on Lasers and Electro-Optics, Shanghai,<br />
2009<br />
2. Chien-Hsing Chen, Yi-Chun Chen, Jian-Neng Wang,<br />
Lai-Kwan Chau, Jaw-Luen Tang and Wei-Te Wu,<br />
“Multimode fiber Mach–Zehnder interferometer for<br />
measurement of refraction index”, IEEE Sensors<br />
2010 Conference - the 9th Annual IEEE Conference<br />
on Sensors, 2010/11/1-2010/11/4, USA.<br />
3. Y. Liu, J. Kim, Numerical investigation of finite<br />
thickness metal-insulator-metal structure for<br />
waveguide-based surface plasmon resonance<br />
biosensing, Sens. and Actu. B, Vol. 148, pp. 23-28,<br />
2010.<br />
4. L. K. Chau, Y. F. Lin, S. F. Cheng, and T. J. Lin,<br />
Fiber-optic chemical and biochemical probes based<br />
on localized surface plasmon resonance, Sens. and<br />
Actu. B, Vol. 113, pp. 100–105, 2006.<br />
5. Ian D. Block, Nikhil Ganesh, Meng Lu, and Brian T.<br />
Cunningham, ”Bulk-Micromachined Optical Filer<br />
Based on Guided-Mode Resonance in Silicon-Nitride<br />
Membrane,” IEEE Sens. J., Vol. 8, pp.274-280, 2008.<br />
6. C. S. Ma, W. B. Guo, D. M. Zhang, K. X. Chen, Y.<br />
Zhao, F. Wang, Z. C. Cui, S. Y. Liu, Analytical<br />
modeling of loss characteristics of a polymer arrayed<br />
waveguide grating multiplexer, Vol. 34 PP. 621-630,<br />
2002.<br />
7. C. Chen, X. Sun, D. Zhang, Z. Shan, S. Y. Shin, D.<br />
Zhang, Dye-doped polymeric planar waveguide<br />
devices based on a thermal UV-bleaching technique,<br />
Optics & Laser Technology, Vol. 41 , pp. 495–498,<br />
2009.<br />
8. A. M. Vengsrlar, P. J. Lemaire, et al. “Long-Period<br />
Fiber Gratings as Band-Rejection Filters,” Journal of<br />
Lightwave Technology, vol. 4, pp. 58-65, 1996.<br />
9. J. W. Chan, T. R. Huser, S. H. Risbud, J. S. Hayden, D.<br />
M. Krol, Waveguide fabrication in phosphate glasses<br />
using femtosecond laser pulses, APPLIED PHYSICS<br />
LETTERS, Vol. 82, pp. 2371-2373, 2003.<br />
10. L. Shah, Y. A. Arai, S. M. Eaton, P. R. Herman,<br />
Waveguide writing in fused silica with a femtosecond<br />
fiber laser at 522 nm and 1 MHz repetition rate,<br />
Optics Express, Vol. 13, pp. 1999-2006, 2005.<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
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11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Capacitive Microphone fabricated with<br />
CMOS-MEMS Surface-Micromachining Technology<br />
Josué Esteves, Libor Rufer, Gustavo Rehder<br />
TIMA Laboratory (CNRS, G-INP, UJF)<br />
46 Avenue Félix Viallet, Grenoble, France<br />
Abstract- This paper presents a standard complementary metal–<br />
oxide–semiconductor (CMOS) technology combined with<br />
sacrificial layers etching used for micro-electro-mechanical<br />
systems (MEMS) fabrication. We will describe the modeling and<br />
the design of an acoustic device represented here by a condenser<br />
microphone with a perforated diaphragm. Models based on the<br />
electromechanical analogy and on the finite element analysis<br />
(FEA) have been used to predict the behavior of the microphone.<br />
These models have taken into account material constants and<br />
dimensions of the AMS 0.35 μm CMOS technology. An effect of<br />
etch holes in the microphone diaphragm on the dynamic response<br />
of the structure was studied and an optimization study has been<br />
done to determine the sensor lateral dimensions and the position<br />
of these holes. We will show simulation results, the microphone<br />
design and the final layout of the structure.<br />
I. INTRODUCTION<br />
Several studies have been carried out since last two decades<br />
with the aim to use a standard CMOS process to fabricate<br />
micro-electro-mechanical systems [1], [2]. This, so called,<br />
CMOS-MEMS process was originally introduced as a<br />
technique using back-side bulk micro-machining (BSBM), and<br />
later front-side bulk micro-machining (FSBM) of a silicon<br />
wafer, thus allowing to free-up layers deposited on the silicon<br />
wafer during the CMOS process. The technologies based on<br />
the bulk etch of silicon made possible to design different<br />
devices with movable elements made of a stack of several<br />
layers deposited on top of silicon wafer during a CMOS<br />
process. Different realizations using this approach can be<br />
found in [3], [4].<br />
More recently, another approach to a CMOS-based MEMS<br />
fabrication was proposed [5]. This technique is based on<br />
surface micromachining applied on specific layers issued from<br />
CMOS process. Thus metal layers can be considered as<br />
sacrificial and can be removed for instance with PAN etch<br />
(Phosphoric, acetic and nitric acids). Remaining structure can<br />
be then composed of silicon nitride, silicon oxide and<br />
polysilicon layers [5]. In some cases, it is more convenient to<br />
keep metal layers for device design. In this case, silicon oxide<br />
is chosen as a sacrificial layer and is removed for example<br />
with buffered oxide etch (BOE) saturated with aluminum<br />
(silox vapox III – transene) [6]. Further, the silicon oxide can<br />
also be etched by vapor HF, which does not attack the<br />
aluminum layers and prevents inter-layers stiction.<br />
Advantages of this new CMOS-MEMS fabrication<br />
technique are easy execution, low-cost maskless etching and<br />
the possibility to integrate the electronic circuit and MEMS<br />
device together on the same chip.<br />
We have considered to use this CMOS-MEMS technology<br />
to fabricate a condenser microphone for the audible frequency<br />
range. Different MEMS microphones have been developed by<br />
different groups and some designs, using dedicated technology<br />
process, have been commercialized until now. Most of the<br />
designs consist of a movable diaphragm and a perforated fixed<br />
electrode, called the backplate, separated by an air gap (see<br />
Fig. 1). Below the backplate, there is a back-chamber that<br />
makes the evacuation of air from the air gap easier.<br />
Fig. 1. Schematic structure of a conventional MEMS condenser microphone.<br />
In this work, we have studied and developed a model of a<br />
condenser microphone with a different structure. Our<br />
microphone structure does not contain a back-chamber, and is<br />
composed of a movable perforated electrode and a fixed<br />
electrode (without holes), separated by an air gap. This kind of<br />
structure allows using of a standard CMOS process with only<br />
one additional post-process step to etch sacrificial layers to<br />
realize the MEMS device. Similar microphone structure with<br />
perforated movable electrode was recently described in [7]. In<br />
this article, a specific dedicated technology was used to create<br />
an aluminum diaphragm, and the modeling does not take into<br />
account the air gap effect that is very important for the<br />
frequency response of the microphone.<br />
The paper is structured as follows. In Section 2, we will<br />
describe the microphone structure and propose an equivalent<br />
circuit model with lumped-parameters for the microphone<br />
modeling taking into account the effect of etch holes on the<br />
microphone behavior. In Section 3, simulations with<br />
CoventorWare, commercial FEA simulation software for<br />
MEMS, are performed in order to determine different<br />
parameters of the equivalent circuit and to estimate the<br />
microphone performance. Next, Section 4 will present the<br />
fabrication using AMS 0.35 µm CMOS standard process.<br />
Finally, Section 5 will provide some conclusions and<br />
directions on our future research.<br />
II. MICROPHONE STRUCTURE<br />
The microphone is fabricated with the AMS 0.35 µm<br />
CMOS back-end process resulting in a passivation layer, four<br />
metal layers, three via layers and several silicon dioxide<br />
layers. In our design, we can create the diaphragm and the<br />
backplate of the microphone with metal layers, the air gap<br />
between the electrodes can be realized by etching the<br />
sacrificial silicon dioxide layer through small holes in the<br />
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microphone diaphragm. The chosen CMOS technology<br />
imposes the vertical structure dimensions i.e. the air gap and<br />
the diaphragm thicknesses. Fig. 2 shows different views of the<br />
microphone structure and the corresponding dimensions are<br />
listed in Table I.<br />
Fig. 2. Microphone cross-sectional view (a), top view of the diaphragm (b).<br />
TABLE I<br />
DIMENSIONS OF THE MICROPHONE ELEMENTS<br />
Elements<br />
Dimension (µm)<br />
Diaphragm side length (L mem) 500<br />
Diaphragm thickness (t mem) 1<br />
Arm length (L arm) 71<br />
Arm width (W arm) 141<br />
Big hole side length (L hole1) 5<br />
Big hole pitch (Pitch1) 10<br />
Small hole side length (L hole2) 1<br />
Small hole pitch (Pitch2) 5<br />
Air gap thickness (h a) 2.64<br />
We have considered a diaphragm supported by four beams<br />
anchored by the oxide layer in order to obtain an optimum<br />
stiffness and thus acceptable sensitivity. The air gap thickness<br />
that was taken into account in this design corresponds to the<br />
distance between the metal layers M4 and M2 of the CMOS<br />
process. Two sets of holes are designed in the microphone<br />
diaphragm. Small holes with the sides of 1 µm are densely<br />
distributed on the diaphragm surface with the aim to allow fast<br />
sacrificial layer etching. Larger holes with the sides of 5 µm<br />
disposed on the diaphragm have the important role of<br />
controlling the diaphragm damping and their dimensions were<br />
optimized in order to obtain flat frequency response close the<br />
resonance.<br />
The microphone must work in the audible frequency range<br />
from 20 Hz to 20 kHz. The required capacitance variation of<br />
the microphone was obtained through the analysis of the<br />
expected noise of the electronic circuit. For a minimum signalto-noise<br />
ratio of 40 dB, the microphone capacitance variation<br />
must be at least 60 fF.<br />
In this section, we will describe the microphone equivalent<br />
circuit with lumped parameters. We present the effect of holes<br />
on the mechanical parameters and on the electrostatic field<br />
distribution, as well as their influence on the mechanical<br />
damping.<br />
A. Equivalent circuit<br />
Lumped parameters equivalent circuit (Fig. 3) is used to<br />
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May 2011, Aix-en-Provence, France<br />
<br />
study the frequency response of the microphone. An analogy<br />
between the acoustic, mechanical, fluidic and electrical<br />
domains is used to build the equivalent circuit.<br />
To characterize the mechanical and fluidic part behavior of<br />
the MEMS microphone, an equivalent spring-mass-damper<br />
system under harmonic excitation is considered. The total<br />
stiffness of the system is given by the rigidity of the<br />
diaphragm and the spring effect of the air gap, k mem and k airgap<br />
respectively. The resistance R airgap represents the damping<br />
caused by the viscous losses in the air gap and M mem is the<br />
effective diaphragm mass.<br />
In the acoustic domain, a sound pressure (P g ) is applied to<br />
the diaphragm through the radiation impedance composed of<br />
the radiation resistance R rad , representing the frictional force,<br />
and by the radiation mass M rad , representing the mass of the air<br />
close to the diaphragm that is vibrating in phase with the plate.<br />
In the electrical domain, the capacitance of the microphone<br />
is represented by C 0 and the parasitic capacitance, due to the<br />
electric field in the oxide (anchor of the microphone) is C p .<br />
The link between the acoustic and mechanical domain is<br />
modeled by the mechano-acoustic transformer with the ratio<br />
A mem , representing the diaphragm area. The second transformer<br />
represents the coupling between electrical and mechanical<br />
domain with the ratio Γ.<br />
Fig. 3. Microphone equivalent circuit.<br />
The following paragraphs discuss the different parameters<br />
of the equivalent circuit taking into account the effect of etch<br />
holes.<br />
B. Mechanical behavior with holes<br />
The analytic description of the microphone mechanical<br />
behavior is quite complex and difficult to solve because of the<br />
diaphragm geometry (arms, holes). For this reason, an<br />
approach using finite element simulations and reduced<br />
elements approximation is used. Indeed, we can determine the<br />
mechanical parameters of the equivalent circuit that represents<br />
the diaphragm, namely its spring coefficient k mem and effective<br />
mass M mem , with FEA performed in CoventorWare and<br />
consider the diaphragm as a spring-mass system governed by<br />
the well-known relations:<br />
mem==<br />
maxwkP (1) AF<br />
kmem<br />
=f<br />
(2)<br />
0<br />
M<br />
mem<br />
Simulations can calculate the resonant frequency (f 0 ) and the<br />
maximum displacement (w max ) of the structure when we<br />
applied a uniform force F corresponding to a pressure P on the<br />
diaphragm. From w max and knowing the diaphragm area (A),<br />
we can calculate the spring coefficient of the diaphragm with<br />
(1). From the resonant frequency and the calculated spring<br />
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coefficient, we can calculate the effective mass of the<br />
diaphragm with (2).<br />
However, the mechanical properties, namely the effective<br />
Young’s modulus and the internal stress of a MEMS structure,<br />
are influenced by its perforation [8]. We have performed<br />
estimations of this effect and confirmed with CoventorWare<br />
the results of [9] showing the stress concentration in the<br />
proximity of holes (see Fig. 4). Simulations for several<br />
diaphragms with different holes configurations have shown<br />
that the resonant frequency for a perforated diaphragm<br />
(f 0WithHole ) can be approximately estimated with the following<br />
relation:<br />
f 0WithoutHole<br />
A<br />
A<br />
WithHole<br />
WithoutHole<br />
≤ f 0WithHole ≤ f 0WithoutHole (3)<br />
where A WithHole is the perforated diaphragm area, A WithoutHole is<br />
the entire diaphragm area and f 0WithoutHole is the resonant<br />
frequency of the diaphragm without holes. According to these<br />
different observations, we have to consider the etch holes in<br />
the calculation of k mem and M mem .<br />
Fig. 4. Simulation showing the stress on the diaphragm with holes.<br />
The simulations of the microphone were performed on the<br />
quarter of the structure using symmetry plane conditions<br />
(Fig. 5) because of the high number of etch holes (several<br />
thousands) that demands intensive computational resources .<br />
CoventorWare calculates the resonant frequency and the<br />
maximum displacement of the structure, and with (1) and (2),<br />
we can calculate the spring coefficient k mem and the effective<br />
mass M mem of the perforated diaphragm (Table II).<br />
Fig. 5. Simulation structure: quarter model of the microphone using symmetry<br />
planes.<br />
TABLE II<br />
MICROPHONE MECHANICAL PARAMETERS<br />
Simulation results Calculated parameters<br />
f 0 = 13491 Hz k mem = 3.8 N/m<br />
w max = 88 nm at 1 Pa M mem = 5.3x10 -10 kg<br />
It can be noticed that the resonant frequency of the<br />
perforated structure obtained from the FEA (13491 Hz)<br />
corresponds to the relation (3).<br />
C. Air gap modeling<br />
When the diaphragm oscillates normally to the backplate,<br />
the air gap between the diaphragm and the backplate is<br />
squeezed causing a lateral fluid motion in the gap. Due to the<br />
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<br />
viscous flow of air, pressure in the air gap changes and creates<br />
forces against the diaphragm movement. This phenomenon<br />
called squeeze film damping is accompanied with two kinds of<br />
forces. One is the damping force, caused by the viscous flow<br />
of air, and the other is the elastic force due to the compression<br />
of the air gap. The squeeze film can be described by the<br />
Navier-Stokes and Reynolds equations taking into account<br />
some effects that are specific to the small film dimensions,<br />
like air rarefaction, compressibility and inertia effects, and air<br />
flow through the holes of a perforated plate. The squeeze film<br />
damping is very important for the microphone operation, in<br />
particular in the high frequencies, in the vicinity of its<br />
resonance [8].<br />
The air gap stiffness k airgap and the damping R airgap are<br />
proportional to the elastic and damping forces respectively.<br />
Several analytic models have been proposed to calculate the<br />
elastic and damping forces and so the corresponding elements<br />
of the model ([10]-[14]). These models take into account<br />
rarefaction, compressibility and inertia effects as well as the<br />
effect of etch holes and are a helpful alternative to the<br />
microphone FEM simulations that can have constraints in<br />
computational requirements.<br />
Although it exists several similar models, we have decided<br />
to choose the model described in [13]. This model provides<br />
relations for damping force (F damp ) and stiffness force for<br />
perforated plate taking into account rarefaction,<br />
compressibility effects and also inertia effects (F stiff+iner ).<br />
2<br />
Fdamp { I=<br />
net<br />
}<br />
a 0)(rPF (4)<br />
{ }<br />
2<br />
Fstiff + iner<br />
R=<br />
net a 0)(rPF (5)<br />
where P a is the atmospheric pressure, r 0 is the outer radius of a<br />
pressure cell (proportional to the pitch), F net is the real<br />
complex force acting on the diaphragm due to the squeeze film<br />
given by the following expression:<br />
sq<br />
sq2<br />
F net = F sq1 + F sq2 + F h (6)<br />
⎡ 2Ri<br />
[<br />
i<br />
i<br />
)j() ] ⎤<br />
1 1<br />
1<br />
1<br />
ΓΓ−<br />
2 jτ KRj(I)<br />
π=F ξ<br />
01 ⎢<br />
1−−<br />
i ⎥e)R(<br />
(7)<br />
⎢⎣<br />
jΓ<br />
0 i 1<br />
1<br />
0<br />
ΓΓΓΓ<br />
i<br />
)Rj() ⎥⎦<br />
Kj(I+<br />
⎡ 2R [<br />
i<br />
i<br />
)j() ] ⎤<br />
i 1<br />
1<br />
1<br />
Γ−ΓΓ<br />
1<br />
jτ (8) KRj<br />
π=F Φ<br />
b ⎢<br />
⎥e<br />
⎢⎣<br />
jΓ<br />
0<br />
Γ<br />
i 1<br />
Γ<br />
1<br />
0<br />
ΓΓ<br />
i<br />
)Rj() ⎥⎦<br />
Kj<br />
h<br />
jτ<br />
[ π ]<br />
2<br />
)(<br />
Φ=<br />
eRF<br />
(9)<br />
bi<br />
where F sq = F sq1 + F sq2 is the complex force due to the<br />
squeeze-film, F h is the force acting on the hole surface, ξ o is<br />
the non-dimensional amplitude, Ф b is the non-dimensional<br />
pressure at the hole/air gap interface, Г is a non-dimensional<br />
complex number that includes the compressibility, inertia and<br />
gas rarefaction effects, R i and R 0 are respectively the nondimensional<br />
inner and outer radii of a pressure cell, I n is the<br />
Modified Bessel function of n th order, K n is the Macdonald’s<br />
function of n th order and τ is the non-dimensional time.<br />
Knowing that F damp and F stiff.+iner. are respectively the<br />
imaginary and real parts of F net , we can calculate the air gap<br />
stiffness k airgap and the damping R airgap .<br />
In order to respect the requirements on the microphone<br />
performance, we must achieve a negligible spring effect and<br />
low damping due to the air gap. The size of the holes and the<br />
pitch have been chosen taking into account the etch time, low<br />
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damping and negligible spring effect in the frequency range of<br />
<br />
Air gap spring coefficient<br />
interest. So, a non-staggered configuration with 5 x 5 µm²<br />
The air gap spring coefficient, k airgap , can be calculated<br />
square holes and 10 µm pitch is used.<br />
from (5):<br />
Air gap damping coefficient<br />
Fstiff<br />
+ iner<br />
k<br />
airgap<br />
= (11)<br />
The damping coefficient, R airgap , can be determined from<br />
ha<br />
(4):<br />
In a similar way, we compare FEM and lumped-model<br />
Fdamp<br />
results for various configurations, but there were important<br />
R<br />
airgap<br />
= (10)<br />
h<br />
errors (more than 80%). Even if this model is not accurate<br />
aω<br />
enough to calculate the air gap spring coefficient, according to<br />
First, we compare FEM and lumped-model results for<br />
the simulation result for the microphone case, in the audio<br />
various configurations. For simulations, we use CoventorWare<br />
frequency range, the air gap spring coefficient,<br />
which provides the damping force and coefficient for squeezefilm.<br />
Next, we apply the lumped-model on the chosen<br />
k airgap = 0.006 N/m (at 20 kHz), is very low comparing to the<br />
diaphragm spring coefficient k mem (Table II). Therefore, the air<br />
configuration. Considering possible device applications,<br />
gap stiffness can be neglected. Indeed, according to the<br />
simulations were performed up to 100 kHz.<br />
obtained air gap spring value, we can suppose a condition of<br />
Table III shows the air gap damping coefficient error<br />
incompressible fluid. This can be also confirmed by the<br />
between the FEM and the lumped-model results for different<br />
squeeze number σ estimation, which characterizes the<br />
diaphragm sizes and thus for different number of holes. We<br />
compressibility effect for a perforated diaphragm:<br />
have compared several diaphragm configurations. In this table,<br />
2<br />
12 μω r0<br />
the air gap value similar to that of the designed microphone<br />
σ = (12)<br />
2<br />
was fixed and we have used the non-staggered (matrix) hole<br />
hP<br />
aa<br />
configuration with 10 µm pitch for each diaphragm.<br />
TABLE III<br />
MODEL/SIMULATION ERRORS<br />
Diaphragm size Damping coefficient (kg/s)<br />
Air gap (µm²) (number<br />
Error (%)<br />
Simulation Analytical<br />
of holes)<br />
100x100 (100) 2.1x10 -6 2.6x10 -6 22.2<br />
200x200 (400) 9.4x10 -6 1.0x10 -6 10.7<br />
300x300 (900) 2.1x10 -5 2.3x10 -5 7.2<br />
2 µm 400x400 (1600) 3.9x10 -5 4.1x10 -5 5.6<br />
500x500 (2500) 6.2x10 -5 6.5x10 -5 4.6<br />
600x600 (3600) 9.0x10 -5 9.3x10 -5 3.9<br />
700x700 (4900) 1.2x10 -4 1.2x10 -4 3.5<br />
In the considered frequency range, the damping coefficient<br />
R airgap is constant and the agreement between simulated and<br />
calculated values varies from 22 % to 3.5 %.<br />
We have found similar results when using the analytical<br />
model of the squeeze-film for the microphone as when<br />
performing the FEA with CoventorWare. Fig. 6 shows the<br />
damping force simulated with CoventorWare and using (10).<br />
We have obtained R airgap = 5.4x10 -5 kg/s, which is within 10 %<br />
of the simulated value (6.1x10 -5 kg/s).<br />
Where μ is the fluid viscosity, h a is the air gap thickness and ω<br />
is the pulsation. If σ
A. Quasi-Static response<br />
In order to judge the microphone performance, we need to<br />
estimate the capacity variation ΔC induced by a known<br />
pressure. Supposing that we know the microphone frequency<br />
characteristics, we can do this estimation for DC values. The<br />
estimation of ΔC was done in three approaches. In all of them,<br />
we have compared the microphone capacity without a pressure<br />
with that with a pressure of 1 Pa applied on the diaphragm.<br />
The results obtained in the three cases are shown in Table IV.<br />
For the first approximation, we have used a non-perforated<br />
diaphragm (Fig. 7 (a)) because the electromechanical FEM<br />
simulations of the microphone with a perforated diaphragm<br />
have important computational time requirements, even if<br />
symmetry is assumed and a quarter of the model is simulated.<br />
Based on the predicted value of the pull-in voltage, V PI ,<br />
given by CoventorWare that is between 4.4 V and 4.5 V, we<br />
have decided to use the bias voltage V 0 of 1V. The<br />
corresponding capacitance variation is in Table IV (“FEM”<br />
line)<br />
We can verify these simulation results using an analytic<br />
approach. Indeed, if we consider just the maximum<br />
displacement of the diaphragm, we have the static equation:<br />
2<br />
k mem w max = F pressure + F electrostatic = PA +<br />
ε<br />
00<br />
AV<br />
−<br />
(14)<br />
2<br />
a max<br />
)<br />
And when P = 0 Pa, we have:<br />
max<br />
a<br />
P is the pressure acting on the diaphragm, A is the area of the<br />
diaphragm, V 0 is the bias voltage and B is a correction<br />
coefficient respecting the diaphragm deformation shape<br />
(Fig. 7 (b)). The pull-in effect occurs when the term on the left<br />
side of (15) reaches a maximum:<br />
∂<br />
∂w<br />
max<br />
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<br />
2<br />
ha<br />
[ B ] =−<br />
0)(<br />
whw<br />
w =<br />
max<br />
When w max = w Pull-in , V<br />
leads to<br />
a max<br />
Pull−in<br />
=<br />
pull<br />
3<br />
8<br />
memhk<br />
a<br />
27ε<br />
BA<br />
0<br />
(16)<br />
− in<br />
3B<br />
(17)<br />
Thanks to (17) and knowing the pull-in voltage by FEM<br />
simulations we can calculate the correction coefficient:<br />
B = 0.385. Once B is calculated, we can solve (14) with the<br />
Cardan method and then calculate the capacitance variation.<br />
Results are summarized in Table IV (“calculation (nonperforated<br />
diaphragm)” line).<br />
simulation. We can see that the calculation and simulation<br />
results are very close. There is a small capacitance deviation is<br />
due to the parasitic capacitance that is taken into account in<br />
the FEM simulation and is not accounted in the analytic<br />
model.<br />
Now, we applied this model for the perforated diaphragm<br />
considering the same value of B that was obtained for the nonperforated<br />
diaphragm. The pull-in voltage is V PI = 4.24 V and<br />
Table IV shows the variation capacitance (“Calculation<br />
(perforated diaphragm)” line).<br />
TABLE IV<br />
CAPACITANCE VARIATION<br />
Approach Conditions w max (nm)<br />
Capacitance<br />
(pF)<br />
FEM<br />
P = 0 Pa<br />
V 0 = 1 V<br />
V 0 = 1 V<br />
53 2.123<br />
P = 1 Pa<br />
133 2.140<br />
V 0 = 1 V<br />
Calculation P = 0 Pa<br />
(nonperforated<br />
51 1.361<br />
P =1 Pa<br />
diaphragm V 0 = 1 V<br />
131 1.378<br />
Calculation<br />
(perforated<br />
diaphragm)<br />
P = 0 Pa<br />
V 0 = 1 V<br />
P = 1 Pa<br />
V 0 = 1 V<br />
57 1.151<br />
147 1.167<br />
ΔC (fF)<br />
(2 B wh<br />
According to the model, the obtained capacitance variation<br />
is very similar as in the non-perforated case. This fact can be<br />
explained either by the estimated value of the correction<br />
2<br />
2 ε<br />
00<br />
AV coefficient B or by a compensation of the decreased stiffness<br />
B<br />
max<br />
)(<br />
=−(15)<br />
whw by a smaller area of a perforated diaphragm.<br />
2kmem<br />
B. Dynamic response<br />
We have used the microphone equivalent circuit, shown in<br />
Fig. 3, to predict the dynamic response of the microphone. The<br />
microphone dimensions considered in the simulations are<br />
shown in Table I. Fig. 8 shows the frequency range and the<br />
open circuit sensitivity of the microphone.<br />
17<br />
17<br />
16<br />
(a)<br />
(b)<br />
Fig. 7. Microphone structure used for simulation (a). Schematic effective<br />
displacement (b).<br />
Fig. 8. Simulated frequency range and open circuit sensitivity of the<br />
microphone.<br />
Fig. 9 shows the spectral density of the output noise voltage.<br />
From the spectral density, we can calculate the signal-to-noise<br />
ratio (SNR) of the considered microphone. The basic<br />
microphone characteristics are summarized in Table V.<br />
Note that k mem and A have been determined for the nonperforated<br />
diaphragm to have the same conditions as for the<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
<br />
313
TABLE V<br />
MICROPHONE CHARACTERISTICS<br />
Characteristics Value<br />
Resonant frequency 10 kHz<br />
Open circuit sensitivity -35 dB.V/Pa<br />
SNR<br />
68 dB<br />
Fig. 9. Simulated spectral density of the output noise voltage.<br />
IV. MICROPHONE FABRICATION<br />
The fabrication of the microphone is based on the AMS<br />
0.35 µm CMOS back-end process that encompass a<br />
passivation layer, four metal layers, three via layers and<br />
several silicon dioxide layers. The metal layers used for the<br />
microphone is M4 for the diaphragm and M2 for the fixed<br />
electrode. The silicon dioxide layer between M4 and M2 is a<br />
sacrificial layer which is removed by HF vapor etching.<br />
Fig. 10 shows the microphone before etching.<br />
Fig. 10. Microphone fabricated with the AMS 0.35 µm CMOS process<br />
before etching.<br />
V. CONCLUSIONS<br />
In this paper, we have proposed a simple model of a MEMS<br />
capacitive microphone to estimate its characteristics for audio<br />
applications. This lumped-parameters reduced model can still<br />
be improved, but it is useful for initial estimations.<br />
The structure of the microphone has been fabricated with a<br />
standard CMOS process (AMS 0.35 µm). A sacrificial etch of<br />
the silicon dioxide layers will be done in the near future to<br />
obtain the working microphone. If successful, it will be<br />
possible to integrate on the same chip the MEMS capacitive<br />
microphone and the electronic circuit.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
REFERENCES<br />
[1] Ristic, L., “CMOS technology: a base for micromachining”,<br />
Microelectronics Journal, Vol. 20, No. 1-2, 1989, pp. 153-169.<br />
[2] Parameswaran, M., Baltes, H. P., Ristic, L., Dhaded, A. C., and<br />
Robinson, A. M. “A new approach for the fabrication of micromechanical<br />
structures”, Sensors and Actuators A, Vol. 19, No. 3, 1989, pp. 289-307.<br />
[3] Rufer, L., Domingues, C., Mir S., Petrini, V., Jeannot, J.-C., Delobelle, P.,<br />
“A CMOS compatible ultrasonic transducer fabricated with deep reactive ion<br />
etching”, IEEE Journal of Microelectromechanical Systems, Vol. 15, No. 6,<br />
2006, pp. 1766-1776.<br />
[4] Maillya, F., Giani, A., Bonnota, R., Temple-Boyerb, P., Pascal-<br />
Delannoya, F., Foucarana, A. and Boyer, A., “Anemometer with hot platinum<br />
thin film”, Sensors and Actuators A, Vol. 94, No. 1-2, October 2001, pp. 32-<br />
38.<br />
[5] Fouladi, S., Bakri-Kassem, M., Mansour, R.R., “An Integrated Tunable<br />
Band-Pass Filter Using MEMS Parallel-Plate Variable Capacitors<br />
Implemented with 0.35 μm CMOS Technology”, IEEE/MTT-S International<br />
Microwave Symposium, 2007, pp. 505-508, 3-8 June 2007.<br />
[6] Dai, C. L., “A maskless wet etching silicon dioxide post CMOS process<br />
and its application”, Microelectronic Engineering, Vol. 83, 2006, pp. 2543-<br />
2550.<br />
[7] Ganji, B.A., Majlis, B.Y;, “Design and fabrication of a new MEMS<br />
capacitive microphone using perforated diaphragm”, Sensors and Actuators A,<br />
Vol. 149, 2009, pp. 29-37.<br />
[8] Rabinovich, V. L., Gupta, R. K. and Senturia, S. D., “The effect of<br />
Release-Etch Holes on the Electromechanical Behavior of MEMS structures”,<br />
International Conference on International Solid State Sensors and Actuators<br />
Conference, Vol. 2, June 1997, pp. 1125-1128.<br />
[9] Sharpe, W. N. Jr., Vaidyanathan, R., Yuan, B., Bao, G.. and Edwards, R.<br />
L., “Effect of etch holes on the mechanical properties of polysilicon” Journal<br />
of Vacuum Science & Technology B (Microelectronics and Nanometer<br />
Structures), Vol. 15, Sep 1997, pp. 1599-1603.<br />
[10] Bao, M., Yang, H., “Squeeze film air damping in MEMS”, Sensors and<br />
Actuators, Vol. 136, 2007, pp. 3-27.<br />
[11] Veijola, T., “Compact model for a MEM perforation cell with viscous,<br />
spring and inertial forces”, Microfluid Nanofluid, 2009, Vol.6, pp. 203-219.<br />
[12] Mohite, S. S., Kesari, H., Sonti, V. R., and Pratap, R., “Analytical<br />
solutions for the stiffness and damping coefficients of squeeze films in MEMS<br />
devices with perforated backplates”, Journal of Micromechanics and<br />
Microengineering, Vol. 15, 2005, pp. 2083-2092.<br />
[13] Mohite, S. S., Venkata, R., Sonti, V. R. and Pratap, R., “A compact<br />
Squeeze-Film Model Including Inertia, Compressibility, and Rarefaction<br />
Effects for perforated 3-D MEMS Structures”, Journal of<br />
Microelectromechanical systems, Vol. 17, No. 3, June 2008, pp. 709-723.<br />
[14] Homentcovschi, D. and Miles, R. N., “Analitycal model for viscous<br />
damping and the spring force for perforated planar microstructures acting at<br />
both audible and ultrasonic frequencies”, Journal of the Acoustical Society of<br />
America, Vol. 124, July 2008, pp. 175-181.<br />
[15] Bendali, A., Labedan, R., Dominique, F., Nerguizian, V., “Hole effects<br />
on RF MEMS Parallel Diaphragms Capacitors”, Canadian Conference on<br />
Electrical and Computer Engineering 2006, May 2006, pp. 2140-2143.<br />
©<strong>EDA</strong> <strong>Publishing</strong>/DTIP 2011<br />
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314<br />
ISBN:978-2-35500-013-3
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May 2011, Aix-en-Provence, France<br />
<br />
A Novel Integrated Solution for the Control and<br />
Diagnosis of Electrostatic MEMS Switches<br />
Carlo Trigona, Norbert Dumas, Laurent Latorre and Pascal Nouet<br />
LIRMM, University Montpellier II / CNRS<br />
161 rue Ada - 34095 Montpellier Cedex 5, France<br />
Abstract- The aim of this paper is to present a novel integrated<br />
solution for the control and diagnosis of electrostatic MEMS<br />
switches. A custom multi-channel integrated circuit (ASIC)<br />
has been designed and fabricated adopting a standard High-<br />
Voltage (HV) CMOS technology with a maximum operating<br />
voltage of 50V. Each channel is composed of a HV driver to<br />
actuate electrostatic switches and a diagnosis element to<br />
monitor the movement of the beam. This control circuit is<br />
particularly interesting for systems based on a large array of<br />
MEMS switches with a certain level of redundancy or fault<br />
tolerance, an active reflect array antenna in our case. The<br />
diagnosis principle has been modeled, simulated and an<br />
experimental campaign validates the principle with real<br />
actuators.<br />
I. INTRODUCTION<br />
During the past decade, RF Micro Electro Mechanical<br />
Systems (RF-MEMS) have been reported in several areas of<br />
engineering and science. The fabrication of MEMS for RF<br />
integrated circuits has seen a rapid rate of expansion within<br />
a broad range of applications such as: resonators [1],<br />
tunable filters [2], radar sensors [3] array, reconfigurable<br />
antennas [4] and RF switches [5].<br />
In particular, reflect array antennas based on phase<br />
shifters to tune the reflecting angle of a wave have attracted<br />
particular interest for applications such as communication<br />
satellite. Main advantages are their small size, high<br />
performance, reduced power consumption and lightness.<br />
Such communication systems are composed of thousands of<br />
RF switches embedded into a single panel and<br />
electrostatically actuated, although magnetic, thermal or<br />
even gas-based forces are alternative solutions to move<br />
micromachined structures. Due to the large number of<br />
control signals (one for each actuators), control architecture<br />
must be embedded in the panel. Basically, the adoption of a<br />
distributed network of ASIC devices presents several<br />
advantages such as small overall dimension, integration<br />
with the phase-shift panel, short distance from the RF<br />
switch to the control driver and reduced interconnection<br />
complexity within the panel.<br />
Due to the high number of RF switches, the phase-shift<br />
panel implements a certain degree of redundancy (i.e.<br />
achieving the same phase shift with various combination of<br />
switches) [6]. Fault-tolerance can then be easily<br />
implemented by monitoring the state of each RF MEMS,<br />
detecting non-working MEMS and configuring the panel<br />
accordingly.<br />
Several diagnostic approaches have been considered:<br />
optical monitoring [7], electromagnetic [8], resistive [9], or<br />
capacitive sensing [10]. These approaches imply complex<br />
diagnosis principles, additive sensing elements and/or<br />
parasitic-capacitance dependent solutions.<br />
The presented approach allows on-line testing of MEMS<br />
switches and exhibits several attractive features:<br />
• fully integrated within the driver,<br />
• active during each pull-in and pull-off events,<br />
• tolerant to large parasitic capacitances.<br />
The two last points are the real innovation of this driver<br />
compared to previously proposed drivers by the same<br />
authors [11]. The diagnosis circuitry can cope with large<br />
parasitic capacitances and can also detect the MEMS<br />
switch-off.<br />
This paper first reports the working principle of the<br />
proposed smart driver with a particular emphasis on<br />
diagnosis. Design and simulation are reported in a second<br />
section while experimental results finally demonstrate that<br />
we can detect both pull-in and pull-off events of the<br />
actuated beam thus demonstrating the suitability of the<br />
proposed solution.<br />
II.<br />
WORKING PRINCIPLE<br />
The diagnosis principle has been derived from an<br />
external test solution proposed in [12]. Here, we propose to<br />
associate an integrated diagnosis unit to each control<br />
channel. It is then composed with a control unit itself and an<br />
add-on circuitry for diagnosis (Figure 1).<br />
n = 10<br />
n∙I act<br />
VDDA<br />
Diagnostic unit<br />
VDDA<br />
n∙I act<br />
I discharge C ramp<br />
Diag1 Diag2<br />
VDD_HV<br />
I1 V th1<br />
I2 V th2<br />
I charge<br />
R1<br />
C int<br />
V ref<br />
Buffer OUT Control unit<br />
IN<br />
I act R s<br />
MEMS<br />
C act<br />
Fig. 1. Schematic diagram of a control channel with embedded diagnosis.<br />
C p<br />
315
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May 2011, Aix-en-Provence, France<br />
The basic idea consists to use a slow ramp-shaped signal<br />
<br />
III. DESIGN AND SIMULATIONS<br />
to control the electrostatic actuator and to analyze the so-<br />
by the driver to<br />
obtained actuation current (I act ) deliveredd<br />
Smart drivers have been designed and fabricated using a<br />
the actuator. To generate the driver HV control signal, the HV 0.35µm CMOS technology<br />
from Austria MicroSystems<br />
digital input (IN) is used to control the charge and discharge<br />
(AMS). This process tolerates voltage drop across transistor<br />
of an integrated capacitance (C ramp ) with two constant<br />
channels (V DS ) up to 50V. Figure 3 (respectively 4) shows the<br />
current generators (I charge and I discharge ). As a result, the<br />
layout (resp. an image) of a control channel. Zone (1)<br />
current buffer output (OUT) delivers a HV signal with a<br />
represents the current source of the ramp module, (2) is a<br />
constant voltage slope. The load is represented by the<br />
current mirror copying a fraction of the current from the<br />
MEMS device, that can be assumed to be a variable source and the circuit to control<br />
the sign of the ramp (rise or<br />
capacitance (C act ), the parasitic contribution which is<br />
fall), (3) is the integrating MIM capacitor, C ramp, used to<br />
capacitive (C p ), and a serial resistance (R s s) coming from the<br />
generate a positive or negative voltage ramp, (4) is the current<br />
routing or wire bonding. It is worth noting that both<br />
source for common drain buffer (5), (6) and (7) are the NMOS<br />
capacitors are connected in parallel and that the parasitic<br />
and PMOS mirror respectively which copy and amplify the<br />
capacitance is large compared to the actuator capacitance. actuation current for the diagnosis, while (8) are the voltage<br />
The diagnosis section copies the actuation current limiters that prevent V I1 and V I2 to go beyond the low voltage<br />
through a serial resistance (R 1 ) to obtain a voltage drop<br />
supply (V DDA = 3.3V). Both rise and fall slopes of the ramp<br />
proportional to the actuation current:<br />
generator can be adjusted thanks<br />
to a pair of external resistors<br />
to allow a parametric study.<br />
V<br />
∂Vact<br />
= n ⋅ R1<br />
⋅ I<br />
act<br />
= n ⋅ R1<br />
⋅ C<br />
p<br />
+ n ⋅ R ⋅V<br />
∂t<br />
I1<br />
1<br />
actt<br />
∂<br />
∂t<br />
C act<br />
= α + β<br />
(1)<br />
This voltage is composed of two terms; the first one (α)<br />
represents the contribution of the parasitic capacitance<br />
(assuming C act is small compared to C p p) and is constant<br />
during the charge (respectively the discharge) of the<br />
actuator, while the second (β) is proportional to capacitance<br />
variations (assuming C p is bias-independent). The latter<br />
term is null except during pull-in or pull-off events. Both<br />
events correspond to a rapid change in the capacitance thus<br />
producing a current peak. Finally, if the ramp is sufficiently<br />
slow, pull-in and pull-off current spikes can be identified<br />
out of the constant current contribution due to α. Finally, a<br />
comparator is used to detect the spikes and to integrate the<br />
actuation current (with capacitor C int ) for quantitative<br />
evaluations. It allows measuring the total variation of<br />
capacitance during the pull-in or the pull-ofoutput). Figure 2 illustrates the main signals: the absence of<br />
events (Diag2<br />
the pull-in and pull-off spikes implies a non-working<br />
MEMS condition. From this diagnostic<br />
“go-nogo”<br />
information (Diag1 output) and the knownn previous state of<br />
the switch the new state of the beam can be deduced. In this<br />
paper, we concentrate on the analysis of the voltage drop in<br />
R 1 (I 1 signal).<br />
(7)<br />
(5)<br />
(2)<br />
(1) (1) (4)<br />
(8)<br />
(3)<br />
(6) (7)<br />
Fig. 3. Layout of a smart-driver channel with a 760x550 µm 2 area occupied<br />
in a HV 0.35µm CMOS technology.<br />
Fig. 2. Main signals of the diagnosis unit: ramp-shaped actuation voltage<br />
V act and actuation current image I 1. For a functional beam, two spikes can<br />
be discriminated through the comparison thresholds.<br />
Fig. 4. Microscope image<br />
of a smart driver channel.<br />
<br />
<br />
316
A rise time of 250ms and a fall time of 500ms are<br />
achievable using 5MΩ resistors. To reduce the total size of the<br />
driver array, the current sources (1) and (4) can be shared<br />
between several drivers. In the fabricated prototype, R 1 , C int and<br />
the comparators for diagnosis are not integrated on-chip to<br />
allow more flexibility. The value of R 1 can vary from 10kΩ to<br />
2MΩ depending on the time response and the capacitance of<br />
the monitored electrostatic actuator. Integration of these<br />
components would not induce a significant area overhead. The<br />
value of C int should be calculated such as: C int· V I2 =<br />
∆C act·n·V act . Considering an actuation voltage of 50V and<br />
expecting a final voltage V I2 = 1V results in an integration<br />
capacitance 500 times higher than the variation of capacitance<br />
to observe, ∆C act . Therefore the capacitance value is typically<br />
in the order of a few tens of picofarads and could be more<br />
problematic to integrate. However, the integration is not<br />
necessary to obtain a binary go-nogo diagnosis. Alternatively<br />
the factor n could be reduced to adapt this solution to the<br />
MEMS switch under test. The static current consumption is<br />
relatively low and makes this driving and diagnosis solution<br />
suitable for controlling a large number of MEMS. It is mainly<br />
due to the current source that biases the common drain buffer<br />
and which consumes 2×366nA at 50V. If the source is shared<br />
by several buffers it reduces the overall consumption per<br />
channel. This small current consumption implies a limitation of<br />
366nA in terms of maximal current that can be sink or source<br />
from the current buffer and an equivalent output resistance of<br />
250kΩ. It should be noted at this point that this output<br />
resistance will have an influence on the cut-off frequency due<br />
to the parasitic capacitance C p . Assuming C p = 1pF, it results in<br />
a cut-off frequency of 640kHz.<br />
The MEMS switch that will be considered in the following<br />
is a gold cantilever beam based on a custom technology with a<br />
pull-in voltage up to 44V [13]. It has been simulated using the<br />
library of beam model provided by CoventorWare ® . Figure 5<br />
illustrates the behavior of the switch when a voltage is applied<br />
on the bottom actuation electrode. A first pull-in occurs<br />
around 29V while a second one occurs around 31V. After the<br />
first pull-in, the beam touches the signal electrode and the<br />
switch is considered to be in the ON-state. The variation of the<br />
actuation capacitance due to this pull-in is in the order of<br />
170fF. The second pull-in, which is not desired, produces a<br />
much larger variation of capacitance (> 6pF) and thus can be<br />
easily experimentally observed as we will see later-on.<br />
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<br />
IV. EXPERIMENTAL RESULTS<br />
The diagnosis approach is investigated and validated in this<br />
section. The designed ASIC and the embedded diagnostic<br />
architecture has been bonded on a test board, as shown in<br />
figure 6. It includes two MEMS prototypes, a switch selector<br />
and resistors R S and R 1 .<br />
It must be noted that for an easy observation of the signal, a<br />
value of 1.2MΩ has been chosen for the resistance R 1 , while<br />
the serial resistance connected with the MEMS (R S )<br />
corresponds to 1kΩ. The high voltage supply is set to the<br />
maximal value 50V. The ASIC prototype has been finally<br />
protected from light by using a cap.<br />
A square waveform having a frequency of 0.5Hz and a<br />
3.3V amplitude has been applied as an input signal while the<br />
measurement has been conducted observing the voltage V I1.<br />
During the experiment a noise level of about 120mV (attributed<br />
to the environment) has been recorded while a main spike<br />
saturating at about 2.4V has been measured as consequence of<br />
the beam movement (figure 7a). This value is due to the<br />
internal voltage limitation of the ASIC. This pull-in spike has<br />
been identified in presence of a working MEMS; a microscope<br />
analysis has been conducted to confirm the diagnosis.<br />
Furthermore, it disappears when the high voltage supply<br />
(V DD_HV ) is lower than 44V, indicating that the MEMS switch<br />
is no longer actuated when the voltage is too low. It should be<br />
noted that the time duration of the pull-in (i.e. the width of the<br />
spike) is about 200µs as previously observed in [11] for a<br />
similar switch. Comparing the response time of the switch to<br />
the cut-off frequency previously calculated in section III<br />
(640kHz), we can conclude that the bandwidth limitation is not<br />
an issue here.<br />
As expected this spike is easily detected by a simple<br />
comparator. Putting a resistor R 1 = 220kΩ reduces the<br />
amplitude of the spike and makes it possible to integrate the<br />
voltage in order to calculate the variation of capacitance. A<br />
value of 760fF is found instead of more than 6pF found with<br />
the simulation. Authors assume that this discrepancy is due, in<br />
order of importance, to the surface roughness of the oxide<br />
coating the electrode, and to the actual dimensions and the<br />
actual initial bending of the switch after fabrication. The<br />
simulation accuracy may also be important because of the large<br />
displacement of the beam and the difficulties to model contacts.<br />
Vact = 0V<br />
Vact = 29V<br />
Vact = 31V<br />
Fig. 5. Simulation of the beam actuation using CoventorWare ® . The z-<br />
dimensions have been multiplied by a factor 10. The unit of the<br />
displacement scale on the left is in micrometer.<br />
Fig. 6. Test board for RF switch actuation and diagnosis.<br />
317
Figure 7b shows the effect of a parasitic load C p = 10pF<br />
connected in parallel to the MEMS device. First, one can notice<br />
that the graph presents a two pull-in spikes, which confirms the<br />
physical simulation presented in section III, and only one pulloff<br />
spike. The presented waveform has been averaged over 32<br />
acquisitions to reduce the noise. It also strongly reduces the<br />
spikes amplitude because of the fluctuation of the pull-in or<br />
pull-off times and the triggering accuracy. Results evince the<br />
presence of spikes also with a large parasitic capacitance. The<br />
pull-in voltage measured are 36.5V for spike #1, 41.8V for<br />
spike #2 and the pull-off voltage is 19.2V (spike #3).<br />
Regarding the effect of the parasitic capacitance we can<br />
observe a stable deviation from the reference voltage, i.e.<br />
1.37V, during the ramp up or the ramp down. The results are in<br />
accordance with the theoretical values. For the ramp up the<br />
deviation should be R 1·n·C p·V DD_HV /∆t up = 4.4mV and for the<br />
ramp down it should be -R 1·n·C p·V DD_HV /∆t down = -2.2mV. The<br />
sensitivity to a parasitic capacitance is negligible for spike #2<br />
which is about 1V high (for the same condition R1 = 220kΩ)<br />
while it is problematic for spike #1 and spike #3 which are<br />
16mV and 60mV high respectively. Several parasitic<br />
capacitance values have been investigated to validate the<br />
diagnostic approach. A maximum capacitive load of about<br />
100pF has been experimentally observed.<br />
V I1 (V)<br />
3,0<br />
2,5<br />
2,0<br />
1,5<br />
1,0<br />
0,5<br />
0,0<br />
V DD_HV =<br />
46V;47V;48V;50V<br />
0,0024 0,0025 0,0026 0,0027 0,0028 0,0029<br />
Time (s)<br />
(a)<br />
V DD_HV = 44V<br />
Spike #2<br />
Pull-in<br />
1,42<br />
1,41<br />
Spike #1<br />
Spike #2<br />
1,40<br />
Pull-in<br />
Pull-in<br />
1,39<br />
1,38<br />
1,37<br />
1,36<br />
Spike #3 Pull-off<br />
1,35<br />
Ramp up No ramp Ramp down<br />
1,34<br />
0 0,2 0,4 0,6 0,8 1<br />
Time (s)<br />
(b)<br />
Fig. 7. Evolution of voltage across R 1: (a) close-up view of pull-in event<br />
(spike #2) for different values of HV power supply and R 1 = 1.2MΩ.<br />
Below 44 V, pull-in spike clearly disappears. (b) Averaged signal (32<br />
acquisitions of 1s are averaged) in presence of a 10pF additional parasitic<br />
capacitance (R 1 = 220kΩ, V DD_HV=50V).<br />
V I1 (V)<br />
averaged<br />
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May 2011, Aix-en-Provence, France<br />
<br />
V. CONCLUSIONS<br />
In this paper, a CMOS HV driver with integrated diagnosis<br />
capabilities for MEMS electrostatic actuators has been<br />
presented. Possible applications concern a wide range of<br />
domain where an array of such actuators could be used (a<br />
reflect array antenna in our case). The paper focuses on the<br />
diagnosis circuitry that has been modeled, simulated and<br />
experimentally validated.<br />
The presented diagnosis approach addresses on-line test of<br />
MEMS switches to detect capacitance variations during pull-in<br />
and pull-off events and to reject parasitic capacitance up to<br />
100pF. The proposed method demonstrates an improvement in<br />
terms of rejection ratio (α/β) of parasitic capacitance of 40 dB<br />
considering the main pull-in event, with respect to the results<br />
obtained in [11]. Furthermore the presented method is able to<br />
detect the pull-off spike thus allowing to determine if a failing<br />
device is stuck down.<br />
In addition, electrical characterization gave some<br />
interesting information on the complete dynamic behavior of<br />
the switch.<br />
ACKNOWLEDGMENT<br />
This work has been carried out under the French ANR<br />
project R3MEMS.<br />
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IEEE MTT-S International, pp. 300 - 303, May 23-28, 2010.<br />
[10] A. Fruehling, A. Khater, B. Jung, D. Peroulis, “Real-time<br />
monitoring of contact behavior of RF MEMS switches with a very<br />
low power CMOS capacitive sensor interface”, MEMS 2010, pp.<br />
775 – 778, January 24-28, 2010.<br />
[11] N. Dumas, L. Latorre, F. Mailly, P. Nouet, “Smart drivers for<br />
online diagnosis of electrostatic MEMS actuators”, Mixed-Signals,<br />
Sensors and Systems Test Workshop (IMS3TW, pp. 1-6, June 7-9<br />
2010.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
[12] B. Caillard, Y. Mita, Y. Fukuta, T. Shibata, and H. Fujita, “A<br />
highly simple failure detection method for electrostatic<br />
microactuators: application to automatic testing and accelerated<br />
lifetime estimation,” IEEE Transactions on Semiconductor<br />
Manufacturing, vol. 19, no. 1, pp. 35–42, 2006.<br />
[13] C. Villeneuve, S. Aouba, M. Dilhan, D. Bourrier, P. Pons, R Plana,<br />
“Low stressed gradient in gold micromachined cantilevers”, 20th<br />
MicroMechanics Europe workshop, September 20-22, 2009.<br />
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<br />
An ultra low power temperature sensor for smart<br />
packaging monitoring<br />
Souha Hacine, Frederick Mailly, Norbert Dumas, Laurent Latorre, Pascal Nouet<br />
University Montpellier 2 / CNRS – LIRMM 161, Rue Ada – 34392 Montpellier Cedex 5 -France<br />
Tel: +33 467.418.665 – Fax: +33 467.418.500 – latorre@lirmm.fr<br />
Abstract- The work presented in this paper concerns the<br />
design and the characterization of a compact and low-power<br />
temperature CMOS sensor as part of an environmental<br />
monitoring scheme for MEMS devices. The proposed sensor is<br />
based on the TCR of available polysilicon resistances and a<br />
previously reported architecture (Active Bridge). It is<br />
composed of a single differential stage that performs both the<br />
resistance biasing and the amplification with the same 2µA<br />
current under 3.3V. Moreover, the Active Bridge offers the<br />
opportunity to implement a very compact Σ∆ modulator, thus<br />
converting the temperature input into digital information.<br />
I. INTRODUCTION<br />
This paper addresses the design of a low-power CMOS<br />
temperature sensor. This study is part of a larger project that<br />
investigates solutions to continuously monitor the<br />
environmental conditions (temperature, pressure, humidity) of<br />
several MEMS sensors packaged together in order to perform<br />
on-line calibration or to anticipate failures [1]. Because the<br />
environment monitoring is not the primary function of the<br />
system, constraints in term of area overhead and power<br />
consumption of dedicated sensors are very tight.<br />
For temperature sensing, an obvious approach is to make use<br />
of thermistors fabricated with standard CMOS materials (e.g.<br />
polysilicon). Still, the conditioning of resistive transducers, by<br />
mean of the well-known Wheatstone bridge is not always a<br />
good solution to minimize the power consumption. Our idea is<br />
to investigate the use of a previously introduced original<br />
structure called “Active Bridge” [2, 3] that combines into a<br />
single stage the resistance biasing and the amplification. Such<br />
approach leads to a small and low-power temperature sensor.<br />
In addition, due to its intrinsic high output resistance, this<br />
structure behaves like an integrator in presence of a capacitive<br />
load. This property offers an excellent opportunity to build a<br />
Σ∆ modulator with the added-value of a digital output.<br />
Therefore, the work presented here first details the design of<br />
an analog temperature sensor based on the Active Bridge.<br />
Then, a Σ∆ modulation scheme is proposed to provide a digital<br />
output to the Active Bridge. Finally, experimental results are<br />
presented and a comparison between the Active Bridge<br />
performances and traditional Wheatstone bridge with same<br />
resistive transducers is made.<br />
II.<br />
TEMPERATURE SENSOR<br />
A. Principle<br />
The ultra-low-power detection of a resistor variation without<br />
noise deterioration is the challenge that we are aiming with the<br />
Active Bridge. The main idea behind this structure is to use the<br />
same current to bias the sensing elements (resistors) and to<br />
achieve the required amplification.<br />
The structure presented in this study is self-biased and does<br />
not require any additional circuitry or reference voltage thus<br />
keeping the power consumption and the silicon area cost very<br />
low. The proposed circuit (Fig.1) uses four sensitive resistors<br />
with two different temperature coefficients (TCR). These TCR<br />
are of opposite signs and have been chosen under<br />
technological constraints: the available materials to make<br />
sensitive resistors with high resistance values and reasonable<br />
area are polyh and poly2 in the AMS (Austria Microsystems)<br />
0.35µm CMOS technology. Their TCR are around 10 -3 /K.<br />
<br />
Out-<br />
<br />
T3<br />
T4<br />
Vdd<br />
Fig. 1 Temperature sensor topology based on the Active Bridge<br />
B. Design of the temperature sensor<br />
In our application, we target a 2µA total current consumption<br />
for the sensing stage. In order to meet this constraint, as well<br />
as good resolution, a (V GS -V T ) equal to 0.4V is chosen. In<br />
addition, this value allows us to reduce the process-induced<br />
mismatches between identically designed devices.<br />
T1<br />
T2<br />
<br />
Out+<br />
<br />
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Besides, in order to reduce the flicker noise’s<br />
corner frequency<br />
<br />
To further validate the proposed circuit, Monte Carlo<br />
and to obtain a white noise floor as close as possible to the simulations have been performed (Fig. 3) using both processes<br />
theoretical resistors’ white noise, the designed PMOS and<br />
(wafer to wafer) and mismatches (intra die) variations (at<br />
NMOS have a large area with a greater length than width.<br />
27°C). Due to its high sensitivity (3.4V/°C), the temperature<br />
sensor amplifies mismatch-induced induced offset and the sensor<br />
<br />
<br />
<br />
<br />
<br />
output always saturates. Due to the random behavior of<br />
mismatches, nearly half of the circuits saturates to V dd while<br />
the other half saturates to V ss . Consequently, the sensor cannot<br />
From the (V GS -V T ) value, the resistor’s value is calculated by<br />
be used in open loop and a digital or analog feedback is<br />
applying a simple Ohm’s law on the Active bridge with<br />
required to control (i.e. center) the operating point.<br />
V dd =3.3V:<br />
<br />
<br />
<br />
With |V GSp |=0.4+V tp (V tp =0.7V) and V GSn =0.4+V tn (V tn =0.5V).<br />
So, the resistors white noise, that gives the intrinsic white<br />
noise of the sensor, is:<br />
=<br />
Where k is the Boltzmann constant and T is the ambient<br />
temperature (27°C in equation 3).<br />
Fig. 2 presents a noise simulation (Cadence<br />
® ) of the<br />
architecture dimensioned above. The flicker noise’s corner<br />
frequency is below 10Hz and the total white noise floor is<br />
close to the circuit intrinsic white noise given by the resistors<br />
white noise.<br />
<br />
<br />
<br />
<br />
/Hz<br />
<br />
<br />
<br />
<br />
<br />
229.3nV/Hz<br />
Fig. 2 Noise simulation results (for T=27°C) of the temperature sensor<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
Fig. 3 Monte Carlo Simulation of the operating point of the differential Active<br />
Bridge (DC output voltage for T=27°C)<br />
<br />
(2)<br />
(3)<br />
<br />
<br />
III. IMPLEMENTATION IN A MODULATOR<br />
A. Principle<br />
The modulators are very efficient for converting analog<br />
output signals resulting from sensor variations with low<br />
bandwidths. The interest of this kind of converter is its high<br />
output resolution.<br />
Since thermal variations lead to resistive variations in our<br />
temperature sensor, we opted for a resistive feedback in the<br />
modulator. The principle is illustrated in Fig. 4. A<br />
temperature variation results in a change in the resistors value<br />
(+R). The difference between this variation R and a<br />
feedback resistor is integrated and compared to zero. The<br />
comparator output is then sampled at the modulator clock<br />
frequency and the feedback resistor is added or subtracted to<br />
the sensing resistors according to the sign of the difference’s<br />
average.<br />
T R Output<br />
Sensor + -<br />
<br />
Clk<br />
+R fb DAC<br />
Fig. 4 Block diagram of the Σ∆ modulator<br />
B. Implementation<br />
Fig. 5 presents the transistors-level implementation of the<br />
modulator. The Active Bridge converts the thermal<br />
variations of the sensitive resistors in voltage variations. The<br />
integrator is implemented as a simple low pass filter composed<br />
of the bridge high output resistance (1.9G) and a 22pF<br />
external capacitance.<br />
The sampled comparator block is a Flip-Flop register with a<br />
clock frequency of 16 kHz. When the Flip-Flop output<br />
toggles, S 0 and S 1 are controlled so that V dd switches from one<br />
side of the feedback resistor to the other. This implements the<br />
digital to analog conversion and allows balancing the two<br />
branches of the structure by adding the feedback resistor in<br />
one branch and subtracting it from the other. Symmetrically,<br />
G nd switches from one side to the other at the bottom of the<br />
structure. The value of the feedback resistors is calculated<br />
from the maximum difference ference between the sensitive resistors<br />
321
over the temperature range (i.e. the full-scale). In our case, it is<br />
equal to about 10%.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
noise floor is evaluated over the 1-30 Hz range and is equal to<br />
2.59m°C/Hz.<br />
Vdd<br />
S 0<br />
S 1<br />
R fb<br />
(4)<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
T3<br />
<br />
T4<br />
T2 C<br />
<br />
<br />
R fb<br />
S 1<br />
S 0<br />
IV.<br />
T1<br />
<br />
Out+<br />
Fig. 5 Implementation of the temperature sensor based<br />
on a Σ∆ modulator topology.<br />
EXPERIMENTAL RESULTS<br />
Comparator<br />
<br />
D Q<br />
<br />
<br />
CK<br />
Fig. 6 presents both simulated and experimental output of<br />
the modulator as a function of temperature.<br />
The simulated output represents the ratio of ‘ones’ in the<br />
bit-stream, within 1024 clock periods for each temperature,<br />
extracted from a transient analysis. Experimentally, the same<br />
ratio is calculated over a 10s time window.<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
°<br />
Fig. 6 Normalized digital output versus temperature<br />
In order to determine the resolution of the sensor, a spectral<br />
analysis of the output bit-steam is made at 16 kHz clock<br />
frequency (Fig. 7). The classical noise shaping of a 1 st<br />
order Σ∆ modulator is well observed with an increase of the<br />
noise at high frequencies. From this spectrum, the average<br />
<br />
<br />
Fig. 7 Bit-stream spectral density (F clk=16 kHz)<br />
Now, let us focus on the linearity performance obtained<br />
with the Active Bridge in a Σ∆ modulator configuration, and<br />
compare it to the one of the Wheatstone bridge (Fig. 8). The<br />
Wheatstone bridge remains the most common approach to<br />
condition resistive sensors [4]. This architecture introduces a<br />
major tradeoff between resolution and power consumption [5].<br />
On the one hand, the smaller the sensor resistor is, the higher<br />
the current in the bridge will be. On the other hand, the higher<br />
the sensor resistor is, the higher the noise floor will be. The<br />
performances are compared with identical resistors and thus<br />
similar silicon surface (MOS transistor surfaces are<br />
negligible). Note that in our case, temperature sensor is highly<br />
sensitive; therefore the SNR is not really an issue.<br />
<br />
<br />
<br />
Vdd<br />
<br />
<br />
<br />
Fig. 8 Wheatstone bridge topology with Rpolyh and Rpoly2<br />
Fig. 9 presents the simulated non linearity for both<br />
architectures. These results are compared to the non linearity<br />
experimentally observed for the Σ∆ modulator. It shows that<br />
the calculated nonlinearities (as a percentage of the full-scale<br />
temperature range, i.e. 140°C) are identical for both<br />
architectures. It can be deduced that this non linearity comes<br />
from the sensing elements themselves and not from the<br />
architectures. Besides, the maximum is reached at both<br />
extremities of the temperature range and the shape of the nonlinearity<br />
reasonably allows suspecting a quadratic effect of the<br />
temperature on the resistance. Regarding experimental non<br />
linearity, we can conclude that the experimental set-up of<br />
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11-13 <br />
May 2011, Aix-en-Provence, France<br />
these preliminary results must be improved but obtained<br />
<br />
results are close to the simulated one.<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
TABLE I: Comparison chart between the<br />
Wheatstone bridge and the Active Bridge performances<br />
Active bridge with<br />
Σ∆ modulator<br />
Wheatstone<br />
bridge<br />
Consumption 1 (µA) 2 5.3<br />
v n (nV/Hz) 101.5<br />
Sensitivity (mV/°C) 2.14<br />
Resolution (°C/Hz) 2.59m 47.42µ<br />
Non linearity 6.9° 7.5°<br />
1<br />
Consumption of the sensor only.<br />
<br />
Fig. 9 Linearity study of the two circuits from -40 to 100°C<br />
To confirm the relationship between linearity and the<br />
resistance dependence to temperature, we have studied the<br />
linearity of the resistance of both materials used to implement<br />
resistance temperature sensors (Fig. 10). It confirms that the<br />
non linearity is basically due to the thermal quadratic terms of<br />
both materials (polyh and poly2).<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
°<br />
<br />
<br />
Fig. 10 Intrinsic linearity with temperature of resistors.<br />
Finally, Table I summarizes performances obtained at 27°C,<br />
for both studied architectures: Wheatstone bridge and Active<br />
Bridge in a Σ∆ modulator. Due to its one-bit digital output, it<br />
is impossible to calculate output noise or sensitivity for the Σ∆<br />
modulator. These notions are only meaningful after<br />
decimation filtering. Noise that has been determined<br />
previously for the Σ∆ modulator output (Fig. 7.) corresponds<br />
to a quantization noise that allows determining a resolution<br />
<br />
proportional to <br />
[6], where f c is the cut-off frequency of<br />
<br />
the low-pass filter implementing the integrator and f ck is the<br />
clock frequency of the modulator. It is then possible to freely<br />
adjust these parameters to reach the targeted resolution down<br />
to the limit fixed by the intrinsic noise of the Active Bridge<br />
(Eq. 3).<br />
The main advantage of the Active Bridge is its ability to<br />
reduce the power consumption while providing a digital output<br />
when used in a Σ∆ modulator.<br />
V. CONCLUSION<br />
This paper presents a temperature sensor with an innovative<br />
structure for the signal conditioning of resistance. A<br />
comparison between the traditional Wheatstone bridge and the<br />
Active Bridge has been made that demonstrates the main<br />
advantages of the proposed solution are its very low power<br />
consumption and its capacity to provide directly a one-bit<br />
digital output. In particular, the high output resistance of the<br />
Active Bridge allows implementing a modulator with very<br />
few additional parts.<br />
ACKNOWLEDGMENT<br />
This work was supported by ANR, the French National<br />
Research Agency, under the project MIDISPPI.<br />
REFERENCES<br />
[1] F. Mailly, N. Dumas, N. Pous, L. Latorre, O. Garel, E. Martincic,<br />
F. Verjus, C. Pellet, E. Dufour-Gergam, P. Nouet. Original<br />
Research Article Sensors and Actuators A: Physical, Volume 156,<br />
Issue 1, November 2009, Pages 201-207<br />
[2] Boujamaa E. M., Dumas N., Mailly F., Latorre L., Nouet P. «The<br />
Active Bridge: an Alternative to the Wheatstone Bridge for<br />
Efficient Conditioning of Resistive MEMS Sensors » Design, Test,<br />
Integration of MEMS/MOEMS (DTIP’09), Avril 2009.<br />
[3] Boujamaa E. M., Alandry B., Hassine S., Mailly F., Latorre L.,<br />
Nouet P. « A Low Power Interface Circuit for Resistive Sensors<br />
with Digital Offset Compensation » IEEE International<br />
Symposium on Circuits and Systems (ISCAS’ 10), Juin 2010.<br />
[4] Luc, Hébrard, Jean-Batiste, Kammerer and Francis, Braun. « A<br />
chopper Stabilized Biasing Circuit Suitable for Cascaded<br />
Wheatstone-Bridge-Like Sensors ». IEEE Transaction on Circuits<br />
and Systems. August 8, 2005, Vol. 52, 8.<br />
[5] Apinunt, Thanachayanont and Suttisak, Sangtong. « Low-Voltage<br />
Current-sensing CMOS Interface Circuit for Piezo-Resistive<br />
Pressure Sensor ». ETRI Journal. February 2007, Vol. 29, 1.<br />
[6] O. Leman, F. Mailly, L. Latorre, P. nouet, «A wide bandwith, wide<br />
dynamic-range thermal Σ∆ architecture for convective<br />
accelerometers», 8 th IEEE Conference on Sensors (SENSORS’09),<br />
October 2009.<br />
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<br />
Accurate Thermal Characterization of Power<br />
Semiconductor Packages by Thermal Simulation and<br />
Measurements<br />
Andras Vass-Varnai (1,2) , Robin Bornoff (2) , Sandor Ress (1,2) , Zoltan Sarkany (2) , Sandor Hodossy (1) , Marta Rencz (1,2)<br />
Budapest University of Technology and Economics (1)<br />
Mentor Graphics Mechanical Analysis Division (2)<br />
@eet.bme.hu<br />
@mentor.com<br />
In this paper the possibility of generating a compact thermal<br />
model based on thermal transient measurements is discussed and<br />
evaluated. A case study of a power diode in a cylindrical-shaped<br />
copper package is shown. The detailed model of the package is<br />
built and simulated in a CFD based thermal simulator software.<br />
The measurement results are compared to the results of the<br />
simulations and after some model refinement we found good<br />
agreement. The compact model of the package is also identified<br />
based on the structure functions generated from the real<br />
measurement. The case-node is determined using the standard<br />
dual-interface method. The resulting compact model has proven<br />
to be a perfect representation of the real package as the structure<br />
functions generated based on measurements and corresponding<br />
simulation results match perfectly.<br />
I. INTRODUCTION<br />
As the functionality of thermal simulators gets more and<br />
more complex, measurement techniques also improve.<br />
Thermal engineers face an increasingly difficult task to<br />
make the right selection from the existing tools. Beside this<br />
problem the precise determination of thermal performance<br />
indicators such as RthJC or RthJB is becoming more and<br />
more difficult as the package geometries become more<br />
complex. The thermal characterization of novel power<br />
packages hosting a number of dies is a major issue where<br />
the standard definitions cannot be applied anymore [1,2].<br />
The answer to these challenges may lie in a combined<br />
measurement and simulation approach. Measurements yield<br />
a structure description of materials having different<br />
conductivities; simulation gives the clue as to what certain<br />
sections in the measured structure correspond to. TIM<br />
materials are very difficult to model, as neither their<br />
conductivity nor their thickness can be determined with<br />
high accuracy even by the designer of a given package.<br />
Well planned thermal measurements are suitable tools to<br />
measure the in-situ resistance of these materials so that they<br />
can be later on used for accurate model creation. Another<br />
example may be the junction-to-case thermal resistance<br />
measurement of power packages where the single R thJC<br />
value obtained by measurements may not be a perfect<br />
indicator due to the complex heat-spreading path. In such<br />
cases steady-state simulations may show the temperature<br />
variation on the case surface which is a good basis to verify<br />
whether the simple R thJC approach is valid for the given<br />
package or not.<br />
As the measurement and simulation techniques mutually<br />
support each-other, the ultimate solution for package<br />
thermal characterization may be the simulation model<br />
creation based on real measurements [3].<br />
II. Experimental<br />
In order to compare simulation results with actual<br />
measurements a power diode package was selected having a<br />
cylindrical package shape and a hexagonal silicon diode<br />
inside.<br />
Fig. 1: Package structure of the studied diode (actual on left, 3D model on<br />
right)<br />
The package is app. 8.4 mm high and has a 12.5 mm<br />
diameter cooling surface on its bottom. The hexagonal die is<br />
located in the top region of the package and it is mounted<br />
using a top and bottom die attach layer to the copper<br />
cylinders. The edges of the hexagon are located on a 8.1<br />
324
mm diameter circle. The package and die geometry as well<br />
as the material properties are easy to identify, however the<br />
thermal properties of the two die attach layers are unknown.<br />
The properties of these layers can be individually set in the<br />
detailed model for furter refinement.<br />
The sample was pressed against a water-cooled cold-plate<br />
while a thermally conductive grease was used to establish<br />
proper thermal contact between the diode package and the<br />
surface of the cold-plate.<br />
The numerical simulation was conducted by FloTHERM<br />
from Mentor Graphics. Convective and radiative heat<br />
transfer were assumed insiginifcant and thus not modelled.<br />
The circulated water was modelled as a region of constant<br />
temperature. The final setup can be seen in Fig. 2.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
The measurement starts at the switching and finishes in the<br />
cold thermal steady-state of the diode as prescribed by the<br />
JEDEC JESD 51-1 test standard. Based on the voltage<br />
change of the diode it is easy to calculate the temperature<br />
change knowing the so-called k-factor, the relationship<br />
between the diode’s forward voltage and the temperature at<br />
a constant, steady current. The sensitivity measured on this<br />
diode was equal approximately to the textbook value,<br />
2mV/°C. Knowing the k-factor and having the temperature<br />
vs. voltage characteristics measured by the transient tester it<br />
is easy to plot the change of the junction temperature of the<br />
device during the test. The power step used in this study<br />
was app. 4.07 W. For comparision purposes the same power<br />
was set in the thermal simulator to the active volume of the<br />
silicon die. For ease of simualtion, the numerical model<br />
considered a step increase in power dissipation. Both the<br />
measured and simulated temperature difference curves can<br />
be viewed in Fig. 4.<br />
Fig. 2: Detailed model resembling the real measurement conditions<br />
As a first step thermal transient measurements were<br />
carried out on the packaged semiconductor device.<br />
Thermal transient measrements require a power step on<br />
the juntion of the semiconductor device, which is usually<br />
supplied by electrical means. In this measurement the<br />
following electrical setup was used:<br />
Fig. 3: Electrical test setup used in this study<br />
Before the actual measurement starts, the sum of a<br />
heating current and a measurement current is forced through<br />
the semiconductor diode. As the temperature of the diode<br />
stabilizes the heating current is suddenly switched off.<br />
Typical fall time is less than a 1µs. As soon as the switching<br />
takes plase from the high current to the measurement<br />
current value, the voltage of the device is monitored with a<br />
time resolution of 1 µs and a voltage resolution of 12 µV.<br />
Fig. 4: Measured and simulated transient responses of the studied structure.<br />
It is easy to observe that the shape of the curves is similar,<br />
at one part of the transient they even fit, however there is a<br />
slight difference between the rise time and the total<br />
elevation of the two transient responses. Important unknown<br />
parameters in the detailed model are the thermal<br />
conductivity and bondline thickness of the die attach layers<br />
and thermal conductivity of the grease at the package<br />
boundary.<br />
In order to identify the source of the differences between<br />
the curves in Fig. 4, the transient responses were turned into<br />
system descriptive structure functions [4,5].<br />
In the structure functions the cumulative thermal<br />
capacitance is plotted as a function of the cumulative<br />
thermal resistance. If the heat-flow path is mainly onedimensional<br />
as in case of most power semiconductor<br />
packages having an exposed cooling tab, these functions<br />
provide a map of the heat-flow path from the heat-source<br />
which is the semiconductor junction to the ambient. This<br />
approach allows the identification of partial thermal<br />
resistances and partial thermal capacitances in the main heat<br />
conduction path.<br />
The structure functions corresponding to the measured<br />
sample and the detailed model can be seen in Fig. 5.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
which that heat flows is also high. The colored sections of<br />
the following figures show areas of increased thermal<br />
bottleneck, with red being the largest. The arrows<br />
correspond to the direction of the heat flow<br />
The short steep section between point 1 and point 2<br />
corresponds to the predominant resistance of the top die<br />
attach layer as felt by the heat wave, see Fig. 7.<br />
Fig. 5: Structure function view of Fig. 4<br />
In practice the curve is a very useful tool for comparison<br />
studies, e.g. for the identification of die attach voids in a<br />
power package by comparing the structure function<br />
corresponding to a good, so-called “golden reference<br />
device” to an unknown device. [6] When comparing a<br />
simulation to a measurement these graphical structures may<br />
help to identify the source of the difference between two<br />
results in terms of which package structures are responsible<br />
for the observed differences. This is impossible to be done<br />
by analyzing the time-domain curves only. In case of<br />
comparing a simulation to a real measurement this approach<br />
may help to fine-tune the simulation model by revealing<br />
material property data of the internal layer structure.<br />
In order to match the layers identified by the structure<br />
functions to the internal layers of the real assembly, all steps<br />
of the transient simulation were saved individually and the<br />
heat-spreading was visualized. Based on the results the<br />
structure function of the measured sample was divided into<br />
four different parts, see Fig. 6.<br />
Fig. 7 : Heat spreading from the die in the simulated package at 1.69E-4<br />
seconds. The arrows show the heat-flux originating from the active layer<br />
Both the lower and upper die attach layers offer a thermal<br />
resistance, however as the time goes towards steady-state<br />
the large amount of copper beneath the die (compared to the<br />
small easily filled amount of copper above it) and the<br />
increased cooling capacity offered by that path is felt,<br />
pulling the heat downwards resulting in a larger thermal<br />
resistance value as the majority of the heat passes through<br />
this area, see the flat section indicated by point 3.<br />
Fig. 8: Heat flux at 0.76 seconds, dominant thermal bottleneck now in the<br />
bottom die attach<br />
This phenomenon is nicely illustrated in Fig. 8. The high<br />
bottleneck red area has shifted to the bottom die attach<br />
which now acts as the significant thermal resistance in the<br />
thermal path.<br />
As time passes the heat fills up the total copper volume of<br />
the package and enters the TIM (thermal interface material)<br />
layer at the package boundary. This state is shown in Fig. 9.<br />
Fig. 6: Structure function corresponding to the measured transient, divided<br />
into four structural elements<br />
As a first step the die itself cools down and appears as a<br />
very steep section in the function indicating that the thermal<br />
capacitance of the layer is very big compared to its thermal<br />
resistance.<br />
The concept of Thermal Bottlenecks and their<br />
visualization in 3D thermal simulations is covered in [7,8].<br />
Areas of increased thermal bottlenecks indicate where both<br />
the heat flux is high and the thermal resistance through<br />
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<br />
Although the two curves fit very well even in the<br />
structure function space where the differences are enhanced<br />
compared to the time domain, minor differences can be<br />
observed. We believe that they originate from the fact that<br />
the simulations are entirely noiseless while the<br />
measurements inherently contain some white noise. In case<br />
of the electric measurement it is practically around 24-36<br />
µV.<br />
A. Compact model generation<br />
Fig. 9: Steady-state heat-flux distribution in the simulated package and its<br />
environment<br />
In thermal steady-state the thermal grease acts as the<br />
biggest thermal bottleneck in the studied system. The long<br />
straight line in Fig. 6 marked by number 4 corresponds to<br />
this layer and shows that it has a very large thermal<br />
resistance compared to its low thermal capacitance.<br />
This example shows the actual conditions in this<br />
particular system, however similar heat-spreading<br />
mechanisms can be identified in other power packages as<br />
well. The results clearly show that there are three main parts<br />
of the structure where the fine-tuning of the geometry and<br />
the material parameters may be necessary, they are the two<br />
die attach layers with special emphasize on the bottom one<br />
and the TIM layer.<br />
Still the number of the variables is large; however after<br />
14 rounds of iterative fine-tuning we achieved a very good<br />
correspondence between the results of the simulation model<br />
and the real device. The corresponding structure functions<br />
can be viewed in Fig. 10. The slight difference between the<br />
shapes of the structure functions at their end, which describe<br />
the ambient, is a result of the simplified model of the water<br />
in the cold-plate. As the goal of the study was the<br />
appropriate modeling of the package, not the ambient, the<br />
flow of the water was not taken into account.<br />
A compact model is a simplified representation of the<br />
thermal behaviour of a semiconductor package. The goal of<br />
the compact model (CTM) creation is not to resemble the<br />
real package geometry, but to allow the prediction of<br />
temperatures at important points of a thermal system such<br />
as the junction itself. If a CTM is dynamic it is also possible<br />
to predict the change of the temperature of the given node<br />
as a function of time. As the CTM is an abstraction of the<br />
component only, it requires less griding thus less<br />
computational efforts. This is in addition to the fact that the<br />
CTM hides proprietory information about the package<br />
construction.<br />
Such a model can be generated based on the thermal<br />
resistance – thermal capacitance values derived from the<br />
strucutre functions if a one-dimensional heat spreading can<br />
be assumed. The resulting RC ladder has to be cut at the<br />
case node. In order to identify it, the dual-interface<br />
methodology was used. [9]<br />
A point of separation of the structure functions was<br />
identified after making two thermal measurements of the<br />
same package with different boundary conditions. Whilst<br />
the heat spreads in the same structure the structure functions<br />
are identical, but as soon as the main trajectory of the heat<br />
leaves the package boundary and enters the two different<br />
interface layers a point of separation can be identified. In<br />
case of this particular package we measured 0.34K/W. All<br />
the derived thermal resistance and thermal capacitance<br />
points are can be used for the model creation up to this<br />
resistance value.<br />
Fig. 11: Identification of the case node using the dual-interface method<br />
Fig. 10: Comparison of the detailed model to the measurement results<br />
In the derived model the real geometry is not considered,<br />
only the total volume and the contact area of the package is<br />
327
given. This allows a CTM model to be provided by package<br />
manufacturers even along with the standard datasheets as<br />
they include no proprietary information of the internal<br />
details of the package at all.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
If one tries to apply the double-interface methodology [9]<br />
to obtain the R thJC of such a package, difficulties may arise.<br />
As the area of one individual chip compared to the area of<br />
the entire package is small, the internal heat-spreader of the<br />
packages may already serve as a cold-plate in the system<br />
[10]. If the boundary condition is thermally good (e.g.<br />
grease is applied between the package surface and the coldplate,<br />
the heat leaves the heat-spreader into the cold-plate<br />
through a relatively narrow spreading cone. On the other<br />
hand if no thermal grease is applied between the two<br />
surfaces, the heat will spread first laterally along the heatspreader<br />
before entering the cold-plate.<br />
This simple behavior will be visible in the structure<br />
functions<br />
Fig. 12: Derived CTM – No real geometry data is revealed<br />
A transient simulation is carried out on this model and,<br />
for CTM accuracy verification, the resulting structure<br />
function was compared to the original experimental model.<br />
The results were satisfactory, as shown in Fig. 13.<br />
T3Ster Master: cumulative structure function(s)<br />
0.06339 0.101006<br />
10000<br />
1000<br />
FMG2G150US60_10A_T25 - Ch. 0<br />
FMG2G150US60_10A_T25_MY - Ch. 0<br />
Cth [Ws/K]<br />
100<br />
10<br />
1<br />
49.6756<br />
Fig. 13: Comparison of the behavior of the CTM and the real structure<br />
A perfect match is obtained between the two structure<br />
functions up to the package boundary, which proves that the<br />
measurement-based CTM worked as expected. The app. 5%<br />
difference in overall thermal resistance is assumed to be a<br />
result of the fixed temperature approach for modeling the<br />
water or the error on the TIM resistance measurement itself.<br />
B. Problems with packages having larger surface area<br />
For packages having multiple heat-sources, such as half-,<br />
or full bridge modules hosting multiple MOSFET or IGBT<br />
devices the 1D heat-flow assumption is in most of the cases<br />
not valid. It is a common practice that engineers try to<br />
characterize these systems with a single R thJC value,<br />
however modeling such a complex package with one single<br />
value may lead to inaccuracies.<br />
0.1<br />
0.01<br />
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4<br />
Rth [K/W]<br />
Fig. 14. Junction-to-case thermal resistance results in case of a 7PM-GA<br />
package hosting 2 IGBT devices<br />
Fig. 14. shows the calculated structure functions of a<br />
power package which has both of the above mentioned<br />
problems. The blue curve shows the case when the device<br />
was put directly on a cold-plate using thermal grease. The<br />
red curve represents the case when no grease was applied. It<br />
is clearly visible that the location of the diverging point is at<br />
0.063K/W. After this point however the curves run parallel,<br />
up to app. 0.16K/W. The parallel running structure<br />
functions describe the heat spreading in the package base;<br />
this is partly proven by the long horizontal part in the red<br />
functions which is the thermal resistance of the air gap<br />
between the contacting surfaces.<br />
328
The fact that the location of the diverging point is so<br />
much indefinite also points out that such complex packages<br />
should not be modeled with a single thermal resistance<br />
value. Based on the fact that the measurement based<br />
compact modeling approach nicely describes discrete<br />
components including the internal details of the package<br />
there is a possibility to make measurements on each<br />
individual chip in a multichip module and cut the RC<br />
ladders at the point where the heat starts to spread along the<br />
internal heat-spreader. The resulting CTM-s which<br />
describes the individual chips can be connected to a detailed<br />
model of the heat-spreader. The resulting ‘hybrid model’<br />
could accurately predict the thermal behavior of each single<br />
chip while taking into consideration the temperature<br />
coupling effects between the chips by modeling the<br />
conduction within the heat-spreader. This is however a topic<br />
of our future research.<br />
III. Conclusions<br />
In this paper the steps of a complex thermal<br />
characterization by a combined measurement and<br />
simulation approach are explained.<br />
In case of the creation of a detailed package simulation<br />
model there may be some unknown parameters such as the<br />
exact bond-line thickness and thermal conductivity of the<br />
thermal interface materials. Taking the structure functions<br />
calculated based on the measured thermal transient response<br />
of the package one can get a good reference to which the<br />
simulation model can be calibrated in an iterative way. With<br />
the help of this approach very good calibration results were<br />
achieved at the package level.<br />
Another option is the generation of a CTM of a power<br />
package based on real thermal measurements. This<br />
approach has the prerequisite of a one-dimensional heatflow<br />
path from the junction towards the package surface.<br />
Such a heat-spreading phenomenon is characteristic to<br />
power packages with one dedicated cooling tab.<br />
This study proves that the structure function approach is a<br />
good tool for compact model thermal model generation as<br />
the simulated results match with the measured results<br />
perfectly. In order to achieve such good results only two<br />
quick thermal measurements have to be carried out saving<br />
both time at the model creation and also at the solution. In<br />
our study we focused on the modeling at package level, the<br />
model of the environment was not fully refined which<br />
results in minor deviations at the parts of the structure<br />
functions describing the ambient. Verifying the effect of<br />
the water-flow in the pipes on the structure functions has<br />
scope for further study.<br />
The perfect match between the simulation results using<br />
the CTM and the measurements at the package level raises<br />
new possibilities for the modeling of more complex multichip<br />
modules, too. If one can work out a suitable algorithm<br />
to cut the structure function at the point where the heat<br />
enters the base-plate of a given package, the chip and its<br />
close thermal environment can be modeled using an<br />
accurate CTM, while the cross-coupling effects could be<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
simulated using a detailed model of the base-plate itself.<br />
This ‘hybrid model’ could be distributed along with the<br />
package itself as it inherently hides proprietary geometry<br />
and material data, however allows the end-users to<br />
accurately model the in-situ thermal behavior.<br />
REFERENCES<br />
1. T. Bohm , T. Hauck, “Scaleable thermal RC-network<br />
approach to analyze multichip power packages”. Proc. 6th<br />
THERMINIC, Budapest, pp. 230-234, 2000.<br />
2. B.S. Lall, B.M. Guenin, R.J. Molnar, “Methodology for<br />
thermal evaluation of multichip modules”, Proc. 21 st<br />
SEMITHERM, San Jose, Page(s): 72-79, 1995<br />
Digital Object Identifier: 10.1109/STHERM.1995.512054.<br />
3. András Poppe, Andras Vass-Varnai, Gábor Farkas, Marta<br />
Rencz, “Package characterization: simulations or<br />
measurements? In: Proceedings of the 10th Electronics<br />
Packaging Technology Conference (EPTC'08). Singapore,<br />
2008.12.09-2008.12.12.pp. 155-160. Paper A6.4. (ISBN:<br />
978-1-4244-2117-6)<br />
4. O. Steffens, P. Szabo, M. Lenz, and G. Farkas, "Thermal<br />
transient characterization methodology for single-chip and<br />
stacked structures", Proc. 21th SEMITHERM, San Jose, pp.<br />
313-321, 2005.<br />
5. V.Szekely, "Identification of RC networks by<br />
deconvolution: Chances and Limits", IEEE Trans. On<br />
Circuits and Systems – I: Fundamental Theory and<br />
Applications, Vol. 45, No. 3, pp. 244-258, 1998.<br />
6. M. Rencz, V. Székely, A. Morelli, C. Villa, “Determining<br />
partial thermal resistances with transient measurements and<br />
using the method to detect die attach discontinuities”, 18th<br />
Annual IEEE SEMI-THERM Symposium, March 1-14<br />
2002, San Jose, CA,USA, pp. 15-20<br />
7. J. Parry, R. Bornoff, B. Blackmore, “Thermal Bottlenecks<br />
and shortcut opportunities; innovations in electronics<br />
thermal design simulation”, Electronics cooling, September<br />
2010<br />
8. R. Bornoff, B. Blackmore, J. Parry, “Heat Sink Design<br />
Optimization Using the Thermal Bottleneck Concept”, 27th<br />
Annual IEEE SEMI-THERM Symposium, March 2011,<br />
San Jose, CA,USA<br />
9. D. Schweitzer, “The junction-to-case thermal resistance: A<br />
boundary condition dependent thermal metric”, 27th<br />
Annual IEEE SEMI-THERM Symposium, March 2011,<br />
San Jose, CA,USA, pp. 151<br />
10. A. Vass-Varnai, S. Gao, Z. Sarkany, J. Kim, S. Choi, G.<br />
Farkas, A. Poppe, M. Rencz “Issues in junction-to-case<br />
thermal characterization of power packages with large<br />
surface area”, 26th Annual IEEE SEMI-THERM<br />
Symposium, March 2010, San Jose, CA,USA, pp. 151<br />
ACKNOWLEDGMENT<br />
The work was partially supported by the SE2A ENIAC JU project (No<br />
120009) of NKTH (No OMFB-00521/2009) and by the JEMSiP_3DENIAC<br />
JU project (No 120016) of NKTH (No OMFB-00166/2010).<br />
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<br />
LINEAR ENERGY CONTROL OF LASER<br />
DRILLING AND ITS APPLICATION FOR<br />
TFT-LCD BRIGHT PIXEL REPAIRING<br />
Taco Chen and Ming-Tzer Lin *<br />
Graduate Institute of Precision Engineering,<br />
National Chung Hsing University, Taichung 402, Taiwan, R.O.C.<br />
Abstract- Laser drilling is energy dependent and linear<br />
proportion to the thickness of the materials. Therefore, a linear laser<br />
power supplier will be convenient applying in extensive aperture<br />
processes. However, practically, the energy output of laser<br />
equipments is non-linear. To obtain a linear energy output, laser<br />
power meter is utilized for energy compensation, but the application<br />
of laser power meter requires ceasing the operation of laser<br />
equipment. In this paper, a linear energy compensation method was<br />
investigated and designed by using a measurement of laser energy<br />
output that provides a stable linear energy laser for processes. In the<br />
method, the laser energy testing only requires a fixed time for<br />
measuring laser energy and changing laser energy compensate table.<br />
Furthermore, the laser equipment doses not need stop during the<br />
laser power meter calibration. In addition, a software method for<br />
linear energy compensation was designed and applied to the laser<br />
equipments which have no laser power meter compensation<br />
practices. The method could control and compensate laser energy in<br />
linear output which the energy linear proportion (R square) reaches<br />
0.9989 and provided a very stable power source. When using this<br />
laser method in the LCD panel design processes, the successes rate<br />
reached 80% in performing the bright pixel repair. In panel defect<br />
repair, it could prevent taking the case apart from module and<br />
fabricate that increases the efficiency in production.<br />
Keyword: non-linear energy, linear energy compensation, bright<br />
pixel repair<br />
I. INTRODUCTION<br />
Laser beam repairing is one of the most widely used<br />
thermal energy based non-contact type advance machining<br />
process which can be applied for almost whole range of<br />
materials. In microelectronics processing and manufacturing, it<br />
has became one of an important processes for increasing the<br />
final yielding and product refinement. Laser beam is focused<br />
for melting and vaporizing the unwanted material from the<br />
parent material. Among various type of lasers used for<br />
machining in industries, CO2 and Nd:YAG lasers are most<br />
established. It has become a common repairing tool for fix the<br />
common defect in TFT-LCD.<br />
In the recent years, TFT-LCD has become one of the key<br />
electronics appliances in our daily life. In addition, because of<br />
its light in weight and thin in panel display TFT-LCD has<br />
become an alternative display to replace traditional CRT TV.<br />
However, TFT-LCD is an advanced industry which costs a lot<br />
in its fabrication and requires an extremely high yield. To fulfill<br />
the increasing green engineering and low price demand of the<br />
TFT-LCD optoelectronic products, the high yield and refurbish<br />
of each useful component have become one of the most<br />
important issues. As a result, the laser beam repairing technique<br />
that is used to repair the driving IC and the liquid crystal display<br />
(LCD) panel becomes one of the key techniques in the<br />
processes of the flat panel display manufacturing.<br />
Previously, researchers have explored a number of ways<br />
to improve the laser repairing process performance by<br />
analyzing the different factors that affect the quality<br />
characteristics. The experimental and theoretical studies show<br />
that process performance can be improved considerably by<br />
proper selection of laser parameters, material parameters and<br />
operating parameters.<br />
In particular, proper selection of laser energy parameters<br />
is the most important. Usually, laser drilling is energy<br />
dependent and laser energy is proportion to the thickness of the<br />
materials. Thus, design and development of a functional linear<br />
laser power supplier is needed to extensive aperture processes.<br />
However, the energy output of laser equipments is<br />
non-linear and is traditionally rely on the on-off process for the<br />
energy control. In order to obtain a smooth linear energy output<br />
of laser drill for TFT-LCD defect repairing. This research work<br />
carried out in the area of energy control study for different<br />
TFT-LCD materials and defect shapes. It reports about the<br />
experimental and theoretical studies of Laser repairing to<br />
improve the process performance. Several modeling and<br />
optimization techniques for the determination of optimum laser<br />
beam cutting condition have been critically examined.<br />
In addition, we will discuss how to repair the bad pixels<br />
of TFT-LCD, including laser cutting and welding. We use the<br />
D.O.E (Design of Experiment) method to obtain the factors of<br />
process parameters in laser repairing, which let us to know the<br />
important factor in the laser energy process parameters.<br />
II. EXPERIMNETAL<br />
In general, there are inevitable defects resulted from the<br />
TFT-LCD fabrication processes. However, some defects such<br />
as the bad pixel can be repaired as dark or lightly bright point.<br />
This paper studies several laser repairing methods to repair the<br />
defects of the pixel which permits those repaired pixels<br />
working normally. Based on the TFT-LCD fabrication process,<br />
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<br />
driving principle, laser repairing and pixel design, we adopt the<br />
adjacent pixel to drive the bad one. The repaired pixel will be<br />
normally operated.<br />
Figure 1~3 show common defect happened in metal and<br />
insulation lines in TFT-LCD after processing.<br />
Figure 4: Schematic view of the laser repairing for a TFT-LCD<br />
pixel.<br />
Figure 1: Common defect in TFT-LCD (Metal)<br />
Figure 5: Internal rescue circuit lines design in TFT-LCD<br />
Figure 2: Common defects in TFT-LCD (Insulation)<br />
In general, proper selection of laser energy parameters is<br />
the most important for TFT-LCD repairing. Usually, laser<br />
drilling is energy dependent and laser energy is proportion to<br />
the thickness of the materials. Thus, design and development of<br />
a functional linear laser power supplier is needed to extensive<br />
aperture processes. In recent years, researchers have explored a<br />
number of ways to improve the laser repairing process<br />
performance by analyzing the different factors that affect the<br />
quality characteristics. Figure 6 shows Example of Nd:YAG<br />
Laser repairing processes [13]. The experimental and<br />
theoretical studies show that process performance can be<br />
improved considerably by proper selection of laser parameters,<br />
material parameters and operating parameters.<br />
Figure 3: Common defects in TFT-LCD (Metal)<br />
In the most common laser repairing for TFT-LCD, one of<br />
best ways for repairing is to use the flowing metal between data<br />
line and ITO. When the pixel has been damaged, it will be<br />
repaired by cutting the TFT of pixel and connecting the data<br />
line with ITO by laser welding; therefore, the damaged pixel<br />
will be lightly bright point as shown in Figure 4. The other way<br />
is to use the diode to replace the flowing metal, the diode is<br />
used as a filter for ac data signals and the damaged pixel will<br />
work normally in checking image on matter what is black or<br />
white screens. Figure 5 shows a general design layout for metal<br />
rescue lines adjacent to the TFT-LCD circuit of a pixel.<br />
Figure 6: Nd:YAG Laser repairing processes [13]<br />
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<br />
Practically, the energy output of laser equipments is<br />
non-linear. In order to obtain a linear energy output, the laser<br />
power meter is utilized for the calibration of energy<br />
compensation. We measured the laser energy output as shown<br />
in Figure 7, and then utilized D.O.E (Design of Experiment)<br />
method to obtain the optimized laser energy parameters for the<br />
laser energy.<br />
Figure 8: The experimental results of linearization<br />
laser Energy of HOYA Laser repairing processes<br />
Figure 7: The laser energy output measurement system<br />
This D.O.E method was used to calibrate the output<br />
energy through 200 energy steps. We used equations 1 to obtain<br />
modified energy output of each step thus to obtain linear energy<br />
output. Where N can be defined as 10 for each 10 energy steps<br />
or 20 for each 20 energy steps.<br />
(Step Power Max — Step Power Min )/ N=Rang Power /step (1)<br />
Thus we can obtain a reference table to linearlize the<br />
laser energy output. In order to refine the output energy for each<br />
step. The additional calibration can be performed using<br />
equation (1) again and set the N number to be 5.<br />
Using this study allowed us to create the linear<br />
compensation table that provides a stable linear energy laser for<br />
processes. The method, used in this experiment, required only a<br />
fixed time for measuring laser energy and allowed laser energy<br />
changed according to the compensate table. As a result, the<br />
laser equipment doses not need to run on-off control and can be<br />
performed no-stop process during the laser drilling.<br />
III. RESULTS AND DISCUSSION<br />
Figure 8 shows the experimental results of linearization<br />
laser Energy of HOYA Laser repairing processes using above<br />
mention method.<br />
Application used linear laser drilling energy<br />
compensation method was performed and tested to the laser<br />
equipments which have no laser power meter compensation<br />
practices. The results indicated that this method controlled and<br />
compensated laser drilling energy output linearly in which the<br />
energy linear proportion reached R square in 0.9989, providing<br />
a very stable power source. Moreover, when using this method<br />
in the TFT- LCD panel repairing processes, the successes<br />
repairing rate reached 80% in performing the bright pixel repair.<br />
In addition, in panel defect repair, it prevents taking the case<br />
apart from module and fabricate that increases the efficiency in<br />
TFT-LCD production drastically. Figure 9 shows the SEM<br />
image of repaired TFT-LCD metal section using proposed<br />
experimental results of laser repairing processes. Figure 10<br />
shows the images of with and without repaired TFT-LCD panel<br />
using proposed method for laser repairing processes.<br />
Figure 9: The SEM image of repaired TFT-LCD<br />
metal section using proposed experimental results of laser<br />
repairing processes<br />
332
Figure 10: The images of without and with repaired<br />
TFT-LCD using proposed experimental results of laser<br />
repairing processes<br />
IV. CONCLUSION<br />
A linear energy compensation method was investigated<br />
and designed by using a measurement of laser energy output<br />
that provides a stable linear energy laser for processes. In the<br />
method, the laser energy testing only requires a fixed time for<br />
measuring laser energy and changing laser energy compensate<br />
table. Furthermore, the laser equipment doses not need stop<br />
during the laser power meter calibration. In addition, a software<br />
method for linear energy compensation was designed and<br />
applied to the laser equipments which have no laser power<br />
meter compensation practices. The method could control and<br />
compensate laser energy in linear output which the energy<br />
linear proportion (R square) reaches 0.9989 and provided a very<br />
stable power source. When using this laser method in the LCD<br />
panel design processes, the successes rate reached 80% in<br />
performing the bright pixel repair. In panel defect repair, it<br />
could prevent taking the case apart from module and fabricate<br />
that increases the efficiency in production.<br />
ACKNOWLEDGMENT<br />
The authors are grateful to the assistance from both<br />
engineers in Shuz Tung Machinery Industrial CO. LTD,<br />
Taiwan and AU Optronics Corp. Taiwan.<br />
REFERENCES<br />
[1] C. Jeong,Y.Jinwoo,P. G. Poo ―A defect inspection method for TFT<br />
panel using the compute unified device architecture (CUDA)‖,<br />
Industrial Electronics, 2009. ISIE 2009. IEEE International<br />
Symposium on, (2009) p779-782.<br />
[2] W. C. Lee ,J. B. Song , B. Y. Kim, S. H. Park, S. M. Lim, W. J.<br />
Lee―Auto Defect Repair Algorithm For LCD Panel Review &<br />
Repair Machine‖, SICE Annual Conference, 2008 , p2200-2203.<br />
[3] J. Lee, J. Ehrmann, D. Smart, J. Griffiths, J. Bernstein, ―Analyzing<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
the Process Window for Laser Copper-link Processing,‖ Solid State<br />
Technology,(2002) p63-66.<br />
[4] J. B. Bernstein, J. Lee, G. Yang, T. Dahmas, ―Analysis of Laser<br />
Metal-cut Energy Process Window,‖ IEEE Semicondut. Manufact.,<br />
Vol. 13, No. 2,(2000) p228-234.<br />
[5] K.J. Chung,‖ Microbridge Formation for Low Resistance Interline<br />
Connection Using Pulsed Laser Techniques‖ Doctor of Philosophy,<br />
2005.<br />
[6] G. Chryssolouris, Laser Machining—Theory and Practice.<br />
Mechanical Engineering Series, Springer-Verlag, New York Inc.,<br />
NewYork,(1991).<br />
[7] J.D. Majumdar, I. Manna, Laser processing of materials,<br />
Sadhana 28 (3–4) (2003) p495–562.<br />
[8] T. Norikazu, Y. Shigenori, H. Masao, Present and future of lasers for<br />
fine cutting of metal plate, Journal of Materials Processing<br />
Technology 62 (1996) p309–314<br />
[9] C. H. Li, M. J. Tsai, R. Chen, C. H. Lee, S. W. Hong, Cutting for QFN<br />
packaging by diode pumping solid state laser system, Proceedings of<br />
IEEE Workshop on Semiconductor Manufacturing Technology<br />
(2004) p123–126.<br />
[10] C. H. Li, M. J. Tsai, S. M. Yao, Cutting quality study for QFN<br />
packages by Nd:YAG laser, Proceedings of the IEEE International<br />
conference on Mechatronics (ICM’05) (2005) p19–24.<br />
[11] C.-H. Li, M.-J. Tsai, C.-D. Yang, Study of optimal laser parameters<br />
for cutting QFN packages by Taguchi’s matrix method, Optics and<br />
Laser Technology 39, (2007) p786–795.<br />
[12] J.K.S. Sundar, S.V. Joshi, Laser cutting of materials, Centre for Laser<br />
Processing of Materials, International Advance Research Centre for<br />
Powder Metallurgy and New Materials, Hyderabad.<br />
[13] A. K. Dubey a and V. Yadava, Laser beam machining—A review,<br />
International Journal of Machine Tools and Manufacture, Vol 48,<br />
Issue 6, (2008) p609-628<br />
[14] D. K. Y. Low, L. Li, P. J. Byrd, The influence of temporal pulse train<br />
modulation during laser percussion drilling, Optics and Lasers in<br />
Engineering 35:149-164,2001<br />
333
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Measurement of Electrical Properties of Materials<br />
under the Oxide Layer by Microwave-AFM Probe<br />
Lan Zhang, Yang Ju*, Atsushi Hosoi, and Akifumi Fujimoto<br />
Department of Mechanical Science and Engineering, Nagoya University<br />
Furo-cho, Chikusa-ku, Nagoya 4648603, Japan<br />
Abstract- The capability of a new AFM-based apparatus<br />
named microwave atomic force microscope (M-AFM) which<br />
can measure the topography and electrical information of<br />
samples simultaneously was investigated. Some special samples<br />
with different thicknesses of dielectric film (SiO 2 ) which plays<br />
the role of oxide layer creating on the material surface were<br />
fabricated. The measurement of electrical properties of<br />
materials under the oxide layer by the M-AFM was studied.<br />
The results indicate that the M-AFM can lead the microwave<br />
signal penetrate the oxide film (SiO 2 ) with a limited thickness of<br />
60 nm and obtain the electrical information of underlying<br />
materials.<br />
I. INTRODUCTION<br />
Scanning probe microscope (SPM) has become increasingly<br />
important, as an evaluation apparatus having a spatial<br />
resolution on nanometer scale. After the scanning tunneling<br />
microscope (STM), the atomic force microscope (AFM) and<br />
the near-field scanning optical microscope (NSOM) were<br />
invented, and various kinds of SPMs based on these<br />
microscopes have been developed. These improved SPMs<br />
made it possible to measure not only the topography of<br />
materials but also the thickness of the oxidized membrane,<br />
the profile of the two-dimensional dopant [1], the<br />
distribution of the electrical potential and the magnetic field<br />
on the material surface [2], the distribution of the hardness<br />
and stiffness on the material surface [3] and so on.<br />
Electrical properties are very important fundamental<br />
properties, since they have an enormous influence on the<br />
functionality of materials. However, development of a<br />
technology which is able to measure electrical properties<br />
such as conductivity, permittivity and permeability on a<br />
nanometer scale is far behind the development of the SPMs<br />
noted above. In particular, the electrical properties of<br />
materials in a nano-region are affected not only by the<br />
structure and composition of these materials, but also by the<br />
mechanical factors of stress and strain due to lattice<br />
vibrations. The measurement of electrical properties in a<br />
nano-region is expected to be used in various fields for the<br />
fabrication of nanomaterials, the development and evaluation<br />
of nanodevices, the elucidation of various mechanisms<br />
within living tissues and so on. On the other hand,<br />
microwave microscopies have been developed for the<br />
measurement of electrical properties and the detection of<br />
defects in the microscopic regime [4-6]. Duewer et al. [7]<br />
developed a scanning evanescent microwave microscope<br />
(SEMM) with which they succeeded in measuring the<br />
resistivities of three kinds of metallic materials by using the<br />
property of microwaves that the resonant frequency changes<br />
depending upon the capacitance between the probe tip and<br />
the material’s surface. Tabib-Azar and Akiwande [8] were<br />
successful in detecting and imaging depletion regions in<br />
solar cell p-n junctions in real time. Ju et al. [9] were<br />
successful in the detection of delamination in integrated<br />
circuit packages by exploiting the properties of microwave<br />
signals that change depending on the electrical properties of<br />
the materials.<br />
To evaluate the electrical properties of materials using<br />
microwaves, it is necessary to keep the standoff distance<br />
between the microwave probe and the sample constant<br />
because microwave signals in the near-field are extremely<br />
sensitive to this distance. Otherwise, it is difficult to<br />
distinguish the change in the signal due to the difference of<br />
the material properties or due to the change of the stand-off<br />
distance. In particular, to evaluate the electrical properties of<br />
materials with high resolution on a nanometer scale, it is<br />
indispensable to control the stand-off distance precisely on<br />
the order of a nanometer. In order to solve these problems, Ju<br />
et al. [10-12] proposed a microwave atomic force microscope<br />
(M-AFM). The M-AFM has the characteristics that it can<br />
maintain a constant stand-off distance as an AFM and it also<br />
can evaluate the electrical properties of materials<br />
quantitatively as a microwave microscope. With this unique<br />
combination, M-AFM is able to evaluate the electrical<br />
properties, as well as the topography, of a material<br />
simultaneously in one scanning process with<br />
nanometer-scale resolution.<br />
Recently, the researches about electrical characteristics<br />
of metallic film and membrane on a nano-scale have become<br />
the hot topics more and more. However, if the surface of the<br />
metallic film or membrane is covered by a thin oxide layer,<br />
the traditional method will fail in evaluating the electrical<br />
property of material under the oxide layer, because it is<br />
difficult to make direct contacts to the material under test. On<br />
the contrary, M-AFM can be used to solve this problem,<br />
because the microwave signals emitted from the tip of<br />
M-AFM can penetrate the dielectric film, and have an<br />
334
interaction with the underlying materials. By analyzing the<br />
reflected microwave signals, the electrical properties of<br />
underlying materials can be evaluated. In this paper, some<br />
special samples with different thickness of dielectric films<br />
which plays the role of oxide layer created on the material<br />
surface were fabricated, and the measurement of electrical<br />
properties of materials under the oxide layer by the M-AFM<br />
was investigated in details.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
A<br />
Microwave generator<br />
Tunable<br />
short<br />
Measurement system<br />
Attenuator<br />
Multiplier<br />
Magic<br />
tee<br />
II.<br />
EXPERIMENTAL PROCEDURE<br />
A. Probe Fabrication<br />
To restrain the attenuation of microwave propagating in<br />
the probe, a non-doped GaAs wafer was used as the substrate<br />
of the probe. Wet etching was used to fabricate the probe<br />
because it is possible to obtain the desired structure by<br />
causing a side etching under the etching mask. The<br />
fabrication method and process of the M-AFM probe have<br />
been studied in details by Ju et al. [10,11].<br />
Topography<br />
B<br />
Tip<br />
Ag film<br />
Laser<br />
Microwave<br />
image<br />
Cantilever<br />
SiO oxide layer<br />
2<br />
Silicon substrate<br />
Detector<br />
Probe holder<br />
AFM<br />
Fig. 2. Schematic diagram of the M-AFM measurement<br />
system.<br />
The images of the cantilever and tip of M-AFM probe<br />
were taken by scanning electron microscopy (SEM), as<br />
shown in Fig. 1. Fig. 1A shows the forepart of cantilever, a<br />
sharp tip with a height of 7 μm was formed at the front of the<br />
cantilever. As shown in Fig. 1B, a nano-slit was introduced<br />
across the cantilever through the center of the tip by focus ion<br />
beam (FIB) fabrication. The width of the nano-slit is<br />
approximately 100 nm.<br />
Fig. 1. SEM images, A: the cantilever of M-AFM probe; B:<br />
a nano-slit across the probe tip introduced by FIB<br />
fabrication.<br />
B. Microwave Measurement System<br />
The measurements in this paper were carried out by a<br />
compact microwave instrument which is composed of an<br />
amplifier, a magic-Tee, an attenuator, a tunable short, and a<br />
diode detector [13], as shown in Fig. 2. Fig. 2A shows the<br />
flow chart of the operating microwave signals for<br />
measurements. The microwave signals working at a<br />
frequency f=94 GHz, which was generated by a microwave<br />
generator. Then the microwave signals were separated into<br />
two branches by the magic-Tee. One branch signal was sent<br />
to the M-AFM probe to sense the samples and then the<br />
reflected signal was received by the probe tip, as shown in<br />
Fig. 2B. Another branch signal was sent to the attenuator and<br />
then to the tunable-short to form a reference signal with a<br />
constant phase difference and a similar amplitude comparing<br />
with the reflected signal from the sample. The reference<br />
signal was determined by setting the output voltage of the<br />
detector to be a definite value when the M-AFM was set in<br />
air without the approaching, and this was carried out by<br />
adjusting the attenuator and the tunable-short. The reflected<br />
signal and the reference signal were finally synthesized by<br />
335
the magic-Tee, and the coherent signals were measured by<br />
the detector. The detector used in the experiment was a<br />
square-law detector, and the output voltage has a linear<br />
relationship with the squared complex modulus of the<br />
reflection coefficient. When the sample was scanned by the<br />
M-AFM probe, the reflected signal which carries some<br />
useful information of the sample’s electrical property was<br />
received and converted to the voltage value.<br />
C. Experimental Conditions and Samples<br />
Fig. 3A indicates the interaction of microwave signals<br />
with the sample under test. The incident microwave signals<br />
propagated in M-AFM probe and then were emitted at the top<br />
of probe tip. Considering the configuration and dimensions<br />
of the probe tip and the nano-slit from which microwave<br />
signals are emitted, the measurement was dominated by the<br />
interaction between the near-field microwave and a shallow<br />
surface layer of the sample. Therefore, if the distance<br />
between M-AFM probe tip and the sample under test exceeds<br />
the interaction range of the near-field microwave, the<br />
microwave can not sense the sample and no useful<br />
information of the sample’s electrical property is presented<br />
in the reflected microwave signals. In this study, we<br />
researched on the relationship between the thickness of oxide<br />
layer on a metallic film and the reflected microwave signals.<br />
It can be demonstrated from our study that the M-AFM is<br />
able to measure the electrical property of material under a<br />
thin oxide layer.<br />
A<br />
M-AFM<br />
probe tip<br />
Incident wave<br />
Reflected wave<br />
B<br />
SiO oxide layer<br />
2<br />
Silicon wafer<br />
Ag film<br />
Silicon wafer<br />
a Ag film<br />
Silicon wafer<br />
Silicon substrate<br />
Ag film<br />
20 nm Silicon dioxide layer<br />
b<br />
Ag film<br />
Silicon wafer<br />
40 nmSilicon dioxide layer<br />
c<br />
Silicon wafer<br />
Ag film<br />
d<br />
100 nm Silicon dioxide layer<br />
f<br />
Silicon wafer Ag film<br />
Fig. 3. Schematic diagrams, A: interaction of near-filed<br />
microwave with the sample under test; B: the fabrication<br />
process of samples used in this study.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Au films<br />
(wave guide)<br />
GaAs<br />
Near-field<br />
microwave<br />
Interface<br />
The samples under test are prepared as follows. A thin<br />
Ag film with thickness of 100 nm on average was deposited<br />
on the Si wafer by electron beam (EB) evaporation. Then,<br />
SiO 2 films with different thickness were evaporated on the<br />
Ag film respectively, as shown in Fig. 3B. These SiO 2 films<br />
play the role of oxide layer created on the surface of the<br />
metallic film. In this way, we obtained five samples with the<br />
SiO 2 films having the thickness from 20 nm to 100 nm with<br />
the increment of 20 nm and another one without the SiO 2<br />
film.<br />
Since the thickness of Ag film in this study was 100 nm,<br />
which is much larger than the skin depth of Ag for<br />
microwave at the frequency of 94 GHz, only the microwave<br />
signal reflected from the top face of the Ag film can affect the<br />
measurement results. The measurements were performed in<br />
air, with a working environment of temperature 25.0 °C,<br />
relative humidity 50%. The M-AFM worked in a non-contact<br />
mode, and the scanning area and scanning speed were 2 µm ×<br />
2 µm and 1000 nm/sec, respectively.<br />
III.<br />
RESULTS AND DISCUSSION<br />
A. Experiment of Measuring the Samples<br />
Fig. 4A depicts the schematic diagram of experiment of<br />
measuring the samples. Since this study is going to use the<br />
M-AFM probe to scan the Ag samples under 6 different<br />
thickness of oxide film (SiO 2 ), the whole experiment was<br />
separated into 6 times. In order to make all the steps can be<br />
kept in same initial measurement conditions, before the<br />
scanning processes for all the samples, we firstly set the<br />
initial voltage to 1.5 V (by the voltage-offset function of<br />
pre-amplifier) at the situation of keeping a constant distance<br />
of 2.6 µm between the probe tip and measured sample. Then,<br />
during the scanning process, the stand-off distance between<br />
the probe tip and scanning surface was fixed in several<br />
nanometers by the atomic force and the voltage<br />
corresponding to the inspected sample was measured and<br />
recorded. Fig. 4B shows the relationship between the<br />
thickness of oxide film and the measured voltage, which was<br />
converted from the reflected microwave signals.<br />
B. Calibration Experiment<br />
In order to verify the creditability of the measurements,<br />
calibration experiment was also carried out. The M-AFM as<br />
well as the AFM is able to adjust the standoff distance to<br />
different values and keep it constant during the scanning<br />
process. By recording the scanning route of scanning<br />
topography, the cantilever of probe could be lifted up a set<br />
value with nano-meter order. Lifting the cantilever on the<br />
each scanning contour, the M-AFM probe performs<br />
up-and-down motion line by line. Then, the stable<br />
topography and microwave image can be acquired in twice<br />
scanning process. In this study, we input the height value to<br />
lift up the M-AFM tip from the normal feedback position of<br />
topography image with a positive value from 20 nm to 1000<br />
nm. All the other experimental conditions are kept the same<br />
as mentioned for the previous works.<br />
336
A<br />
60 nm SiO 2 100 nm<br />
film<br />
SiO 2 film<br />
Ag film<br />
Ag film<br />
Ag film<br />
Silicon substrate Silicon substrate<br />
Silicon substrate<br />
B<br />
Measured voltage (V)<br />
A<br />
Tip of M-AFM<br />
C<br />
C<br />
20 nm SiO<br />
Microwave oxide layer<br />
image<br />
Microwave<br />
image<br />
Oxide film thickness (nm)<br />
Fig. 4. A: schematic diagram of the scanning process of<br />
samples covered by oxide films with different thickness; B:<br />
the relationship between the measured voltage and the<br />
thickness of oxide film. Inset C and D are the microwave<br />
images of Ag film and Ag film covered by a 60 nm oxide<br />
layer measured by the M-AFM.<br />
Ag film<br />
Silicon substrate<br />
B<br />
Measured voltage (V)<br />
Tip of M-AFM<br />
1.342 V 1.349 V<br />
20 nm SiO<br />
oxide layer<br />
D<br />
D<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Fig. 5A depicts the schematic diagram of calibration<br />
experiment. The M-AFM probe scanned the samples of Ag<br />
film 20 times with different standoff distances ranging from<br />
0 to 1000 nm. Fig. 5B shows the relationship between<br />
measured voltage values and the standoff distances.<br />
1.453 V 1.462 V<br />
1000 nm<br />
200 nm<br />
Ag film<br />
Ag film<br />
Silicon substrate<br />
Silicon substrate<br />
Stand-off distance (nm)<br />
C. Discussion<br />
As shown in Fig. 4B, the measured voltage is monotone<br />
increasing when the thickness of oxide layer is smaller than<br />
60 nm. However, when the thickness of oxide layer covered<br />
on the Ag film becomes larger than 60 nm, the measured<br />
voltage almost keeps constant regardless of different<br />
thickness of oxide layer. This result illustrates that the<br />
electrical property of Ag film under the oxide layer would<br />
affect the reflected microwave signals and thus can be<br />
extracted from the measured voltage when the SiO 2 layer is<br />
thinner than 60 nm. However, if the thickness of oxide layer<br />
is larger than 60 nm, the microwave signals will spread to<br />
other directions rather than penetrate the oxide layer to sense<br />
the covered sample. Thereby, the electrical property of the<br />
sample under a thick oxide layer can not be extracted from<br />
the measured voltage. In the calibration experiment, the<br />
similar phenomenon was observed. When the standoff<br />
distance is larger than 200 nm, the change in the measured<br />
voltage becomes very small. It means that the effective<br />
detection range of the M-AFM probe tip in air is almost 3<br />
times larger than that in the SiO 2 layer. The reason can be<br />
explained as that microwave signals can propagate more<br />
easily in the air than in the oxide layer due to the dielectric<br />
attenuation.<br />
The results suggest that the M-AFM can be used to<br />
measure the electrical property of material under a thin oxide<br />
layer, but the thickness and electromagnetic parameters of<br />
the oxide layer should be considered in a quantitative<br />
measurement.<br />
IV. CONCLUSION<br />
We carried out a group of experiment to verify the<br />
M-AFM with the capacity of measuring the electrical<br />
information of underlying materials. Some special samples<br />
with different thickness of dielectric films (SiO 2 ) which<br />
plays the role of oxide layer creating on the material surface<br />
were fabricated. The thickness of oxide-layer is from 20 nm<br />
to 100 nm with 20 nm increase in this work. As the results<br />
shown, the M-AFM probe can sense the electrical<br />
information of measured materials under the oxide layer with<br />
a limited thickness of 60 nm.<br />
ACKNOWLEDGMENT<br />
This work was supported by the Japan Society for the<br />
Promotion of Science under Grants-in-Aid for Scientific<br />
Research (A) 20246028 and (S) 18106003.<br />
Fig. 5. A: schematic diagram of the scanning process for Ag<br />
film with different standoff distance; B: the relationship<br />
between the measured voltage values and the standoff<br />
distances.<br />
REFERENCES<br />
[1] J.J. Kopanski, J.F. Marchiando, and J.R. Loweny, Scanning capacitance<br />
microscopy measurements and modeling: Progress towards dopant profiling<br />
of silicon, Journal of Vacuum Science and Technology B: Microelectronics<br />
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and Nanometer Structures, Vol. 14, No. 1, pp. 242-247, (1996). <br />
[2] Y. Martin, D.W. Abraham and H. K. Wickramasinghe, High-resolution<br />
capacitance measurement and potentiometry by force microscopy, Applied<br />
Physics Letters, Vol. 52, No. 13, pp. 1103-1105, (1988).<br />
[3] M. Nonnenmacher, M.P.O’ Boyle and H.K. Wickramasigh, Kelvin<br />
probe force microscopy, Applied Physics Letters, Vol. 58, No. 25, pp.<br />
2921-2923, (1991).<br />
[4] Y. Martin and H. K. Wickramasinghe, Magnetic imaging by “force<br />
microscopy” with 1000 Å resolution, Applied Physics Letters, Vol. 50, No.<br />
20, pp. 1455-1457, (1987).<br />
[5] M. Petzold, J. Landgraf, M. Füting and J.M. Olaf, Application of atomic<br />
force microscopy for microindentation testing, Thin Solid Films, Vol. 264,<br />
No. 2, pp. 153-158, (1995).<br />
[6] K. Yamanaka and S. Nakano, Ultrasonic atomic force microscopy with<br />
overtone excitation of cantilever, Japanese Journal of Applied Physiscs Part<br />
1, Vol. 35, No. 6B, pp. 3787-3792, (1996).<br />
[7] F. Duewer, C. Gao, I. Takeuchi, and X.D. Xiang, Tip-sample distance<br />
feedback control in a scanning evanescent microwave microscope, Applied<br />
Physics Letters, Vol. 74, No. 18, pp. 2696-2698, (1999).<br />
[8] M. Tabib-Azar and D. Akiwande, Real-time imaging of semiconductor<br />
space-charge regions using high-spatial resolution evanescent microwave<br />
microscope, Review of Scientific Instruments, Vol. 71, No. 3, pp. 1460-1465,<br />
(2000).<br />
[9] Y. Ju, M. Saka and H. abé, NDI of delamination in IC packages using<br />
millimeter-waves, IEEE Transactions on Instrumentation and Measurement,<br />
Vol. 50, No. 4, pp. 1019-1023.<br />
[10] Y. Ju, H. Sato and H. Soyama, Fabrication of the tip of GaAs<br />
microwave probe by wet etching, Proceedings of interPACK2005, Paper No.<br />
73140, CD-ROM (2005).<br />
[11] Y. Ju, T. Kobayashi and H. Soyama, Fabrication of a GaAs microwave<br />
probe used for atomic forcemicroscope, Proceedings of interPACK2007,<br />
Paper No. 33613, CD-ROM (2007).<br />
[12] Y. Ju, T. Kobayashi and H. Soyama, Development of a nanostructural<br />
microwave probe based on GaAs, Microsystem Technologies, Vol. 14, No. 7,<br />
pp. 1021-1025, (2008).<br />
[13] L. S. Liu and Y. Ju, Nondestructive measurement and high-precision<br />
evaluation of the electrical conductivity of doped GaAs wafer using<br />
microwaves, Review of Scientific Instruments, Vol. 81, No. 124701, pp.<br />
124701-1-4, (2010).<br />
338
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<br />
Amplitude Enhancement Using Vibration Mode<br />
Localization with A Single Micro-mechanically<br />
Coupled Beam-shaped Resonator Array<br />
Keisuke Chatani 1 , Dong F. Wang 1 , Tsuyoshi Ikehara 2 , and Ryutaro Maeda 2<br />
1<br />
Micro Engineering & Micro Systems Laboratory, Ibaraki University (College of Eng.), Hitachi, Ibaraki 316-8511, Japan<br />
(Tel: +81-294-38-5024; Fax: +81-294-38-5047; E-mail: dfwang@mx.ibaraki.ac.jp)<br />
2<br />
Ubiquitous MEMS and Micro Engineering Research Center (UMEMSME), AIST, Tsukuba, Ibaraki 305-8564, Japan<br />
Abstract- The use of vibration mode localization in arrays of<br />
micro-mechanically coupled, nearly identical beam-shaped resonators<br />
has been studied for ultrasensitive mass detection and analyte<br />
identification. Eigenstate shifts that are 3 to 4 times (compared to<br />
single resonator), and orders (compared to resonator array) of<br />
magnitude greater than corresponding shifts in resonant frequency for<br />
an induced mass perturbation are theoretically analyzed, from the view<br />
points of geometrical design of the coupling overhang, cantilever<br />
length, as well as number of the identical coupled cantilevers.<br />
Furthermore, the shifts in eigenstates are unique to the resonator to<br />
which the stiffness or mass perturbation is induced, therefore<br />
providing a characteristic “fingerprint” that identifies the particular<br />
resonator where the stiffness or mass perturbation is induced.<br />
Keywords- Vibration mode localization, Eigenstate shifts,<br />
Amplitude enhancement, Ultrasensitive mass detection, Analyte<br />
identification, Coupled resonator array, Coupling overhang<br />
I. VIBRATION MODE LOCALIZATION<br />
In resonant frequency based sensors the output corresponds<br />
to a shift in the resonant frequency of a vibrating<br />
micromechanical structure when subjected to small<br />
perturbations in either its stiffness or mass. The most sensitive<br />
micro cantilever based mass detection experiments using the<br />
frequency-shift approach have reported attogram level<br />
detection in ultrahigh vacuum environment [1-3] and<br />
femtogram level detection under ambient conditions [4-5].<br />
In contrast, the concept of using Anderson or vibration<br />
mode localization [6-13] in any array of nearly identical<br />
coupled resonators has also been proposed as a eigenstate-shift<br />
based sensing mechanism in recent years in coupled micro<br />
cantilevers under ambient conditions [6, 14-15].<br />
Some advantages of mode localized sensing can be listed<br />
below. Firstly, times or orders of magnitude in parametric<br />
sensitivity of micromechanical mass detection compared to the<br />
conventional frequency-shift approach can be obtained.<br />
Secondly, such sensors can offer the important advantages to<br />
intrinsic common mode rejection that renders it less susceptible<br />
to false-positive readings that frequency-shift based sensors.<br />
Thirdly, both the ultra sensitive detection and analyte<br />
identification of small perturbation can be achieved at same<br />
time with a single coupled resonator array.<br />
While many studies of mode localization in coupled<br />
structures and arrays of coupled resonators have been<br />
performed, the question of whether this phenomenon can be<br />
used in a sensing capacity has not been examined<br />
systematically.<br />
This work however, first theoretically studies the effects of<br />
geometrical design of the coupling overhang, cantilever length,<br />
as well as number of the identical coupled cantilevers on the<br />
magnitude enhancement by means of hypothesizing a small<br />
mass perturbation, which binds to cantilever surface due to<br />
molecule specific interactions. A preliminary evaluation has<br />
been then carried out by using microfabricated coupled<br />
beam-shaped resonator arrays.<br />
II.<br />
PHYSICS OF THE AMPLITUDE ENHANCEMENT<br />
A. Vibration localization in coupled two-resonator<br />
array<br />
A schematic and a discretized model of two identical<br />
beam-shaped cantilevers coupled by an overhang are shown in<br />
Fig. 1(a) and 1(b), respectively. Each cantilever is modeled as a<br />
damped simple harmonic oscillator, while the effect of the<br />
overhang coupling is modeled as spring connecting the two<br />
oscillators.<br />
Considering first the case of two initially identical<br />
cantilevers, the eigenvalue governing the undamped free<br />
oscillations of the system can be written as follows [6]:<br />
⎡<br />
⎢<br />
⎣<br />
−<br />
+<br />
/1/<br />
1<br />
− KK<br />
1 ⎤ cKK c<br />
= λuu<br />
+ KK + δ )1<br />
⎥<br />
(1)<br />
2<br />
2 ⎦ cKK c /<br />
where K 1 (=K), M 1 (=M) and K 2 (=K), M 2 (=M) are, respectively<br />
the bending stiffness and suspended mass of the two cantilevers,<br />
while δ represents the ratio of the effect mass ( Δ M) being<br />
detected to the single cantilever mass (M) . Kc is the stiffness of<br />
the overhang coupling the two cantilevers.<br />
339
After analyzing the two conditions of δ = 0 and δ ≠ 0, it can<br />
be seen that the relative change in normalized eigenstate is<br />
given by<br />
0<br />
−uu<br />
ii ⎛ 1 1 ⎞<br />
⎜ += ⎟δ<br />
0<br />
, i = 2,1 (2)<br />
u ⎝ 4 /4 KK ⎠ c<br />
i<br />
while the relative change in the eigenvalue or resonance<br />
frequency of a single cantilever is given by<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
−1 ~~ = ω<br />
2<br />
where u is a normalized eigenstate ( u i<br />
= 1) of the system<br />
representing the tip amplitudes of each cantilever of the array at<br />
the corresponding eigenfrequency, ω is an eigenfrequency of<br />
the system, and M and K are the mass and stiffness matrices of<br />
the system, respectively, given by the following Equations (5)<br />
and (6), respectively.<br />
uu (4) K<br />
− λλ−<br />
=<br />
λ 2<br />
0<br />
0 δ<br />
Noted that the perturbed eigenstates U i , i=1, 2 start becoming<br />
localized in the sense that in each eigenstate one cantilever<br />
oscillates more than the other.<br />
Equation (2), which defines the sensed quantity in this<br />
sensing paradigm, suggests that simply by decreasing the<br />
scaled coupling between the two cantilevers Kc/K, the relative<br />
changes in eigenstates can be made orders of magnitude greater<br />
than the relative change in eigenvalue of a single cantilever.<br />
(a)<br />
(3)<br />
⎡M<br />
⎢<br />
⎢<br />
0<br />
M =<br />
⎢<br />
⎢<br />
⎣<br />
1<br />
~ 2<br />
⎡<br />
⎢<br />
K = ⎢<br />
⎢<br />
⎢<br />
⎣<br />
1<br />
−<br />
~ c 2<br />
0 L 0 ⎤<br />
M L 0<br />
⎥<br />
⎥<br />
⎥<br />
(5)<br />
MOM<br />
⎥<br />
00<br />
L n Δ+ MM<br />
⎦<br />
−+<br />
KKK<br />
cc<br />
L 0 ⎤<br />
+ KKK<br />
⎥<br />
c L 0<br />
⎥ (6)<br />
MOM<br />
⎥<br />
⎥<br />
00<br />
+−<br />
KKK<br />
cnc<br />
⎦<br />
(b)<br />
K1<br />
Kc<br />
K2<br />
where M i and K i represent the mass and stiffness, respectively,<br />
of each cantilever. Solving Equation (4) when Δ M = 0 and M i<br />
= M, K i = K yields n eigenstates of the initially perfectly<br />
ordered system, while the primary mode consists of all<br />
cantilevers vibrating in phase with identical amplitude.<br />
Kc<br />
Kc<br />
C1<br />
M1<br />
X1<br />
Fig. 1. (a) Schematic of the coupled 2-cantilever resonator array with a<br />
mass perturbation placed at the end of one beam-shaped cantilever, and (b)<br />
simplified model of the coupled 2-cantilever oscillator array.<br />
B. Vibration localization in coupled n-resonator array<br />
A discretized model of identical cantilevers coupled by<br />
overhangs in a large array is shown in Fig. 2. Considering a<br />
perfect array of identical spring-mass oscillators (cantilevers)<br />
with each oscillator connected to its neighbor by a coupling<br />
spring, the sensitivity to mass added of the eigenstates of the<br />
coupled array can then be estimated as follows [7]:<br />
C2<br />
M1+ΔM<br />
X2<br />
III.<br />
K1 K2 Kn<br />
M1<br />
C1 C2 Cn<br />
X1 X2 Xn<br />
Fig. 2. Schematic of the coupled n-cantilever resonator array.<br />
THEORETICAL ANALYSIS OF COUPLED BEAM-SHAPED<br />
RESONATOR ARRAY<br />
Fig. 3 shows the schematic of a single weakly coupled array,<br />
corresponding to a perfect array of 15 identical spring-mass<br />
beams with each beam connected to its neighbor by an<br />
overhang (coupling spring). The geometrical size of the<br />
overhang is defined as a times b, as defined also in Fig. 3.<br />
M2<br />
Mn+⊿M<br />
340
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
the amplitude enhancement. The design No.4 (4a : 4b)<br />
undergoes 3.81 times (compared to single resonator) of<br />
magnitude greater than corresponding shifts in resonant<br />
frequency of cantilever No.14 at mode 15. While for cantilever<br />
No.13 at mode 1, still 2.5 times of magnitude can be obtained.<br />
(a)<br />
Fig. 3. The schematic of coupled beam-shaped 15-resonator array with a<br />
mass perturbation of 10 pg applied to cantilever No.14, and the size definition<br />
of coupling overhang.<br />
(b)<br />
Fig.4. Micromechanically coupled beam-shaped 15-resonator array<br />
simulated by using CoventorWare TM software, where (a) and (b): 1 st mode and<br />
15 th mode before mass perturbation; ( a’) and (b’ ): 1 st mode and 15 th mode after<br />
mass perturbation.<br />
In order to estimate the sensitivity to small mass<br />
perturbation of the eigenstates of the coupled array, theoretical<br />
analysis has been conducted for eigenstate shifts with and<br />
without small perturbation using Conventor Ware TM software.<br />
Fig. 4 shows the eigenstate shifts in vibration mode 1 and mode<br />
15, respectively, before and after a mass perturbation of 10 pg<br />
was induced. As shown in Fig. 3, the mass perturbation was<br />
induced on the tip of the cantilever No.14, and the vibration<br />
localization at cantilever No.13, No.14, and No.15 were thus<br />
analyzed. Fig. 5 (a) and (b) show the relative changes of<br />
amplitude due to the mass perturbation as a function of<br />
vibration modes for a fifteen coupled resonator array with a<br />
coupling overhang of design No.4 as defined in Table 1. The<br />
resonant frequency of a single cantilever as a function of<br />
vibration mode is also drawn in Fig. 5 for enhancement<br />
comparison. It is noticed that mode 15 undergoes the greatest<br />
change of 67.43 %, and both the mode 1 and mode 2 also show<br />
a relatively great changes. Table 1 summarizes the effects of<br />
five kinds of geometrical designs of the coupling overhang on<br />
Fig. 5. The relative change (%) of eigenstate (amplitude) due to the mass<br />
perturbation as a function of vibration modes, with a relation to a geometrical<br />
design of coupling overhang as 4a : 4b, where 5 (b) is a magnified figure of 5 (a)<br />
for a lower range of relative change. The resonant frequency of a single<br />
cantilever as a function of vibration mode is also drawn for comparison.<br />
Table 1. The effect of different geometrical design of coupling overhang on the<br />
relative change (%) of amplitude before and after a small mass perturbation.<br />
341
IV.<br />
MICRO FABRICATION<br />
For fabricating the above mechanically-coupled<br />
beam-shaped resonator arrays, an SOI (silicon on insulator)<br />
wafer with a 2.5-μm-thick top silicon layer, 300-nm-thick SiO 2<br />
layer, and 400-μm-thick silicon substrate was used as a starting<br />
material, as shown in Fig. 6.<br />
Fig. 6. Typical process chart<br />
for the coupled bema-shaped resonator array system.<br />
The topside silicon was first thinned to 500 nm by reactive<br />
ion etching (RIE) using SF 6 , and then patterned by lithography<br />
and etched using a deep reactive ion etching (deep RIE) to form<br />
the mechanically-coupled cantilever pattern. The substrate<br />
silicon was isotropically etched by RIE through the etching<br />
window of insulating SiO 2 . The coupled cantilever structures<br />
were released by wet etching of SiO 2 in HF and following<br />
supercritical point drying. Several resonator arrays are<br />
fabricated with micromechanically-coupled overhangs (design<br />
No. 4), as typically shown in Fig. 7.<br />
Fig. 7. Typical micrograph of the fabricated coupled 5-resonator arrays.<br />
Spectrum<br />
Analysis<br />
Function<br />
Generater<br />
TV<br />
monitor<br />
Laser<br />
Doppler<br />
Oscillator<br />
PZT Plate<br />
CCD<br />
Vacuum<br />
(~7.5Pa)<br />
Fig. 8. Experimental setup for vibration mode localization characterizations.<br />
Lens<br />
1<br />
2<br />
3<br />
4<br />
5<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
V. PRELIMINARY CHARACTERIZATIONS<br />
All measurements were performed in a vacuum of ~7.5 Pa<br />
at room temperature, as shown in Fig. 8. The sample was<br />
mounted on a piezoelectric ceramic plate, which can be<br />
vibrated by applying an AC voltage using a signal generator.<br />
The vibration of the cantilevers was measured using a laser<br />
Doppler system. The frequency responses were measured using<br />
frequency counters.<br />
The measured resonant frequency and corresponding<br />
amplitude under vibration mode localization has been typically<br />
shown in Fig. 9 for cantilever No.5. It can be seen that the<br />
amplitude response (12.7 dBm) was greatly magnified by<br />
vibration mode localization when driven by its resonant<br />
frequency of 209.50 kHz. As summarized in Fig. 10, although<br />
cantilevers No.3, No.4, and No.5 show the same resonant<br />
frequency of 209.50 kHz and are thus believed to be fabricated<br />
geometrically perfect, cantilever No.5 is connected by one<br />
coupling overhang rather than two like the others. This is the<br />
only different point among the above three cantilevers and<br />
might account for why cantilever No.5 displays a relatively<br />
great amplitude by vibration mode localization. Therefore,<br />
cantilever No.5 is suitable to be used as a detecting cantilever,<br />
and the neighbored cantilever No.4 can then be used as a<br />
sensing one for adding a small mass perturbation. A similar<br />
result can also be observed between cantilever No.1 and No.2,<br />
which needs further studied. However, micro-fabrication errors<br />
will inevitably cause differences in the cantilevers.<br />
Amplitude (dB)<br />
Vibration power [dBm]<br />
10<br />
0<br />
-<br />
10<br />
-<br />
20<br />
-<br />
30<br />
-<br />
40<br />
-<br />
50<br />
-<br />
60<br />
-<br />
70<br />
-<br />
80<br />
-<br />
90<br />
Cantilever No.5<br />
Driving frequency [kHz]<br />
Fig. 9. The amplitude response (12.70 dBm) of cantilever No. 5 was greatly<br />
magnified by vibration mode localization when driven by its resonant<br />
frequency of 209.50 kHz.<br />
15<br />
10<br />
5<br />
0<br />
-5<br />
-10<br />
-15<br />
209<br />
0 1 2 3 4 5 6<br />
210.5<br />
209.5<br />
Cantilever's number<br />
Fig. 10. The amplitude response and the resonant frequency corresponding to<br />
cantilever’s number summarized from the measurements typically shown in the<br />
above Fig. 9.<br />
1st<br />
2nd<br />
3rd<br />
Ave<br />
Freq<br />
211<br />
210<br />
Frequency (kHz)<br />
342
VI.<br />
CONCLUSIONS<br />
The structural design of coupled beam-shaped resonator<br />
arrays to achieve a 5 times (compared to single resonator) of<br />
amplitude enhancement has been performed theoretically. The<br />
effect of different geometrical designs of the coupling overhang<br />
on the relative change (%) of amplitude shifts has been studied<br />
for each vibration mode before and after a small mass<br />
perturbation of 10 pg.<br />
A preliminary characterization using a micro-fabricated<br />
5-resonator array without a small mass perturbation has been<br />
further conducted. The measured amplitude response (12.7<br />
dBm) of cantilever No. 5 was greatly enhanced by vibration<br />
mode localization and can thus be used as the detecting<br />
cantilever, while the neighbored cantilever No.4 can then be<br />
used as the sensing one for next examination.<br />
ACKNOWLEDGEMENT<br />
Part of this work was supported by MEMS Inter<br />
University Network and performed in the Ubiquitous MEMS &<br />
Micro Engineering Research Center (UMEMSME) of National<br />
Institute of Advanced Industrial Science & Technology (AIST).<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
REFERENCES<br />
[1] K. L. Ekinci, X. M. H. Huang, and M. L. Roukes, Appl. Phys. Lett.<br />
84, 4469, 2004.<br />
[2] T. Ono, X. Li, H. Miyashita, and M. Esashi, Rev. Sci. Instrum. 74,<br />
1240, 2003.<br />
[3] Z. Davis and A. Boisen, Appl. Phys. Lett. 87, 013102, 2005.<br />
[4] H. Sone, Y. Fujinuma, and S. Hosaka, Jpn. J. Appl. Phys. Part 1 43,<br />
3648, 2004.<br />
[5] B. Ilic, D. Czaplewski, M. Zalalutdinov, and H. G. Craighead, J.<br />
Vac. Sci. Technol. B 19, 2825, 2001.<br />
[6] M. Spletzer, A. Raman, A.Q. Wu, X. Xu, and R. Reifenberger, Appl.<br />
Phys. Lett. 88, 254102, 2006.<br />
[7] M. Spletzer, A. Raman, H. Sumali and J.P. Sullivan, Appl. Phys.<br />
Lett., 92, 114102, 2008.<br />
[8] P. W. Anderson. Phys. Rev. 109, 1492, 1958.<br />
[9] C. Pierre, D. M. Tang, and E. H. Dowell, AJAAJ. 25, 1249, 1987.<br />
[10] O. O. Bendiksen, AJAAJ. 25, 1492, 1987.<br />
[11] M. Sato, B. E. Hubbard, A. J. Sievers, B. Ilic, D. A. Czaplewski, and<br />
H. G. Craighead, Phys. Rev. Lett. 90, 044102, 2003.<br />
[12] E. Buks and M. L. Roukes, J. Microelectromech. Syst. 11, 802,<br />
2002.<br />
[13] M. Napoli, W. H. Zhang, K. Turner, and B. Bamieh, J.<br />
Microelectromech. Syst. 14, 295, 2005.<br />
[14] L. Nicu and C. Bergaud. J. Micromech. Microeng. 14, 727, 2004.<br />
[15] A. Qazi, D. Nonis, A. Pozzato, M. Tormen, M. Lazzarino, S.<br />
Carrato, and G. Scoles, Appl. Phys. Lett. 90, 173118, 2007.<br />
343
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Stress Identification of Thin Membrane Structures by<br />
Dynamic Measurements<br />
Steffen Michael 1 , Christoph Schäffel 1 , Sebastian Voigt 2 , Roy Knechtel 3<br />
1 IMMS GmbH, Ehrenbergstr. 27, 98693 Ilmenau, Germany<br />
2 TU Chemnitz, Chair in Microsystems and Precision Engineering, Reichenhainer Str. 70, 09107 Chemnitz, Germany<br />
3 X-FAB Semiconductor Foundries AG, Haarbergstr. 67, 99097 Erfurt, Germany<br />
A fast identification method of membrane stresses is<br />
investigated for an early stage of the manufacturing process.<br />
The approach consists of performing optical measurement of<br />
the out-of-plane modal responses of the membrane. This<br />
information is used in an inverse identification algorithm<br />
based on a FE model by an optimization.<br />
I. INTRODUCTION<br />
The development of the two criteria costs and reliability<br />
is essential for the further growth of the MEMS market like<br />
microphones. Efficient test procedures on wafer level can<br />
reduce costs significantly by the detection of faulty sensors<br />
before the subsequent packaging and assembly steps. The<br />
presented method deals with an approach for a fast and<br />
accurate stress identification of thin membranes by using<br />
the sensitivity of their modal frequencies versus the stress.<br />
MEMS devices usually do not permit direct parameter<br />
measurement of mechanical parameters. The indirect<br />
parameter identification by modal frequencies was first<br />
presented in [1], [2]. Up to now the approach is used mostly<br />
for the identification of geometrical parameters like<br />
membrane thicknesses [3], [4]. In this case the approach<br />
competes against other methods like optical ones. In<br />
contrast the method has a unique feature with regard to the<br />
identification of tensile stressed membranes like<br />
microphones – another non-destructive method on wafer<br />
level is not known.<br />
Perforated circular SiN membranes with a thickness of<br />
300 nm and a diameter of 1000 µm are investigated. The<br />
perforation is required by the technology – the membrane<br />
structure is deposited on a sacrificial layer which is<br />
removed at the end of the processing through the<br />
perforation holes. The formed cavity with a height of 1µm<br />
causes a squeeze film damping in conjunction with an<br />
absent resonance rice under ambient atmosphere.<br />
Correspondingly the measurements are done in a vacuum<br />
prober.<br />
II. HARDWARE SETUP<br />
The measurement setup consists on a vacuum probe<br />
station from Cascade and a laser Doppler vibrometer<br />
integrated in the Micro System Analyzer MSA500 from<br />
Polytec. The laser beam of the vibrometer scans<br />
automatically over a user defined grid at the surface of the<br />
membrane.<br />
Fig. 1: Measurement setup<br />
The vibration of passive devices like the membrane<br />
structures is realized by electrostatic forces. A probe needle<br />
is connected to a high voltage (up to 400V) excitation signal<br />
controlled by a chirp signal of the measurement system. The<br />
needle is positioned above the device surface. With respect<br />
to a high excitation force the gap between the needle and<br />
the membrane is smaller than 100µm. The setup permits the<br />
excitation of modal frequencies up to 4MHz.<br />
III. IDENTIFICATION ALGORITHM<br />
The approach can be subdivided into three different<br />
phases. First of all a sensitivity analysis has to be done to<br />
check whether the modal frequencies are sensitive versus<br />
the interesting parameters. In case of a multidimensional<br />
problem the orthogonality of the parameter space has to be<br />
tested furthermore.<br />
Following to the sensitivity analysis a characterization<br />
phase is done. Frequency response functions (FRF) are<br />
measured with a fine grid of measurement points to check<br />
the mode shapes and adapt the finite element (FE) model if<br />
needed.<br />
The results shown here refer to measurement data of the<br />
characterization phase. In case of testing complete wafers<br />
the measurement time should be minimized. The<br />
measurement time depends proportional on the number of<br />
measurement points. The identification approach is based<br />
on frequency values which permits the reduction of<br />
measurement points to one.<br />
344
o<br />
o<br />
o<br />
o<br />
o<br />
o<br />
o<br />
o<br />
Check applicability<br />
Analytic / FE- modelling<br />
Check sensitivity<br />
Check orthogonality<br />
Development of test structures<br />
Characterization<br />
Fine grid of measurement points<br />
Selection of frequency modes for<br />
identification<br />
Parametet identification & validation<br />
Adaption of FE model<br />
Wafer-Test<br />
Fig. 2: Phases of the parameter identification<br />
The measurement time of a one point measurement is 2<br />
seconds. The measurement respectively software system is<br />
not yet optimized, the lower measurement time limit given<br />
by physics is about 200 milliseconds.<br />
Precondition for the identification is on one hand the<br />
measurement unit which delivers a FRF, and the simulation<br />
unit with a parameter matrix as result on the other hand. The<br />
automatic identification is done by a tool implemented in<br />
C++ with respect to a fast data processing. The<br />
identification tool can be structured into three submodules.<br />
The frequency values has to be extracted from the measured<br />
FRF which is done in one submodule, and the parameter<br />
matrix is approximated by usually polynomials in another<br />
submodule due to a fast and efficient data handling. Based<br />
on an user defined accuracy (default value 0.1%) the degree<br />
of the polynomial is selected by the program.<br />
Finally the optimization respectively identification is<br />
realized by the nonlinear least square method.<br />
Measurement<br />
system<br />
Frequency<br />
response<br />
Peak detection<br />
Identification tool<br />
Optimization<br />
FE-Simulation<br />
Polynomial<br />
approximation<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
first step a conventional algorithm searches for local maxima<br />
considering the estimated signal-to-noise ratio (SNR). At the<br />
peaks found, starting values for a nonlinear least square fit to<br />
the Lorentzian function<br />
Parameter<br />
matrix<br />
i<br />
2<br />
, ih<br />
LfL<br />
, ip 2 2<br />
,<br />
)(<br />
+−<br />
, i hi<br />
)(<br />
f<br />
= (1)<br />
fff<br />
with the peak amplitude L p,i , the peak frequency f p,i and the<br />
half-width f h,i. of the ith peak are estimated. The iterative<br />
fitting procedure based on Levenberg-Marquardt algorithm<br />
eliminates wrongly preselected peaks and delivers the peak<br />
parameter including the quality factor.<br />
A. FE Modeling and Simulation<br />
The FE model which delivers the parameter matrices is<br />
implemented in Ansys. The ratio thickness to lateral<br />
dimension of the membrane leads to a modeling by twodimensional<br />
shell elements. The default mesh of the<br />
membrane perforated by several thousand holes will be<br />
irregular. To prevent such an inefficient irregular mesh<br />
substructures are generated. Square areas with a centered<br />
hole permit a regular meshing.<br />
Fig. 4: FE modell with prestructured membrane elements<br />
A prestressed modal analysis as well as a prestressed<br />
harmonic analysis is performed. The multitude of small<br />
structures causes a large number of finite elements<br />
respectively nodes. With regard to the measurement time the<br />
membrane symmetry is used by the calculation of a quarter<br />
model. Symmetric boundary conditions are applied to the<br />
static analysis. The modal analysis is executed with three<br />
load steps with different symmetry conditions at the x and y<br />
axes (symmetric/symmetric, asymmetric/symm. and<br />
asym./asym.) to deliver all modal frequencies .<br />
For the modeling of the squeeze film damping the<br />
corresponding element types of Ansys are used. The macro<br />
RMFLVEC.MAC which extracts the damping parameters<br />
from the modal frequencies is adapted to the quarter model<br />
with the multiple loadsteps.<br />
Sensor parameter<br />
Fig. 3: Structure of the parameter identification<br />
From the measured FRF, the peak frequency values are<br />
extracted automatically by a two level algorithm. Within a<br />
a) f 11 b) f 12<br />
Fig. 5: Simulated modal frequencies<br />
345
IV.<br />
A. Simulation and Sensitivity Analisys<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
STRESS IDENTIFICATION<br />
With respect to the identification phases a sensitivity<br />
analysis is done for the membrane structure. Parameters<br />
which have to be considered beside the interested ones are<br />
parameters with relevant tolerance ranges. The membrane<br />
thickness is such a parameter – due to technological reasons<br />
the thickness varies within a range of ±5%.<br />
(∂ f 1<br />
/∂ h) ∆ h/f 1<br />
[%]<br />
6<br />
5<br />
4<br />
3<br />
2<br />
1<br />
f 1<br />
[kHz]<br />
60<br />
50<br />
40<br />
30<br />
20<br />
10<br />
0<br />
0.42<br />
Fig. 6: First modal frequency versus membrane thickness and stress<br />
As is apparent from Fig. 6 which show the results of the<br />
two dimensional parameter simulation for the first modal<br />
frequency the most sensitive parameter is the stress. An<br />
approximation of the functional dependency is done with<br />
regard to a quantitative analysis. The default expansion is a<br />
polynomial one. In this case rational functions are used for<br />
the stress motivated by the plate theory [5] on the one hand<br />
and the curve characteristic of Fig. 6 on the other hand The<br />
frequency mode f i,j is given by<br />
with the membrane thickness h and the stress s.<br />
Based on partial derivatives of the approximated course<br />
of the function the sensitivity of the modal frequencies<br />
versus the parameters is determined. Fig. 7 shows the<br />
sensitivity normed on the maximum thickness variation of<br />
5%. In case of a tensile stressed membrane the varying<br />
thickness can be neglected – a relevant sensitivity of the<br />
modal frequencies versus the thickness is given only in case<br />
of a stress-free or compressive stressed membrane.<br />
0<br />
0 2 4 6 8 10 12 14 16 18 20<br />
s [MPa]<br />
Fig. 7: Normed sensitivity of the first modal frequency versus membrane<br />
thickness<br />
B. Measurement Results<br />
Measurements are done at three different wafers at a<br />
pressure range between 0.005 mbar and 0.1 mbar. The<br />
pressure range is determined by the resonance rice on one<br />
hand and a minimal peak width to be detectable by the FFT<br />
on the other hand. The measured quality factors show a<br />
good accordance with the simulated ones given by the<br />
harmonic analysis of the FE model.<br />
0.41<br />
20<br />
0.4<br />
15<br />
10<br />
0.39<br />
5<br />
z [µm] 0.38 0<br />
s [MPa]<br />
10 5 measurement data<br />
simulated data<br />
10 4<br />
10 3<br />
2/1<br />
ji 1,<br />
2<br />
++=<br />
3<br />
),(),(),(<br />
sjipsjipjip f<br />
3/1<br />
+<br />
4<br />
5<br />
6<br />
⋅++<br />
),( shjiphjip sjip 10 2<br />
(2)<br />
10 -5 10 -4 10 -3 10 -2 10 -1 10 0<br />
2/1<br />
3/1<br />
p [mbar]<br />
7<br />
8<br />
),(<br />
⋅+⋅+<br />
shjipshjip<br />
Q-factor<br />
Fig. 8: Q-factor versus ambiance pressure<br />
The first three modal frequencies are used for the<br />
identification of the membrane stress. Mode shapes are<br />
investigated at some samples to guarantee the right<br />
classification of the frequency peaks to the corresponding<br />
modes.<br />
Fig. 9: Measured mode shape f 1,1<br />
346
C. Identification Results<br />
The identified tensile stresses at 36 measured dies vary<br />
between 24MPa and 81MPa due to their different position at<br />
the test wafers.<br />
TABLE 1<br />
IDENTIFIED STRESS OF MEMBRANE SAMPLES<br />
f 1,1<br />
[kHz]<br />
f 1,2<br />
[kHz]<br />
f 2,2<br />
[kHz]<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
s r [MPa]<br />
55.2 88.1 117,3 28.75 ± 0.33<br />
60.9 97.2 130.5 34.99 ± 0.16<br />
69.9 109.7 147.1 45.57 ± 0.98<br />
69.1 110.2 148.0 45.50 ± 0.10<br />
[4] Michael, S. at al, “MEMS parameter identification on wafer level<br />
using laser Doppler vibrometry”, Smart Systems Integration 2007,<br />
Editor T.Gessner, VDE Verlag, 2007, pp. 321-328<br />
[5] Dickinson, S.M., “The Buckling and Frequency of Flexural<br />
Vibration of Rectangular Isotropic and Orthotropic Plates Using<br />
Raleigh’s Method”, Journal of Sound and Vibration, 1978, 61(1),<br />
pp. 1-8<br />
The identification is based on the first three modal<br />
frequencies which results in an over-determined problem<br />
which permits a quantitative evaluation of the identification<br />
results. Theoretically the stress values should be identically;<br />
practically measurement and modeling errors will cause<br />
different values. The particular stress values differ within a<br />
range of 2% which shows a good model quality.<br />
90<br />
80<br />
Wafer 1<br />
Wafer 2<br />
Wafer 3<br />
70<br />
s r<br />
[MPa]<br />
60<br />
50<br />
40<br />
30<br />
20<br />
0 20 40 60 80 100<br />
Die index<br />
Fig. 10: Identified stress at test wafers across the x axes<br />
V. CONCLUSION<br />
We have presented an approach for the fast and accurate<br />
stress identification of thin membranes which the uses the<br />
sensitivity of their modal frequencies versus stress. The<br />
approach is well suited for an efficient process control of<br />
stress sensitive membranes like microphones on wafer level.<br />
REFERENCES<br />
[1] Smith, N.F. et al, “Non-Destructive Resonant Frequency<br />
Measurement on MEMS Actuators”, 39 th Annual International<br />
Reliability Physics Symposium, Orlando, FL, USA, 2001,<br />
Proceedings, pp. 99-105<br />
[2] Tanner, D.M. et al: “Resonant frequency method for monitoring<br />
MEMS fabrication”, Reliability, Testing and Characterization of<br />
MEMS/MOEMS II, San Jose, CA, USA, 2003, Proceedings, pp.<br />
220-228<br />
[3] Gerbach, R. et al: „Numerical Identification of Geometric<br />
Parameters from Dynamic Measurement of Grinded Membranes<br />
on Wafer Level”, 7 th Conf. on Thermal, Mec.l and Multiphysics<br />
Simulation and Experiments in Micro-Electronics and Micro-<br />
Systems EuroSimE, Como, Italy, 2006, Proceedings, pp. 223-228<br />
347
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
Electrical and Mechanical Characterization<br />
of Lateral NEMS Switches<br />
R. Hinchet 1 , L. Montès 1 , G. Bouteloup 1 , G. Ardila 1 ,<br />
R. Parsa 2 , K. Akarvardar 2,3 , R. T. Howe 2 , H.-S. Philip Wong 2<br />
1 IMEP-LAHC, Grenoble Institute of Technology, MINATEC, 3 Parvis Louis Néel, 38016 Grenoble, France<br />
2 Department of Electrical Engineering, Stanford University, CA 94305, USA<br />
3 Present affiliation: Globalfoundries/SEMATECH, Albany, NY 12203, USA<br />
Ronan.Hinchet@minatec.inpg.fr Ph : +33 456 529 516<br />
Abstract: In this paper we present a study on the<br />
electrical and mechanical characterization of NEMS<br />
Nano Switches. Pull-in, pull-out voltages are in good<br />
agreement with the theoretical values. However some<br />
reliability and sticking problems are identified. To<br />
investigate furthermore the mechanical properties of<br />
the nanoswitch (NS) beam, we have developed a<br />
dedicated AFM methodology to extract the Young's<br />
modulus. A further improvement of the design is<br />
then realized based on the results of this study.<br />
Fig. 1. Structure of a NS. (Source (S), Gate (G), Drain (D)).<br />
Beam: 300nm width, 1100 nm thick and 20 µm length.<br />
Keywords: NEMS, Nano-switches, Young's<br />
modulus, mechanical characterization, AFM.<br />
I. INTRODUCTION<br />
Alternative device architectures and solutions might be<br />
required to continue the trend described by Moore's Law.<br />
NEMS Nano-Switch (NS) [1,2], may open ways for even<br />
more complex circuits using fewer components, while<br />
lowering the energy consumption due to the complete absence<br />
of electrical leakage in the off-state [3]. In this paper we study<br />
electrical and mechanical properties of such NS (Fig. 1). The<br />
NS concept is based on controlling the current through a<br />
mobile beam between electrodes similar to source (S) and<br />
drain (D) in a standard MOS device. A gate (G) voltage<br />
controls the lateral switch by electrostatically attracting the<br />
beam. According to theory, the non-linear behavior of<br />
electrostatic forces and Van der Waals forces between the<br />
beam and the drain should lead to an hysteresis of the I-V<br />
characteristics. The NS was fabricated using one-mask<br />
process. It started with a 1.5µm thick LTO and 1.1µm thick<br />
doped-poly-silicon deposition, followed by an RTA at<br />
1075°C. The poly-silicon was patterned using deep-UV<br />
lithography and a reactive-ion-etch (RIE) [4,5]. The<br />
underlying SiO2 sacrificial layer was then removed using a<br />
49% wet HF release to free-stand the beam. Critical Dry Point<br />
(CDP) was used to avoid sticking. We checked the beam<br />
release and CDP process with Scanning Electron Microscope<br />
(SEM) observations (Fig. 2). In most of the cases beams were<br />
correctly release. We just noted an over-etching (Fig. 3) on the<br />
edge of the connection pads which depends on the length and<br />
the concentration of the wet HF release step.<br />
Fig. 2. Scanning Electron Microscopy image of a NS.<br />
Fig. 3. Scanning Electron Microscopy image of the side of a NS.<br />
II.<br />
ELECTRICAL CHARACTERIZATION<br />
Electrical results clearly demonstrated the lateral actuation<br />
of the beam (Fig. 4), while showing slightly lower pull-in<br />
voltage Vpi and pull-out voltage Vpo (Fig. 5) than expected<br />
by theory [1,2]. We noted the importance of the atmospheric<br />
environment: care must be taken especially with humidity,<br />
working under nitrogen or vacuum ambient lead to an<br />
improved yield. We also experienced, on some NS, difficulties<br />
to switch on or off (Fig. 6), with part of the NS beam sticking<br />
or staying stuck to the drain.<br />
348
11-13 May 2011, Aix-en-Provence, France<br />
<br />
Experimentally, we first made a topographic image of the<br />
NS device. Then a force was applied at different points along<br />
the beam axis (Fig. 8). We used a specific XYZ feedback loop<br />
to ensure very precise XY localization of the AFM tip on the<br />
NS beam, the Z loop allowing for very precise control of the<br />
applied mechanical force on the NS beam.<br />
Fig. 4. SEM image of a NS at the ON state.<br />
Fig. 7. Schematic of the young<br />
modulus extraction model.<br />
Fig. 5. Experimental electrical characterization of a NS.<br />
Measurement parameters: V S=0V, V D=3V, compliance 1nA.<br />
Fig. 8. AFM image of a NS with the position of approach-retract curves.<br />
Fig. 6. Experimental electrical characterization of a NS.<br />
The beam staying stuck to the drain after ON state, as shown in Fig. 4.<br />
Measurement parameters: V S=0V, V D=3V, compliance 1nA.<br />
III. AFM MECHANICAL CHARACTERIZATION<br />
Mechanical properties of the polysilicon beam were<br />
investigated by vertical and local approach-retract curves<br />
obtained with an AFM. From the Euler-Bernoulli equation<br />
(eq. 1) we calculated the relation between the<br />
force applied by<br />
the AFM tip and the beam deflection ∆z (Fig. 7) [6], where x<br />
is beam axis, u is the position along x where the force is<br />
applied, E is the Young's modulus, F is the applied force, Iy is<br />
the beam moment (considering a rectangle cross section), l is<br />
the beam length, w is the beam width and t is the beam<br />
thickness. To calculate the solution to the equation 1 we<br />
considered the particular case where we applied the force at<br />
the end of the beam.<br />
Fig. 9. Approach-retract cur<br />
rve at point 1 (cf. Fig. 8).<br />
Fig. 10. Approach curves at<br />
points 1 to 7 (cf. Fig. 8).<br />
<br />
349
11-13 May 2011, Aix-en-Provence, France<br />
<br />
35% (Fig. 12). Finally the side and the edge are not exactly<br />
The approach-retract curves (Fig. 9 & 10) show the tip straight, thus changing the beam<br />
section shape and also Iy.<br />
cantilever deflection in function of the AFM<br />
Z piezoelectric<br />
position. During the approach, the AFM head goes down on<br />
the sample. The tip cantilever does not move until it touches<br />
the sample. In effect we observe a flat part on curves point 1<br />
to 5 (Fig. 8). The shift and the absence of<br />
this flat part on<br />
curves point 6 and 7 indicate that the sample<br />
was tilted. After<br />
touching the sample, the AFM head approach is compensated<br />
by the tip cantilever deflection at point 1 and by the beam<br />
deflection at point 7. So at point 1 the slop angle of the curve<br />
is about 45° because there is no beam to absorb the tip<br />
cantilever deflection which is the same that the AFM head<br />
position (after had touched the sample). Conversely at point 7<br />
a part of the AFM head movement is compensated by the<br />
beam deflection. So the tip cantilever deflection is lower and<br />
the slop angle of the curve is lower than at previous points.<br />
From these results, we calculated the force applied by the<br />
AFM tip on the beam (F AFM TIP ) and the deflection of beam<br />
(∆z beam ). We considered that we measured the Young modulus<br />
of the bending part of the beam (i. e. from the beam anchor to<br />
the AFM probe contact point). Thus in our model, the beam<br />
length depends on where we measured the<br />
Young modulus<br />
along the beam. From the curve ∆z beam (F AFM TIP) and using the<br />
beam design dimensions (400nm width and 1100 nm<br />
thickness), we extracted a beam Young's modulus E of 70 GPa<br />
(Fig. 11A). This value is much lower than<br />
the mean value<br />
from the literature, between 130 and 170 GPa depending on<br />
measurement methods, crystal orientation and testing devices<br />
(bulk, thin film, beam) [7,8,9].<br />
Fig. 11. Extraction of the beam young modulus. A)<br />
based on design<br />
dimensions. B) based on measured dimensions.<br />
Such a difference can be explained by<br />
the polysilicon<br />
deposition process [10] but the main reason<br />
is the difference<br />
on the real dimensions of the beam compared to the design.<br />
SEM images showed that the HF etching step to release the<br />
beam created an over-etching on the edge of the connection<br />
pads (Fig. 3) giving them flexibility. Moreover we extracted<br />
the accurate dimensions of beams which are thinner and<br />
narrower than expected due to fabrication process (polysilicon<br />
deposition and etching). Therefore the beam<br />
is more flexible,<br />
which explains the lower Vpi and Vpo observed previously.<br />
The beam length is an important parameter (see eq. 1) but<br />
above all the thickness and the width (Fig. 7) are very critical<br />
(affecting Iy) (eq. 1). A beam 110nm thinner<br />
than designed (-<br />
10%) generates a Young modulus miscalculation higher than<br />
Fig. 12. Young modulus miscalculation depending of the beam thickness<br />
error.<br />
Therefore, taking into account the real beam dimensions<br />
obtained by SEM and AFM, we found a beam Young's<br />
modulus of 140 GPa (Fig. 11B) ), which is in better agreement<br />
with the literature [11,12,13]. We noted the apparent variation<br />
of the Young modulus as a function of the distance from the<br />
beam anchor, even though the Young modulus is a material<br />
property supposed to be constant. This could be due to the<br />
influence of the indentation phenomenon and the elastic<br />
deformations of the material in our measurement method. To<br />
measure the Young modulus of the beam, we measured the<br />
absorption of the AFM cantilever deflection by the beam. At<br />
the end of the beam, this absorption is mainly due to the beam<br />
bending. The indentation phenomenon and the elastic<br />
deformations are negligible. On the other end, near the anchor,<br />
they are less negligible compared to the beam bending. Thus<br />
the total absorption is higher than the absorption due to the<br />
beam bending only, so that the<br />
beam appears more flexible<br />
than in reality, explaining the lower than expected extracted<br />
Young modulus. This explains the constant increase of the<br />
Young modulus with the distance from the beam anchor.<br />
Based on the same principle, we<br />
think that the influence of the<br />
over-etching is more pronounced near the anchor than at the<br />
end of the beam. It makes the beam appear more flexible than<br />
expected and therefore underestimation of the Young modulus<br />
again. Nevertheless this Young modulus measurement method<br />
allowed to improve the beam<br />
design and structure with<br />
different materials, to study the<br />
mechanical reliability of the<br />
beam, and is helpful to improve the electromechanical<br />
behavior of the NS device.<br />
350<br />
<br />
IV.<br />
IMPROVEMENT AND PERSPECTIVES<br />
AFM electrical characterization of the NS beam is<br />
important to better understand what happens at the nano-scale<br />
during a switching event. To perform in-situ AFM<br />
electromechanical characterization [14,15] we have improved<br />
the design of the chip, by routing the electrical signals of the<br />
different pads (Source, Gate and<br />
Drain) to the edge of the chip<br />
in order to observe the beam working without disturbing the<br />
AFM tip (Fig. 13). In addition, we have also changed the<br />
structure of the beam with the deposition of a thin layer of<br />
platinum (Fig. 14) on the pads and on all the sidewalls of the
eam to improve the electrical contacts and to solve the<br />
problems of oxidation caused by humidity which disturb and<br />
stick beams (the metal wall covering is explained in [4,5]).<br />
Thanks to platinum the reliability and the yield of NS devices<br />
was clearly improved [5].<br />
11-13 May 2011, Aix-en-Provence, France<br />
<br />
<br />
ACKNOWLEDGEMENT<br />
Samples were fabricated and designed by Stanford<br />
University. They were finalized and customized at the<br />
Grenoble Up-line Technological Platform (PTA) which is cooperated<br />
by CEA & CNRS within the framework of the<br />
French Basic Technology Research (BTR) network. This work<br />
has been partly supported by the EU through the Network of<br />
Excellence NANOFUNCTION FP7/ICT/NoE (95145). The<br />
author would like to thank Xavier Mescot, Martin Gri, Remy<br />
Lefevre and Xin Xu for their valuable help.<br />
REFERENCES<br />
Fig. 13. Optical image of the improved design with electrical contacts routed<br />
to the edge of the chip.<br />
Fig. 14. SEM image showing thin platinum layer coating on the NS side walls<br />
to improve the device structure.<br />
V. CONCLUSION<br />
The NS electrical characterization showed a correct<br />
behavior, while highlighting a few problems. A deeper<br />
analysis using AFM and SEM techniques showed small<br />
differences from the designed structure. NS mechanical<br />
properties were investigated and the beam Young's modulus<br />
was extracted taking account of the real sample characteristics<br />
and dimensions which are relevant. NS structure and design<br />
was improved. The reliability has been increased and the<br />
electrical behavior was better. Finally this innovative<br />
technique of characterization will help us to explore new<br />
fields of NEMS [16].<br />
[1] K. Akarvardar, et al., "Analytical Modeling of the Suspended-Gate FET<br />
and Design Insights for Low-Power Logic," IEEE Transactions On<br />
Electron Devices, vol. 55, no. 1, p. 48, 2008.<br />
[2] K. Akarvardar, et al., "Design Considerations for Complementary<br />
Nanoelectromechanical Logic Gates," IEDM, pp. 299-302, 2007.<br />
[3] S. Chong, et al., "Nanoelectromechanical (NEM) Relays Integrated with<br />
CMOS SRAM for Improved Stability and Low Leakage," in ICCAD<br />
International Conference on Computer-Aided Design, 2009.<br />
[4] D. Lee, et al., "Titanium nitride sidewall stringer process for lateral<br />
nanoelectromechanical relays," in MEMS 2010, IEEE 23rd International<br />
Conference, Hong Kong, 2010, pp. 456-459.<br />
[5] R. Parsa, et al., "Composite Polysilicon-Platinum Lateral<br />
Nanoelectromechanical Relays," in 14th Solid-State Sensors, Actuators,<br />
and Microsystems Workshop, Hilton Head, South Caroline, 2010, pp. 7-<br />
10.<br />
[6] W. C. Young and R. G. Budynas, Roark's Formulas for Stress and<br />
Strain, 7th ed., McGraw-Hill, Ed. 2002 .<br />
[7] K. R. Virwani, A. P. Malshe, W. F. Schmidt, and D. K. Sood, "Young’s<br />
modulus measurements of silicon nanostructures using a scanning probe<br />
system: a non-destructive evaluation approach," Smart Materials and<br />
Structures, vol. 12, no. 6, pp. 1028-1032, 2003.<br />
[8] M. A. Hopcroft, W. D. Nix, and T. W. Kenny, "What is the Young’s<br />
Modulus of Silicon," Journal Of Microelectromechanical Systems, vol.<br />
19, no. 2, pp. 229-238, 2010.<br />
[9] C. S. Oh, H. J. Lee, S. G. Ko, S. W. Kim, and H. G. Ahn, "Comparison<br />
of the Young’s modulus of polysilicon film by tensile testing and<br />
nanoindentation," Sensors and Actuators A, no. 117, p. 151–158, 2005.<br />
[10] S. Lee, et al., "The effects of post-deposition processes on polysilicon<br />
Young's modulus," Journal of micromechanics and microengineering,<br />
vol. 8, no. 4, pp. 330-337, 1998.<br />
[11] A. San Paulo, J. Bokor, and R. T. Howe, "Mechanical elasticity of single<br />
and double clamped silicon nanobeams fabricated by the vapor-liquidsolid<br />
method," Applied Physics Letters, no. 87, p. 053111, 2005.<br />
[12] J. Wang, Q. A. Huang, and H. Yu, "Young’s modulus of silicon<br />
nanoplates at finite temperature," Applied Surface Science, no. 255, p.<br />
2449–2455, 2008.<br />
[13] C. H. Cho, "Characterization of Young’s modulus of silicon versus<br />
temperature using a ‘‘beam deflection” method with a four-point bending<br />
fixture," Current Applied Physics, vol. 9, p. 538–545, 2009.<br />
[14] L. Montès, et al., "AFM Measurement of the Piezoelectric Properties of<br />
GaN and GaN/AlN/GaN Individual Nanowires," in MRS, San Francisco,<br />
2010.<br />
[15] L. Montès, et al., "Enhancing piezoresistivy and piezoelectricity in<br />
nanowire devices," in IEEE Nano 2010, 12th Nanowire Research Society<br />
Meeting, Seoul, 2010.<br />
[16] X. Xu, et al., "An improved AFM cross-sectional method for<br />
piezoelectric nanostructures properties investigation: application to GaN<br />
nanowires.," Nanotechnology, vol. 22, no. 10, p. 105704, 2011.<br />
351
11-13 May 2011, Aix-en-Provence, France<br />
<br />
A Dielectrophoretic Preconcentrator with Circular<br />
Microelectrodes for Biological Cells in Stepping<br />
Electric Fields<br />
Chun-Ping Jen and Ho-Hsien Chang<br />
Department of Mechanical Engineering,<br />
National Chung Cheng University,<br />
Abstract- The ability to enrich rare cells, e.g. circulating tumor<br />
cells (CTC), circulating fetal cells, and stem cells, has been an<br />
important issue in medical diagnostics and characterization. The<br />
main purpose of this investigation was to develop a handheld<br />
microdevice capable of the effective preconcentration of rare cells.<br />
Circular microelectrodes were designed to generate the stepping<br />
electric field by switching the electric field to an adjacent electrode<br />
pair by relays. The cancerous cells with positive dielectrophoretic<br />
response were not only conveyed but also concentrated toward the<br />
center of the circular microelectrodes because the<br />
high-electric-field region between the adjacent electrodes was<br />
gradually decreased in the direction of the stepping electric field.<br />
Numerical simulations of the electric fields were performed to<br />
demonstrate the concept of the proposed design. Moreover,<br />
enrichment of cervical cancer cells was successfully achieved and<br />
took about 160 seconds in the experiment with an approximate<br />
efficiency of 75%, when the peak-to-peak voltage of 16 volts, a<br />
frequency of 600 kHz and the time interval of relay switching with<br />
20 seconds were applied.<br />
Keywords: handheld; dielectrophoresis; enrichment; stepping<br />
electric field.<br />
I. INTRODUCTION<br />
Biological manipulation is essential to numerous<br />
biomedical applications, such as: the isolation and detection<br />
of rare cancer cells, concentration of cells from dilute<br />
suspensions, separation of cells according to specific<br />
properties, and trapping or positioning of individual cells for<br />
characterization. Among these applications, concentrating<br />
rare cells, such as circulating tumor cells (CTC), circulating<br />
fetal cells, and stem cells, has been an important technique in<br />
biological and clinical studies [1,2]. A highly sensitive and<br />
specific identification of CTC could prove helpful in the<br />
early diagnosis of invasive cancers [3]. The methods of CTC<br />
detection are generally divided into cytometric- and<br />
nucleic-acid-based techniques; however, both of these<br />
techniques require an enrichment and detection procedure<br />
[1,4]. Numerous methods for concentrating biological cells<br />
have been addressed in the relevant literature [5], such as<br />
immuno-affinity, filtration (ISET, Isolation by Size of<br />
Epithelial Tumor cells), fluorescent- (FACS,<br />
fluorescence-activated cell sorting) and magnetic-activated<br />
cell sorting (MACS, magnetic activated cell sorting), cell<br />
surface markers, optical tweezers, and dielectrophoresis.<br />
Dielectrophoresis (DEP) is achieved under a non-uniform<br />
Chia Yi, Taiwan, R.O.C.<br />
electric field generated by various electrode patterns.<br />
Previous studies on dielectrophoretic response adopted large<br />
electrodes, such as needles, pins, wires and sheets [6, 7].<br />
Microfabrication technology has been employed to create the<br />
microelectrode patterns in the studies on electrophoresis;<br />
thereby, sufficiently large DEP forces were generated to<br />
manipulate particles with small applied voltages. The<br />
different patterns of microelectrodes used for DEP have been<br />
reviewed in the relevant literature [8, 9]. The contactless and<br />
gentle forces on cells are produced by dielectrophoresis;<br />
therefore, it is particularly suitable for cell manipulation in a<br />
microchip [9]. The main aim of this study was to design a<br />
handheld device providing the stepping electric fields and a<br />
dielectrophoretic microchip with circular microelectrode for<br />
cellular preconcentration. Moreover, the preliminary<br />
experiment also aimed to demonstrate the feasibility of<br />
enriching cells with the proposed device.<br />
II. THEORY AND DESIGN<br />
The DEP force (F DEP ) acting on a spherical particle of<br />
radius R suspended in a fluid of permittivity ε , is given as:<br />
m<br />
3<br />
2<br />
DEP<br />
= 2 επ<br />
m<br />
Re(<br />
CM<br />
) ∇EfRF<br />
(1)<br />
where Re( f CM<br />
) is the real part of the Clausius-Mossotti<br />
factor; the magnitude of the electric field, E, may be replaced<br />
by E rms , which is the root-mean-square of the external field,<br />
in an alternating field. The Clausius-Mossotti factor (f CM ) is<br />
a parameter of the effective polarizability of the particle; it<br />
varies as a function of the frequency of the applied field (ω)<br />
and the dielectric properties of the particle and the<br />
surrounding medium. The Clausius-Mossotti factor for a<br />
spherical particle is represented as:<br />
**<br />
⎡ − εε ⎤<br />
mp<br />
f = ⎢ * * ⎥<br />
(2)<br />
CM<br />
⎣ p<br />
+ 2εε<br />
m ⎦<br />
*<br />
*<br />
where ε and<br />
p<br />
ε are the complex permittivity of the<br />
m<br />
particle and the medium, respectively. The complex<br />
permittivity is related to the conductivity σ and angular<br />
frequency ω through the formula:<br />
σ<br />
* εε j−≡<br />
ω<br />
( j 1−= ) (3)<br />
Therefore, the DEP force is dependent upon the dielectric<br />
properties of the particles and the medium solution, particle<br />
352
m<br />
m<br />
respectively, of the suspension medium. Based on the<br />
protoplast model, viable HeLa cells in a sucrose medium<br />
(ε r =78; σ=1.76×10 -3 S/m) exhibit a strongly positive<br />
dielectrophoretic response; i.e. the Clausius-Mossotti factor<br />
is 1.0, at high frequencies of 600 kHz [11].<br />
The circular microelectrodes were designed, and the<br />
operational concept of the cellular enrichment is illustrated in<br />
Fig. 1. When the electric field is applied to two adjacent<br />
microelectrodes, a high-electric-field region is generated<br />
between the electrode pair. The applied electric field is<br />
subsequently switched to the electrode pair next to the<br />
previous pair by relays from the peripheral to the center pair<br />
of microelectrodes; therefore, the stepping electric field is<br />
generated. The positive dielectrophoretic cells are conveyed<br />
along the direction of the stepping electric field due to the<br />
movement of the high-electric-field region to the center of<br />
the circular electrodes. The area of the high-electric-field<br />
region between adjacent electrodes is decreased gradually<br />
towards the center. As a result, the cells are not only<br />
conveyed but also concentrated on the central<br />
microelectrodes.<br />
III.<br />
EXPERIMENTAL SECTION<br />
A. Chip Fabrication<br />
A biocompatible material of polydimethylsiloxane<br />
(PDMS) was adopted as a microchamber in the microchip for<br />
cellular preconcentrating. The cellular microchip was<br />
designed, and a schematic illustration of the device is shown<br />
in Fig. 2a. The radius of the microchamber made of PDMS<br />
was 600 μm and the cells were introduced into the chamber<br />
using a pipette. Both the width and space of the electrodes<br />
were 30 μm, and the radius of curvature for the circular<br />
microelectrodes was 495 μm, as shown in Fig. 2b. The<br />
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<br />
size, and frequency of the applied electric field. It can be<br />
either a positive DEP, which pulls the particle toward the<br />
location of a high-electric field, or a negative DEP, which<br />
repels the particle away from the high-electric-field region.<br />
Dielectric properties of viable mammalian cells can be<br />
formulated by the protoplast model, which is based on a<br />
spherical particle consisting of a cytoplasm and a lossless<br />
capacitive membrane [10, 11]. The effective permittivity<br />
device was fabricated in-house, using standard soft<br />
lithography techniques. The cells were introduced into the<br />
chamber using a pipette. The electrodes were made by<br />
depositing chrome (20 nm) and gold (100 nm) sequentially<br />
on the slides, which were cleaned in a piranha solution (a<br />
mixture of sulfuric acid and hydrogen peroxide). The PDMS<br />
prepolymer mixture (Sylgard-184 Silicone Elastomer Kit,<br />
Dow Corning, Midland, MI, USA) was diluted with hexane<br />
could be derived by neglecting the conductance of the with a 1:5 (PDMS to hexane) weight ratio. The<br />
membrane in the protoplast model; thus, the hexane-diluted PDMS prepolymer, 3 μm in thickness, was<br />
Clausius-Mossotti factor for viable cells can be rewritten as: spin-coated onto the electrodes to avoid electrolysis and<br />
*2 *<br />
*<br />
ω ( τ τ τ τ jω()1)<br />
τ τ τ −−−+−<br />
mccm<br />
cmm<br />
(4) adherence of cells on the electrodes. The mold master was<br />
f ω)( −=<br />
CM<br />
2 * *<br />
*<br />
*<br />
2(<br />
mccm<br />
j<br />
m<br />
τττ<br />
cm<br />
−++−+<br />
2)2()<br />
fabricated by ωτ spinning SU8-50 (MicroChem ττ Corp. Newton, τω<br />
where mc<br />
Rcστ<br />
/ τ = ε / σ are the time MA, USA) onto the silicon wafer (around 100 μm in height)<br />
ccc to define the microchamber. The undiluted PDMS<br />
constants, while σ c<br />
and ε<br />
c<br />
are the electrical conductivity prepolymer mixture was poured and cured on the mold<br />
and permittivity of the cytoplasm, respectively. The master to replicate the microchamber. After the PDMS<br />
parameters of c<br />
m<br />
and R represent the effective capacitance<br />
replica had been peeled off, the inlet and outlet ports were<br />
made by a puncher, and the replica was bonded with the glass<br />
of the membrane and the radius of the cell, respectively. substrate after treatment of the oxygen plasma in the O 2<br />
*<br />
Moreover, the constants of τ m<br />
and τ<br />
m<br />
can be defined as plasma cleaner (Model PDC-32G, Harrick Plasma Corp.<br />
= /σετ<br />
mmm<br />
and Ithaca, NY, USA). The fabricated chip for the cellular<br />
= Rcστ<br />
/<br />
mmm<br />
, respectively, where<br />
σ and ε are the electrical conductivity and permittivity,<br />
preconcentration is shown in Fig. 3a. The image of the<br />
microelectrodes and microchamber taken by the optical<br />
microscope is shown in Fig. 3b.<br />
B. Apparatus<br />
Two 12-voltages DC (direct current) sources converted<br />
from the ordinary house current (110 VAC) by a transformer<br />
were used as the power supply for the handheld device. An<br />
IC MAX 038 (MAXIM, USA) voltage-frequency converter<br />
and an AD817 amplifier (Analog Devices, USA) were<br />
employed to design an AC signal source to apply the electric<br />
fields required for the dielectrophoretic enrichment in the<br />
microchamber. Moreover, an 8-bit microcontroller with a 4k<br />
byte flash (AT89C51, Atmel) was adopted to control the<br />
eight relays providing the stepping electric fields. The circuit<br />
module was made on a printed circuit board (PCB). The<br />
microchip was mounted on the handheld module, which<br />
generated the stepping electric field for the cellular<br />
preconcentration, as shown in Fig. 3c. The cells were<br />
observed and recorded by an inverted fluorescence<br />
microscope (Model CKX41, Olympus, Tokyo, Japan), a<br />
mounted CCD camera (DP71, Olympus, Tokyo, Japan), and<br />
a computer with Olympus DP controller image software.<br />
C. Cell Treatment<br />
A human cervical carcinoma cell line (HeLa cells) was<br />
cultured for an experimental demonstration of cellular<br />
preconcentration by the handheld microdevice proposed<br />
herein. The cells were serially passaged as monolayer<br />
cultures in DMEM Medium (Gibco, Grand Island, NY, US),<br />
with 3.7 g of NaHCO 3 per liter of medium added,<br />
supplemented with 10% fetal bovine serum (FBS, Gibco,<br />
Grand Island, NY, US), and 1% penicillin/streptomycin<br />
(Gibco, Grand Island, NY, US). The cell culture dish<br />
(Falcon, Franklin Lakes, NJ, US) was incubated in a<br />
humidified atmosphere containing 5% carbon dioxide at<br />
37°C, and the medium was replaced every 1 to 2 days. Cells<br />
353
grown to sub-confluence were washed with<br />
phosphate-buffered saline (PBS, Biochrome, pH 7.4) and<br />
harvested by a 5-min treatment with 0.25% Trypsin and<br />
0.02% EDTA (Sigma, US). The cells for the experiments<br />
were then suspended in a sucrose solution with an 8.62 wt%<br />
and a measured conductivity of 1.76×10 -3 S/m. For the<br />
dielectrophoresis of cells, the sucrose solution was employed<br />
to raise the osmolarity to the normal physiological level.<br />
IV. RESULTS AND DISCUSSION<br />
Numerical simulations of the electric field were performed<br />
using commercial software CFDRC-ACE + (ESI Group,<br />
France). The effect of the presence of particles on the<br />
electric field was not considered in the simulation for the<br />
sake of simplification. After applying 16 volts from the outer<br />
microelectrode pair to the central pair of microelectrodes<br />
subsequently, the simulated square of the electric field (E 2 )<br />
was revealed, as shown in Fig. 4. The numerical results<br />
indicate that the high-electric-field region moved to where<br />
the electric field was applied, and its area decreased<br />
according to the pattern of the microelectrodes. Therefore,<br />
HeLa cells with positive dielectrophoretic response were<br />
conveyed in the direction of the stepping electric field, which<br />
is toward the center of the microchamber and concentrated at<br />
the central pair of electrodes, demonstrating the realization of<br />
the concept of the present design. The preliminary<br />
experiment for cellular enrichment was investigated herein,<br />
as depicted in Fig. 5. The HeLa cells, with a concentration of<br />
5×10 5 cells/mL, were introduced into the microchamber<br />
using a pipette. About 50 cells were in the microchamber.<br />
The peak-to-peak voltage of 16 volts and a frequency of 600<br />
kHz were applied to the electrodes. The electric field at the<br />
peripheral pair of electrodes was turned on to aggregate the<br />
cells. The electric field was held for about 20 seconds and<br />
then switched to the next adjacent pair of electrodes. The<br />
duration of cellular concentration from the outermost to the<br />
central pair of microelectrodes was about 160 seconds. The<br />
experimental results indicate that the enrichment of HeLa<br />
cells were successfully exhibited at the center pair of<br />
microelectrodes. Furthermore, the experimental results<br />
indicate that the HeLa cells were successfully concentrated at<br />
the center pair of microelectrodes with an approximate<br />
efficiency of 75%.<br />
V. CONCLUSIONS<br />
A handheld device providing the stepping electric fields<br />
and a dielectrophoretic microchip with circular<br />
microelectrode for cellular enrichment was proposed and<br />
demonstrated in the present study. The positive<br />
dielectrophoretic cells were conveyed due to the movement<br />
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<br />
of the high-electric-field region. Furthermore, the cells were<br />
concentrated at the center pair of microelectrodes due to the<br />
fact that the area of the high-electric-field between the<br />
adjacent electrodes was gradually decreased from the<br />
peripheral to the center pair of microelectrodes. The<br />
preliminary experiment for cellular enrichment also<br />
indicated that the HeLa cells were successfully concentrated<br />
at the center pair of microelectrodes with an approximate<br />
efficiency of 75% when the time interval of relay switching<br />
was set at 20 s. The operational method of the cellular<br />
enrichment herein could enhance the sensitivity of further<br />
CTC detections. Besides, the handheld device presented in<br />
this work can be applied in point-of-care applications.<br />
ACKNOWLEDGMENT<br />
The authors would like to thank the National Science<br />
Council of the Republic of China for its financial support of<br />
this research under contract Nos.<br />
NSC-99-2923-E-194-001-MY3<br />
and<br />
NSC-99-2221-E-194-014. In addition, the National Center<br />
for High-Performance Computing for the use of computer<br />
time and its facilities is also acknowledged.<br />
REFERENCES<br />
[1] Mostert B., Sleijfer S., Foekens J.A., Gratama J.W.: “Circulating<br />
tumor cells (CTCs): Detection methods and their clinical relevance<br />
in breast cancer,” Cancer Treat. Rev., 2009, 35, pp. 463-474<br />
[2] Cheng X., Gupta A., Chen C., Tompkins R.G., Rodriguez W. and<br />
Toner M.: “Enhancing the performance of a point-of-care CD4+<br />
T-cell counting microchip through monocyte depletion for<br />
HIV/AIDS diagnostics,” Lab Chip, 2009, 9, pp. 1357-1364<br />
[3] Paterlini-Brechot P.and Benali N.L.: “Circulating tumor cells (CTC)<br />
detection: Clinical impact and future directions,” Cancer Lett., 2007,<br />
253, pp. 180-204.<br />
[4] Grodzinski P., Yang J., Liu R.H. and Ward M.D.: “A modular<br />
microfluidic system for cell pre-concentration and genetic sample<br />
preparation,” Biomed. Microdevices, 2003, 5, pp. 303-310<br />
[5] Dharmasiri U., Witek M.A., Adams A.A., and Soper S. A.:<br />
“Microsystems for the capture of low-abundance cells,” Ann. Rev.<br />
Anal. Chem., 2010, 3, pp. 409-431.<br />
[6] Pohl H.A.: Dielectrophoresis, Cambridge University Press, New<br />
York, 1978.<br />
[7] Jones T.B. and Kraybill J.P.: “Active feedback-controlled<br />
dielectrophoretic levitation,” J. Appl. Phys., 1986, 60, pp.<br />
1247-1252.<br />
[8] Ramos A., Morgan H., Green G.N. and Castellanos A.: “AC<br />
Electrokinetics: A review of forces in microelectrode structure,” J.<br />
Phys. D: Appl. Phys., 1998, 31, pp. 2338-2353.<br />
[9] Voldman J.: “Electrical forces for microscale cell manipulation,”<br />
Annu. Rev. Biomed. Eng., 2006, 8, pp. 425-454.<br />
[10] Jones T.B.: Electromechanics of particles, Cambridge University<br />
Press, New York, 1995.<br />
[11] Jen C.P. and Chen T.W.: “Selective trapping of live and dead<br />
mammalian cells using insulator-based dielectrophoresis within<br />
open-top microstructures,” Biomed. Microdevices, 2009, 11, pp.<br />
597-607.<br />
354
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<br />
Figure 1: Operation concept of the cellular enrichment by dielectrophoresis in<br />
the stepping electric field.<br />
Figure 4: Simulated results of square of the electric field (E 2 ) while the electric<br />
field was applied from the peripheral to the center pair of microelectrodes. The<br />
voltage applied to the electrode pair is 16 volts.<br />
(a)<br />
(b)<br />
Figure 2: (a) Schematic diagram of the cellular microchip using<br />
dielectrophoresis and (b) the dimensions of the circular microelectrodes.<br />
(a)<br />
(b)<br />
Figure 5: Experimental results of enrichment for HeLa cells. The peak-to-peak<br />
voltage and frequency applied was 16 volts and 600 kHz, respectively. The<br />
time interval of relay switching was 20 seconds. The duration from (a) to (i)<br />
was about 160s.<br />
(c)<br />
Figure 3: The images of: (a) the fabricated chip for cellular enrichment, (b)<br />
circular microelectrodes and (c) the complete handheld microchip with the<br />
electric module providing the stepping electric field.<br />
ISBN:978-2-35500-013-3<br />
<br />
355
11-13 <br />
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<br />
A Novel SU-8 Microgripper with External Actuator<br />
for Biological Cells Manipulation<br />
M. Mehdi S. Mousavi 1, 2 , Giorgio De Pasquale 1 , Aurelio Somà 1 , Eugenio Brusa 1<br />
1 Department of Mechanics, Politecnico di Torino, Corso Duca degli Abruzzi 24, 10129, Torino, Italy<br />
2 Italian Institute of Technology, Center for Space Human Robotics, Corso Trento 21, 10129, Torino, Italy<br />
mehdi.mousavi@polito.it, giorgio.depasquale@polito.it, aurelio.soma@polito.it, eugenio.brusa@polito.it<br />
Abstract- Specification and target impose several limitations<br />
and difficulties in micro manipulators design. These obstacles<br />
are even more important when the target of microgripping is<br />
biological cells. Even though a variety of designs and solutions<br />
has been proposed in the literature, the problems of<br />
temperature at microgripper tip and applied voltage in<br />
gripping jaws are still unsolved. In this paper a new approach<br />
to eliminate this kind of problems in biological cells<br />
manipulation is introduced. The proposed kinematics is<br />
externally actuated and an optimization procedure based on<br />
FEM simulations has been performed to improve the micro<br />
gripper design. Some considerations on fabrication process<br />
show that the new approach can sensitively decrease the cost<br />
and time of fabrication processes, as well as the complexity of<br />
the technologies involved.<br />
I. INTRODUCTION<br />
Common micro grippers cannot be used to manipulate<br />
biological samples, such as living cells, because of their<br />
actuation methods. The actuation mechanism shall be<br />
suitable for operating in electrolytic aqueous media because<br />
of ionic environment of cells [1, 2]. This prerequisite limits<br />
the application of high voltage to actuator that is necessary<br />
in piezo-actuated grippers since bubble formation, caused<br />
by electrolysis, occurs at 1.5–2 volts in water [3]. Moreover,<br />
any exposure to magnetic or electrical fields may have some<br />
negative effects on biological cells. This also limits the<br />
application of electrostatic or electromagnetic actuated<br />
micro grippers. In addition, shape memory alloy (SMA)<br />
actuators are not a good candidate for micro grippers due to<br />
lack of reliability for a reasonable number of operations.<br />
Furthermore, the maximum allowed temperature for<br />
manipulation of human cells in many applications such as<br />
Intracytoplasmic Injection or Pro-nuclei DNA injection is<br />
around 37°C that is quite lower than the required high<br />
temperature (more than 100° C for bare extended arms of<br />
gripper) in an electro-thermal gripper. Therefore, even<br />
though electro-thermal actuators are of great interests<br />
among researchers for cell manipulation, they show many<br />
difficulties when are used [4-6]. Another point is biocompatibility<br />
of gripper materials that places some<br />
restrictions in choosing of actuation method and fabrication<br />
process.<br />
From above explanation, it is clear that whatever the<br />
actuation method is, there are many points that must be<br />
taken into consideration for microgripper design. This work<br />
proposes a survey of the literature about the kinematic and<br />
actuation solutions adopted for microgrippers; then a new<br />
design approach is proposed and the FEM simulation of few<br />
candidate devices for cells manipulation are reported.<br />
II. DESIGN ISSUES IN MICROGRIPPING<br />
A survey of the literature reveals the following key<br />
features in design of microgrippers for biological cells:<br />
- actuation principle<br />
- kinematics<br />
- fingertips shape<br />
- force feedback<br />
- releasing strategy<br />
The actuation strategy is usually determined by selecting<br />
internal or external actuators. About the first category, it is<br />
possible to build some specific parts of the gripper with<br />
piezoelectric (PZT) material to generate a localized force<br />
when an electric voltage is provided [7]. The electrostatic<br />
force can be used as an actuation by applying a voltage<br />
difference on a capacitor with movable armature [8]. The<br />
thermal actuation, widely used for both biological and nonbiological<br />
manipulation, is based on the thermal expansion<br />
of the gripper arms due to the Joule effect in presence of<br />
electric currents [9]. A faster response of the arms can be<br />
achieved with shape memory alloys (SMA) [10]: they are<br />
able to restore almost immediately the memorized shape<br />
when a threshold temperature is passed. The<br />
electromagnetic actuation is based on micro-coils and is<br />
able to generate weak confined magnetic fields [11].<br />
Hydraulic and pneumatic actuation can be used to<br />
manipulate bio-cells with micro-pipes integrated in small<br />
circuits including micro-pumps and valves [12]. There are<br />
strong limitations in using internal actuators for the<br />
manipulation of biological particles. PZT actuators have<br />
strong nonlinear output, high supply voltage, small motion<br />
range and other problems such as creep, mechanical fatigue,<br />
hysteresis and biocompatibility; as a consequence, they<br />
require an embedded force feedback control. The<br />
electrostatic actuators are generally disadvantaged by the<br />
small dimensions of the capacitors. To increase to the force,<br />
very complicated shapes of the gripper are necessary by<br />
introducing many comb drives; then, the motion range is<br />
strongly reduced by the small gaps between the armatures<br />
and the applied voltage easily causes electrolysis of watered<br />
356
environments. Despite thermal actuation is one of the most<br />
common way of bio-manipulation, thermal actuators may<br />
induce high temperature in the region close to the cells; few<br />
solutions were proposed with long grippers used to dissipate<br />
the heat produced by the actuators. About SMA materials,<br />
the main problem is related to their low fatigue resistance<br />
that causes very limited cycle time; then, SMA materials<br />
have small strain capability, strong nonlinearity and<br />
hysteresis and their fabrication process is usually very<br />
complicated in the microscale. The limitations of electromagnetic<br />
actuators are related to their small dimension that<br />
implies fast heating of the coil due to the Joule effect and<br />
low allowable currents; the resulting magnetic field is<br />
generally weak and subjected to relevant leakages, giving<br />
small power per unit volume. Hydraulic and pneumatic<br />
actuators are limited to pipe based devices; usually they are<br />
not suitable for precision operations involving more than<br />
one cell and the hydraulic solution only works in wet<br />
environments [13].<br />
More promising opportunities for bio-cells manipulators<br />
are offered by external actuators, which preserve the<br />
thermal insulation of the gripper and avoid contaminations<br />
or biocompatibility problems. The most suitable solutions<br />
are electric motors (DC motors and stepper motors) and<br />
piezoelectric motors [13]. The first category is affected by<br />
undesired heat generation, relatively low motion precision<br />
and quite large size for micro manipulation; furthermore,<br />
stepper motors are not able to provide smooth motion. The<br />
piezoelectric motors are the most promising solution for this<br />
application, due to their small size and high accuracy [14];<br />
they also have very high response and wide speed range<br />
(since few micrometers/second to few millimeters/second).<br />
The thermal heating is also negligible. The only problems<br />
are related to the interface between the motor and the<br />
microgripper, where interferences and frictions must be<br />
considered. Other less investigated strategies for the internal<br />
and external actuation includes ultrasonic motors,<br />
picomotors, stick-slip and inchworm actuators. The<br />
Internal actuation<br />
External actuation<br />
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<br />
advantages and limitations of fundamental internal and<br />
external actuation strategies are summarized in Table 1.<br />
The kinematic solutions adopted in the literature are very<br />
numerous and strongly related to the field of application of<br />
the gripper and to the actuation strategy used; however,<br />
most of them can be summarized in few kinematic schemes,<br />
which are reported in Table 2. Compliant structures (some<br />
examples are reported in [13, 15]) belongs to a particular<br />
class of grippers where the material elasticity is used to<br />
transfer the force from the actuator to the cell, possibly by<br />
amplifying its effect with the geometry characteristics and<br />
the structure deforming shape. This solution is very<br />
interesting because of its simplicity, and the possibility to<br />
increase the gripping surface and limiting the local pressure<br />
on the cell membrane.<br />
Normally, the fingertip shape must be carefully designed<br />
according to the conformation of the grip site. The less<br />
elaborated kinematics of arms lead to gripping by only two<br />
points; this solution is not indicated for the manipulation of<br />
bio-cells because very high pressures may interest the<br />
gripping site and compromise the cell integrity. Thus it is<br />
preferable to choose rigid translating fingers instead of<br />
rotating fingers (see the example in the Table 2) and to<br />
provide a proper shape to the clamps. The most diffused<br />
fingertips shapes can be divided in flat, angular carved,<br />
cylindrical carved and non-standard fingertips, as<br />
represented in Table 3 [16]. Another advantage of<br />
compliant structures is the possibility to slightly adapt the<br />
shape of the fingertip to the conformation of the cell: thanks<br />
to the high deformability of the structure, it is possible to<br />
embrace the cell and to distribute almost uniformly the<br />
gripping force on its surface.<br />
The force feedback measurement is often crucial for biomanipulation,<br />
due to the small mechanical resistance of the<br />
cells [7,.17-19]. The most suitable strategy is the<br />
displacement control through optical detection, which is<br />
contactless and very accurate. Other methods were<br />
explored, for instance by using integrated<br />
TABLE 1<br />
ADVANTAGES AND LIMITATIONS OF FUNDAMENTAL INTERNAL AND EXTERNAL ACTUATION STRATEGIES<br />
Actuation strategy Advantages Limitations<br />
Piezoelectric Thermal stability, high accuracy, high response.<br />
Nonlinearity, high supply voltage, small motion<br />
range, creep, fatigue, hysteresis, low biocompatibility.<br />
Electrostatic<br />
capacitive<br />
Consolidated manufacturing process, direct motion feedback.<br />
Complicated geometry, small motion range,<br />
electrolysis and bubbles formation.<br />
Thermal Consolidated manufacturing process. High temperature, low response.<br />
SMA actuators Faster response then thermal actuation, large motion range.<br />
Fatigue, small motion range, nonlinearity, hysteresis,<br />
hard manufacturing process, high cost.<br />
Electromagnetic Preservation of cells integrity. Coil heating, magnetic field weakness, field leakage,<br />
Hydraulic and<br />
pneumatic<br />
Reliability, preservation of cells integrity.<br />
Limited applicability.<br />
DC motors Thermal insulation, high speed, high accuracy.<br />
Heat generation, dimensions, hysteresis, interface<br />
connection, feedback control needed.<br />
Step motors Thermal insulation, very large motion range.<br />
Heat generation, low precision, dimensions, unsmooth<br />
motion, interface connection, noise.<br />
Piezoelectric<br />
motors<br />
Large force, high accuracy, high response, thermal insulation,<br />
small size, no wear and tear, low power consumption.<br />
Interface connection.<br />
357
Clamps<br />
rotation<br />
[21]<br />
Clamps<br />
translation<br />
[21]<br />
Compliant<br />
structures<br />
[15]<br />
Thermal<br />
expanded<br />
arms<br />
Electromagnetic<br />
actuation<br />
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May 2011, Aix-en-Provence, France<br />
<br />
exploit the features of the gripper surface (shape, material,<br />
TABLE 2<br />
MOST COMMON KINEMATIC STRATEGIES<br />
coatings, etc.), while the last ones make use of external<br />
Open state<br />
Closed state<br />
actions (forces, pressures, vibrations, etc) [20]. A list of<br />
releasing strategies is reported in Table 4.<br />
TABLE 3<br />
MOST COMMON STANDARD TYPOLOGIES OF FINGERTIP SHAPES<br />
Flat Angular carved Cylindrical carved<br />
piezoresistive transducers or micro capacitive sensors;<br />
however these approaches are usually limited by the low<br />
biocompatibility of materials and the possibility to induce<br />
electrolysis of water with related bubbles formation.<br />
Due to the small effect of gravity force compared to the<br />
adhesion and capillary forces in microgripping of bio-cells,<br />
the releasing is generally problematic. Many releasing<br />
strategies were tested in the literature and they can be<br />
divided in passive and active strategies; the first ones<br />
Passive releasing<br />
Active releasing<br />
Releasing strategy<br />
Rough surfaces<br />
Hydrophobic coating<br />
Conductive coating<br />
Vacuum environment<br />
Fluid environment<br />
Ionized air<br />
Vibrations<br />
Air pressure<br />
Heating<br />
Electrostatic control<br />
Adhesion to the<br />
substrate<br />
Additional tools<br />
TABLE 4<br />
MOST COMMON PASSIVE AND ACTIVE RELEASING STRATEGIES<br />
III. DESIGN AND FEM OPTIMIZATION<br />
As mentioned before, different kind of problems in living<br />
cells manipulation are related to the health of cells and<br />
caused by the actuator part of microgrippers. Therefore, in<br />
all kind of actuators any undesirable effect of actuator on<br />
biological cells has to be avoided. In this paper a kind of<br />
gripper is considered that uses external actuator approach as<br />
its actuation method. In this approach the actuation part<br />
completely separates from gripping jaws. It seems that it is<br />
the first time that external actuation method is proposed for<br />
biological cells manipulation in order to solve all kind of<br />
problems that are related to the actuation mechanism of<br />
microgrippers. Furthermore, the fabrication process of the<br />
gripper can be considerably simplified by using this<br />
approach. This separation allows using different kind of<br />
actuators as driving part of the gripper without any<br />
consideration to their undesired effects on cells health.<br />
Several configurations were considered to define a shape<br />
suitable for this application. Figures 1 to 4 show the<br />
improvement procedure of microgripper design in this<br />
work. For all those layouts tip of the gripper was conceived<br />
to be proper for gripping of a cell with 35 µm diameter.<br />
Moreover, the total length of all proposed configurations<br />
and their out-of-plane thickness are 1 mm and 20 µm,<br />
respectively, and was considered fixed parameters during<br />
optimization procedure. Furthermore, the applied<br />
displacement to the moving arm of the gripper was<br />
considered 20 µm when it was pulled and 10 µm when it<br />
was pushed. In all FEM analysis of the structures we<br />
changed the geometrical parameters so that the maximum<br />
stress did not exceed of 34 MPa. Four dimensional models<br />
of the proposed layouts were developed and then<br />
investigated by the finite element code ANSYS.<br />
Description<br />
The contact area is reduced, as the electrostatic adhesion force.<br />
It reduces the surface tension effect.<br />
The electrostatic forces are reduced by conductive coatings or materials with small potential difference<br />
with the object.<br />
It reduced the surface tension effect.<br />
It eliminates the surface tension effect and reduces electrostatic forces.<br />
It reduces electrostatic forces.<br />
The acceleration imposed causes the object releasing due to inertial force.<br />
A pressurized air flow overcomes the adhesion forces.<br />
The temperature increasing reduces the capillary forces.<br />
The electrostatic force is controlled by shorting the gripper electrodes or inverting the polarity.<br />
The object adheres to the substrate due to higher adhesion forces, or gluing on the substrate, or engagement<br />
by the substrate.<br />
Additional tools are used to detach the object.<br />
358
The FEM model allowed calculating the structural<br />
stiffness, as well as the accuracy of kinematics. An<br />
optimization work was done by changing the configuration,<br />
length, and width of the different parts of the grippers. The<br />
results of all analysis can be seen in Table 5. To select the<br />
material, bio-compatibility of the gripper was taken into<br />
account. After performing a comprehensive survey in the<br />
literature, the SU-8 polymeric material was selected because<br />
of its properties, which are particularly suitable for cell<br />
manipulation devices. For the FEM simulations, the<br />
following properties of SU-8 were considered: coefficient of<br />
thermal expansion α = 52x10 -6 K -1 , Young’s modulus E =<br />
4.02 GPa, Poisson’s ratio ν = 0.22 and ultimate tensile stress<br />
of 34 MPa.<br />
Since in all microgrippers previously proposed in the<br />
literature, due to angular motion of the jaws, the gripping<br />
planes at the tip of the tweezers do not remain parallel<br />
during cell manipulation (see Table 2), in this study an<br />
attempt was performed to change this strategy, by leading to<br />
surround a cell and keep it within the room defined by the<br />
gripper tips, in closed configuration. Furthermore, by<br />
increasing the contact area of the cell and gripper tip, this<br />
approach causes less stress on the cells membrane during<br />
gripping. Above all, a single mask layer is enough to<br />
fabricate this layout and there is no need to dope any kind of<br />
metal on substrates and use several mask layers in<br />
fabrication process that is usual in all thermal actuators.<br />
The first configuration in Table 1 is our first idea to<br />
achieve the above mentioned goals. One upper and one<br />
lower arm which are connected together with some inclined<br />
connecting parts compose this simple configuration. The<br />
upper arm is fixed and the lower arm is attached to an<br />
external actuator. In this configuration the most important<br />
problem was the y-direction movement of the upper arm of<br />
the gripper due to its bending. Figure 1 shows the deformed<br />
and undeformed shape and also stress distribution of this<br />
layout. The results of all analyses can be seen in Table 5. By<br />
applying 20 µm displacement to the lower arm in FEM<br />
simulation, more than 100 µm displacement in y direction<br />
occurred. To solve this problem a wider upper arm and<br />
longer connecting parts were used in second proposed<br />
configuration. Moreover, the inclined connectors were<br />
changed to vertical position in order to increase the<br />
movement in direction of x in comparison to y and to<br />
decrease the amount of force required by the lower arm. To<br />
keep the gripper tip dimensions proper for the considered<br />
cell size (35 µm) the shape of the tip was changed<br />
accordingly (Fig. 2).<br />
As can be seen in Table 5 these modifications did not<br />
solve the problem. Still were found about 20 µm of<br />
displacement along y direction when displacement was<br />
equal along direction x. Even though the second<br />
configuration decreased the movement along y direction in<br />
comparison to the first layout, long connecting parts<br />
increased the dimension of whole gripper structure. The<br />
next idea to solve these difficulties is shown in Fig. 3. An<br />
additional arm was joined to the structure. Displacement in<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
FEM simulation was induced on the upper arm while the<br />
lower two arms were fixed. This improvement allowed<br />
finding good results in terms of stress and opening range of<br />
the gripper tip. As it can be seen in Fig. 3 length of the<br />
upper arm can be a problem for the microfabrication,<br />
because of its compliance. This effect can be seen in Fig. 3,<br />
since a significant deflection of the upper arm occurred after<br />
applying displacement. To solve this final problem the<br />
upper arm was supported by another structure to prevent its<br />
deflection. Figure 4 shows the final configuration. In this<br />
layout a large opening at the gripper tips occurs if only 10<br />
µm of displacement are applied (half of the first two<br />
layouts). As can be seen in Table 5, the most opening range<br />
with minimum stress can be achieved by this layout.<br />
Furthermore, the least stiffness is related to this<br />
configuration while the width of the gripper does not exceed<br />
140 µm.<br />
Fig. 1. First configuration, left: deformed and undeformed shape,<br />
right: stress analysis<br />
Fig. 2. Second configuration, left: deformed and undeformed shape,<br />
right: stress analysis<br />
Fig. 3. Third configuration, left: deformed and undeformed shape,<br />
right: stress analysis<br />
Fig. 4. Fourth configuration, left: deformed and undeformed shape,<br />
right: stress analysis<br />
359
Layout<br />
Total dimensions<br />
of the gripper<br />
(length, width,<br />
thickness)<br />
(µm)<br />
Minimum<br />
width<br />
(µm)<br />
Maximum<br />
width<br />
(µm)<br />
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<br />
TABLE 5<br />
RESULTS OF FEM ANALYSIS FOR EACH LAYOUT<br />
Applied<br />
displacement<br />
to the moving<br />
arm a<br />
(µm)<br />
Equivalent<br />
force<br />
(µN)<br />
Stiffness<br />
(µN/µm)<br />
Tip<br />
displacement in<br />
direction of y<br />
(fixed arm,<br />
moving arm)<br />
Ux,Relative<br />
displacement<br />
of two tips in<br />
direction x b<br />
(µm)<br />
Uy, Relative<br />
displacement<br />
of two tips in<br />
direction y b<br />
(µm)<br />
Maximum<br />
stress<br />
(Mpa)<br />
1 1000 x 127 x 20 18 40 -20 28 3.1 (106, 102) 10 -4 33.93<br />
2 1000 x 390 x 20 20 100 -20 25 2.6 (20, 19) 10 -1 32.94<br />
3 1000 x 70 x 20 5 30 10 5.5 0.14 (4, 42) -0.5 -38 30.95<br />
4 1000 x 140 x 20 5 30 10 4.2 0.11 (5, 41) -6 -35 32.55<br />
a Positive values mean tension and minus values mean compression<br />
b Minus values mean the two arms moved away of each other<br />
Fig. 5. Approaching<br />
Fig. 6. Opening<br />
Fig. 7. Closing<br />
Fig. 8. Keeping<br />
The gripping procedure that was explained here is divided<br />
in four steps; approaching to the cell, opening of the jaws,<br />
closing of the jaws and finally, keeping the cell. These steps<br />
are shown in Fig. 5 to 8. All relative sizes are according to<br />
the real condition. In Fig. 1 a 35 µm cell has been shown in<br />
front of the gripper tip. This is after the process of<br />
approaching to the cell. Fig. 6 shows the maximum opening<br />
range of the gripper tip that can be achieved by applying 10<br />
µm or 4.2 µN force to the moving arm. As can be seen the<br />
opening range in this configuration is more than the other<br />
layouts. This also can be confirmed by the stiffness numbers<br />
of Table 5. Stiffness was calculated by dividing the amount<br />
of forces to the received relative displacement of the gripper<br />
tips. Fig. 7 is the returning process of the gripper to its<br />
unloaded condition. This figure shows the jaws when 2 µm<br />
displacement is applied to the moving arm of the gripper.<br />
Finally, Fig. 8 shows the relax condition of the gripper<br />
without any applied load or displacement. In this figure the<br />
gripper keep the cell inside its tips. One of the most<br />
advantages of this approach is increasing the contacts area<br />
between cell membrane and gripper tip that causes less<br />
stress on the membrane of the cell. Also in the case of a<br />
precise manipulation it can be completely a non-contact<br />
gripping.<br />
IV. RESULTS AND CONCLUSION<br />
A new approach in designing of microgrippers for<br />
biological cells manipulation was proposed and optimized<br />
by finite element method. By separating the actuator part of<br />
the microgripper from gripping jaws, many existing<br />
problems in cell manipulation were solved. This layout<br />
assures significant benefits such as the reduction of micro<br />
fabrication steps, time and cost of building process. By<br />
using this approach, there is no need to dope any kind of<br />
metal on substrates and use several mask layers in<br />
fabrication process that is usual in all thermal actuators.<br />
Some attempts to design an innovative configuration were<br />
performed and herein documented, by distinguishing<br />
advantages and disadvantages of each new layout proposed.<br />
The main innovation was strategy of surrounding the cells<br />
and keeping them instead of using tweezers for gripping,<br />
which make a local contact with the cells limited to two or<br />
few points. A finite element optimization was then<br />
performed to improve the performance of the proposed<br />
360
configuration of microgrippers and their functionality. The<br />
authors, after this work, where they have proposed and<br />
simulated this structure, will concentrate the attention on the<br />
shape of the cavity for the cell in order to correlate the<br />
contact force acting on the cell and the actuation force<br />
required to drive the manipulator.<br />
ACKNOWLEDGMENT<br />
The work presented was supported by the Italian Institute<br />
of Technology (IIT) at Politecnico di Torino, Center for<br />
Space Human Robotics.<br />
REFERENCES<br />
[1] N. Chronis and L.P. Lee, “Electrothermally activated SU-8<br />
microgripper for single cell manipulation in solution,” J.<br />
Microelectromech. S., vol. 14, pp. 857–863, 2005.<br />
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micro-gripper with a force sensor for manipulating a cell,” proc. of<br />
SICE-ICASE, Busan, Korea, pp. 5833-5836, 2006.<br />
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electrolysis of water: an actuation principle for MEMS with a big<br />
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[4] W.J. Li and N. Xi, “Novel micro gripping, probing, and sensing<br />
devices for single-cell surgery,” proc. of 26th Ann. Int. Conf. of the<br />
IEEE Eng. in Medicine and Biology Soc., San Francisco, USA, vol.<br />
1, pp. 2591-2594, 2004.<br />
[5] B. Solano and D. Wood, “Design and testing of a polymeric<br />
microgripper for cell manipulation,” Microelectron. Eng., vol.<br />
84, pp. 1219-1222, 2007.<br />
[6] R.E. Mackay, H.R. Le, K. Donnelly, and R.P. Keatch, “Microgripping<br />
of small scale tissues,” proc. of 4th Europ. Conf. of the<br />
Int. Feder. for Medical and Biological Eng., vol. 22, pp. 2619-<br />
2622, 2009.<br />
[7] R. Pérez, N. Chaillet, K. Domanski, P. Janus, and P. Grabiec,<br />
“Fabrication, modeling and integration of a silicon technology<br />
force sensor in a piezoelectric micro-manipulator,” Sensor. Actuat.<br />
A - Phys., vol. 128, pp. 367-375, 2006.<br />
[8] F. Beyeler, A. Neild, S. Oberti, D.J. Bell, Y. Sun, J. Dual, and B.J.<br />
Nelson, “Monolithically fabricated microgripper with integrated<br />
force sensor for manipulating microobjects and biological cells<br />
aligned in an ultrasonic field,” J. Microelectromech. S., vol. 16, pp.<br />
7-15, 2007.<br />
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[9] N.T. Nguyen, S.S. Ho, and C.L. Low, “A polymeric microgripper<br />
with integrated thermal actuators,” J. Micromech. Microeng., vol.<br />
14, pp. 969-974, 2004.<br />
[10] M. Kohl, B. Krevet, and E. Just, “SMA microgripper system,”<br />
Sensor. Actuat. A - Phys., vol. 97-98, pp. 646-652, 2002.<br />
[11] I. Giouroudi, H. Hötzendorfer, J. Kosel, D. Andrijasevic, and W.<br />
Brenner, “Development of a microgripping system for handling of<br />
microcomponents,” Precis. Eng., vol. 32, pp. 148-152, 2008.<br />
[12] V. Seidemann, S. Büterfisch, and S. Büttgenbach, “Fabrication and<br />
investigation of in-plane compliant SU8 structures for MEMS and<br />
their application to micro valves and micro grippers,” Sensor.<br />
Actuat. A - Phys., vol. 97-98, pp. 457-461, 2002.<br />
[13] P.R. Ouyang, R.C. Tjiptoprodjo, W.J. Zhang, and G.S. Yang,<br />
“Micro-motion devices technology: the state of arts review,” Int. J.<br />
Adv. Manuf. Technol., vol. 38, pp. 463-478, 2008.<br />
[14] S.K. Nah and Z.W. Zhong, “A microgripper using piezoelectric<br />
actuation for micro-object manipulation,” Sensor. Actuat. A -<br />
Phys., vol. 133, pp. 218-224, 2007.<br />
[15] J.A. Martinez and R.R. Panepucci, “Design, fabrication, and<br />
characterization of a microgripper device,” proc. of FCRAR,<br />
Tampa, USA, pp. 1-6, 2007.<br />
[16] P. Pedrazzoli, R. Rinaldi, and C.R. Boër, “A rule based approach<br />
to the gripper selection issue for the assembly process,” proc. of<br />
4th IEEE Int. Symp. on Ass. and Task Plan., Fukuoka, Japan, pp.<br />
202-207, 2001.<br />
[17] S. Fahlbusch and S. Fatikov, “Implementation of self-sensing SPM<br />
cantilevers for nano-force measurement in microrobotics,”<br />
Ultramicroscopy, vol. 86, pp. 181-190, 2001.<br />
[18] X. Liu, K. Kim, Y. Zhang, and Y. Sun, “Nanonewton force sensing<br />
and control in microrobotic cell manipulation,” Int. J. Robot Res.,<br />
vol. 28, pp. 1065-1076, 2009.<br />
[19] M.C. Carrozza, A. Eisinberg, A. Menciassi, D. Campolo, S.<br />
Micera, and P. Dario, “Towards a force-controlled microgripper<br />
for assembling biomedical microdevices,” J. Micromech.<br />
Microeng., vol. 10, pp. 271-276, 2000.<br />
[20] G. Fantoni and M. Porta, “A critical review of releasing strategies<br />
in microparts handling,” in Micro-Assembly Technologies and<br />
Applications, vol. 260, S. Ratchev and S. Koelemeijer, Eds.<br />
Boston: Springer, 2008, pp. 223-234.<br />
[21] B. Hoxhold and S. Büttgenbach, “Easily manageable,<br />
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<br />
Particle Focusing in a Contactless Dielectrophoretic<br />
Microfluidic Chip with Insulating Structures<br />
Chun-Ping Jen 1 *, Hsin-Yuan Shih 2 , Yung-Chun Lee 3 and Fei-Bin Hsiao 2<br />
1* Department of Mechanical Engineering, National Chung Cheng University, Chia Yi, Taiwan, R.O.C.<br />
2 Institute of Aeronautics and Astronautics, National Cheng Kung University, Tainan, Taiwan, R.O.C.<br />
3 Department of Mechanical Engineering, National Cheng Kung University, Tainan, Taiwan, R.O.C.<br />
Abstract- The main purpose of the present study was to investigate<br />
the feasibility of applying the technique of contactless<br />
dielectrophoresis (cDEP) on an insulator-based dielectrophoretic<br />
(iDEP) microdevice with effective focusing of particles. The<br />
particles were introduced into the microchannel and pre-confined<br />
hydrodynamically by the funnel-shaped insulating structures<br />
close to the inlet. The particles were, therefore, repelled toward<br />
the center of the microchannel by the negative dielectrophoretic<br />
forces generated by the insulating structures. The microchip was<br />
fabricated by the technique of cDEP. The electric field in the main<br />
microchannel was generated by using electrodes inserted into two<br />
conductive micro-reservoirs, which were separated from the main<br />
microchannel by thin insulating barriers made of 20 μm-width of<br />
PDMS. The impedance of the PDMS barrier under different<br />
frequencies was measured by an impedance analyzer and the<br />
fitting curve to experimental data using the least-squares method<br />
were also addressed. The results revealed the capacitive behavior<br />
of the PDMS, in which the impedance decreased with the<br />
frequency. The numerical simulations indicated that an increase<br />
in the strength of the applied electric field significantly enhanced<br />
the performance of focusing. The preliminary experiments<br />
employing latex particles with 10 μm in diameter were conducted<br />
to demonstrate the feasibility of the present design. The usage of<br />
contactless DEP technique makes the insulator-based<br />
dielectrophoretic microchip mechanically robust and chemically<br />
inert. Furthermore, the voltage applied was also reduced rather<br />
than conventional iDEP microchip.<br />
Keywords: contactless dielectrophoresis; insulating structures;<br />
microfluidic; focusing.<br />
I. INTRODUCTION<br />
Dielectrophoresis (DEP) has been widely used for the<br />
manipulation of particles in various microdevices.<br />
Contact-free and gentle forces on cells are produced;<br />
therefore, conditions are particularly suitable for cell<br />
manipulation in a microchip. Dielectrophoresis is achieved<br />
under a non-uniform electric field generated by various<br />
electrode patterns. Furthermore, the use of geometrical<br />
constrictions in the insulating structures, which produce<br />
non-uniform electric fields by squeezing the electric field in<br />
a conductive medium, has been proposed. These structures<br />
are termed insulator-based or electrodeless DEP (iDEP or<br />
EDEP) [1]. The direction of the DEP force is dominated by<br />
the dielectric properties of the particles as well as of the<br />
medium, which are functions of frequency. Particles<br />
experiencing positive DEP forces move to the local electric<br />
field maxima; however, those experiencing negative DEP<br />
forces will be driven toward the local electric field minima.<br />
The focusing of biological cells in microdevices is a<br />
prerequisite for medical applications, such as cell sorting,<br />
counting or flow cytometry [2]. Throughput and sensitivity<br />
can be greatly enhanced by effective focusing.<br />
Hydrodynamic focusing by sheath flow [3,4] has been a<br />
widely-used method for particle focusing; however,<br />
additional buffer inlets and precise flow control are required.<br />
The dielectrophoretic confinement of particles generated by<br />
microelectrodes on the top and bottom of the channel has<br />
been combined with hydrodynamic focusing by sheath flow<br />
to improve the performance of particle focusing and<br />
throughput for detecting and counting [5]. Furthermore, the<br />
position of particles in a microchannel can be accurately<br />
manipulated by electrodeless dielectrophoresis generated by<br />
the electric field between the so-called liquid electrodes [6].<br />
Recently, a method of contactless DEP [7], so called cDEP,<br />
was proposed due to its repeatability, high-efficiency and<br />
easy fabrication. An insulator-based dielectrophoretic<br />
microdevice with effective focusing of particles was<br />
designed and fabricated in our previous work [8]. A lower<br />
conductive material of polydimethylsiloxane (PDMS) was<br />
adopted as a structure in the microchip for particle focusing,<br />
instead of a metallic pattern, to squeeze the electric field in a<br />
conducting solution and generate the regions of high field<br />
gradient. However, the metallic electrodes were made on the<br />
glass substrate to reduce the applied voltage in our previous<br />
study. The main purpose of the present study was to<br />
investigate the feasibility of applying the technique of cDEP<br />
on our design of the insulator-based dielectrophoretic<br />
microdevice with effective focusing of particles.<br />
II. THEORY AND DESIGN<br />
The DEP Force (F DEP ) acting on a spherical particle of<br />
radius R suspended in a fluid of permittivity ε is given as:<br />
m<br />
3<br />
2<br />
DEP<br />
= 2 επ<br />
m<br />
Re(<br />
CM<br />
) ∇EfRF<br />
(1)<br />
rms<br />
where Re( f CM<br />
) is the real part of the Clausius-Mossotti<br />
factor, and E rms is the root-mean-square of the external<br />
electric field, in an alternating field. The Clausius-Mossotti<br />
factor (f CM ) is a parameter of the effective polarizability of<br />
the particle. It varies as a function of the frequency of the<br />
applied field (f), as well as the dielectric properties of the<br />
particle and the surrounding medium. The<br />
Clausius-Mossotti factor for a spherical particle is<br />
represented as:<br />
362
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
**<br />
⎡ − εε ⎤<br />
mp<br />
f = ⎢ ⎥<br />
(2)<br />
CM<br />
**<br />
⎢⎣<br />
+ 2εε<br />
mp<br />
⎥⎦<br />
*<br />
*<br />
where ε and<br />
p<br />
ε are the complex permittivity of the<br />
m<br />
particle and the medium, respectively. The complex<br />
permittivity is related to the conductivity σ and angular<br />
frequency ω, which relates to ω = 2πf<br />
, through the<br />
formula:<br />
σ<br />
* εε j−≡<br />
ω<br />
(3)<br />
where j equals − 1 . Therefore, the DEP force is<br />
dependent mainly on difference between the dielectric<br />
properties of particles and the suspending medium solution.<br />
It can be either a positive DEP, which pulls the particle<br />
toward the location of the high-electric field gradient, or a<br />
negative DEP, that repels the particle away from the region<br />
of high-electric-field gradient.<br />
The schematic diagram of the cDEP microfluidic chip is<br />
depicted in Fig. 1. Four insulating structures which formed<br />
an X-pattern in the microchannel, as shown, were employed<br />
to squeeze the electric field in a conducting solution, thereby<br />
generating the regions high-electric-field gradient. The inlet<br />
flow field and the electric field were applied vertically. The<br />
negative-dielectrophoretic particles were repelled from the<br />
high- electric-field region, moving to the center of the<br />
microchannel where the flow velocity was higher. The<br />
insulator was designed 60 μm in width and 200 μm in length.<br />
The distance of the insulators along the direction of flow, as<br />
well as along the direction of the electric field, was 120 μm.<br />
The inclined angle of the insulator was 45 degrees. The<br />
particles were introduced into the microchannel and<br />
pre-confined hydrodynamically by the funnel-shaped<br />
insulating structures close to the inlet. The particles with a<br />
negative dielectrophoretic response were repelled toward the<br />
center of the constricting region. The electric field in the<br />
main microchannel was generated by using electrodes<br />
inserted into two conductive micro-reservoirs, which were<br />
separated from the main microchannel by thin insulating<br />
barriers made of 20 μm-width of PDMS [7].<br />
III.<br />
EXPERIMENTAL SECTION<br />
A. Chip Fabrication<br />
A biocompatible material of polydimethylsiloxane<br />
(PDMS) was adopted for cDEP microfluidic chip. The mold<br />
master was fabricated by using the inductively coupled<br />
plasma (ICP) dry etching technique on the silicon wafer<br />
(around 100 μm in height) to define the micropatterns, as<br />
showed in Fig. 2a. The PDMS prepolymer mixture was<br />
poured and cured on the mold master to replicate the<br />
structures (Fig. 2b). After the PDMS replica had been peeled<br />
off, the replica was bonded with the glass substrate after<br />
treatment of the oxygen plasma in the O 2 plasma cleaner<br />
(Model PDC-32G, Harrick Plasma Corp. Ithaca, NY, USA).<br />
The image of the fabricated microchip for particle focusing<br />
taken by the optical microscope was revealed in Fig. 2c.<br />
B. Apparatus and Materials<br />
A function/arbitrary waveform generator (Agilent<br />
33220A, Agilent Technology, Palo Alto, CA, USA) was<br />
employed as the AC signal source and connected to an RF<br />
amplifier (HSA-4011, NF corporation, Japan) to apply<br />
electric fields required for dielectrophoretic manipulation in<br />
the microchannel. Polystyrene particles, 10 mm in diameter,<br />
(G1000, Thermo Scientific Inc., USA) were used to<br />
investigate the efficiency of focusing. A sample of<br />
polystyrene particles with a concentration of 10 6<br />
particles/mL was injected using a syringe pump (Model KDS<br />
101, KD Scientific Inc., Holliston, MA, USA). The<br />
dielectric permittivity and conductivity of polystyrene<br />
particles at the frequency of 1 MHz are about ε r =2.6; σ=10 -16<br />
S/m, respectively [9]. The particles are suspended in a<br />
sucrose solution with an 8.62 wt% and 2.74×10 -2 wt% of<br />
K 2 HPO 4 (ε r =78; σ=4.50×10 -2 S/m). The dielectrophoretic<br />
focusing of particles was observed and recorded by an<br />
inverted fluorescence microscope (model CKX41, Olympus,<br />
Tokyo, Japan) mounting a CCD camera (DP71, Olympus,<br />
Tokyo, Japan) and a computer with Olympus DP controller<br />
image software.<br />
IV. RESULTS AND DISCUSSION<br />
To investigate the capacitive behavior of the PDMS barrier,<br />
the Wayne Kerr precision impedance analyzer 6420 was<br />
used to measure the impedance of the PDMS in DMEM<br />
medium (Gibco, Grand Island, NY, USA), which was<br />
commonly used in cell culture. The micro-reservoirs were<br />
filled with DMEM medium with an electric conductivity of<br />
0.8 S/m; besides, this medium in light red color could be used<br />
for confirm that there was no leakage from insulating barriers.<br />
The experimental set up was depicted in Fig. 3. The<br />
measured impedance of the PDMS barrier under different<br />
frequencies and the fitting curve to experimental data using<br />
the least-squares method were depicted in Fig. 4. The results<br />
revealed the capacitive behavior of the PDMS barrier. The<br />
impedance decreased with the frequency. The relative<br />
dielectric permittivity and conductivity of PDMS at a<br />
frequency of 1 MHz were 10.46 and 7.6×10 -4 S/m,<br />
respectively. The numerical simulations of the electric and<br />
flow fields, as well as the particle trajectory, were performed<br />
using the commercial software CFDRC-ACE + (ESI Group,<br />
France). Fig. 5 showed the transient simulation of the tracks<br />
of latex particles (ε r =2.6; σ=10 -16 S/m) [9] under varying<br />
electric field strengths and inlet velocity. The increase in the<br />
applied electric field significantly enhances the performance<br />
of focusing. Furthermore, decreasing inlet velocity increases<br />
the efficiency of focusing because the higher velocity results<br />
in more lateral expansion. Experimental results of focusing<br />
of fluorescent latex particles at different inlet flow rates and<br />
under varying electric field strengths of a frequency of 1<br />
MHz were demonstrated the performance of focusing, as<br />
showed in Fig. 6. The latex particles of 10 μm in diameter<br />
suspended in the sucrose medium with 8.62 wt% and<br />
2.74×10 -2 wt% of K 2 HPO 4 (ε r =78; σ=4.50×10 -2 S/m) were<br />
used to investigate the efficiency of focusing. The sample of<br />
latex particles was injected using a syringe pump. The<br />
experimental results showed that the performance of<br />
363
focusing increased both as the strength of the applied electric<br />
field increased and as the inlet velocity decreased.<br />
V. CONCLUSIONS<br />
The feasibility of applying the technique of cDEP on our<br />
design of the insulator-based dielectrophoretic microdevice<br />
with effective focusing of particles was successfully<br />
demonstrated. The results of measurement revealed the<br />
capacitive behavior of the PDMS barrier, which indicated the<br />
impedance of the PDMS decreased with the frequency. The<br />
preliminary experiments employing latex particles were<br />
conducted to demonstrate the feasibility of the present<br />
design. The design proposed herein has no need for<br />
complicated flow controls for focusing cells. Moreover, the<br />
usage of contactless DEP technique makes the<br />
insulator-based dielectrophoretic microchip mechanically<br />
robust and chemically inert. The microdevice is easy to<br />
operate and to integrate into further biomedical applications.<br />
ACKNOWLEDGMENT<br />
The authors would like to thank the National Science<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Council of the Republic of China for its financial support<br />
under contract No. NSC-99-2923-E-194-001-MY3. In<br />
addition, the National Center for High-Performance<br />
Computing for the use of computer time and its facilities is<br />
also acknowledged.<br />
REFERENCES<br />
[1] B. H. Lapizco-Encinas, B. A. Simmons, E. B. Cummings, Y.<br />
Fintschenko, Electrophoresis 25 (2004) 1695-1704.<br />
[2] S. Gawad, L. Schild and P. Renaud, Lab Chip 1 (2001) 76-82.<br />
[3] G. B. Lee, C. I. Hung, B. J. Ke, G. R. Huang, B. H. Hwei and H. F.<br />
Lai, J. Fluids Eng. 123 (2001) 672-679.<br />
[4] L. Lei, Y. L. Zhou, Y. Chen, Microelectron. Eng. 86 (2009)<br />
1358-1360.<br />
[5] D. Holmes, H. Morgan and N. G. Green, Biosens. Bioelectron. 21<br />
(2006) 1621-1630.<br />
[6] N. Demierre, T. Braschler, P. Linderholm, U. Seger, H. van Lintel<br />
and P. Renaud, Lab Chip 7 (2007) 355-365.<br />
[7] H. Shafiee, J. L. Caldwell, M. B. Sano, R. V. Davalos, Biomed.<br />
Microdevices 11 (2009) 997-1006.<br />
[8] C. P. Jen, C. T. Huang and C. H. Weng, Microelectron. Eng. 87<br />
(2010) 773-777.<br />
[9] Y. Kang, B. Cetin, Z.Wu, D. Li, Electrochim. Acta, 54, 1715 (2009).<br />
364
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Figure 1: Schematic diagram of the contactless DEP microchip<br />
for particle focusing using insulating structures.<br />
(a)<br />
(b)<br />
Figure 5: Transient simulation of tracks of negative<br />
dielectrophoretic latex particles under varying electric field<br />
strengths and inlet velocity. The relative dielectric permittivity<br />
and conductivity of particles were 2.55 and 9.28×10 -4 S/m,<br />
respectively.<br />
(c)<br />
Figure 2: (a) The scanning electron microscopy (SEM) image<br />
of the silicon mold and (b) PDMS replica. (c) The image of the<br />
chip for particle focusing taken by the optical microscope.<br />
Figure 3: The schematic illustration of the experimental set up<br />
for measuring the impedance of the PDMS barrier.<br />
Figure 6: Experimental results of focusing of fluorescent latex<br />
particles at different inlet flow rates and under varying electric<br />
field strengths (at a frequency of 1 MHz). The images of the<br />
last set of X-patterned insulating structures (the region marked<br />
by the dash line in the layout) demonstrate the performance of<br />
focusing.<br />
Figure 4: The measured impedance of the PDMS barrier under<br />
different frequencies and the fitting curve to experimental data<br />
using the least-squares method.<br />
365
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May, 2011, Aix-en-Provence, France<br />
<br />
Increasing Density of Antibody-Antigen Binding on a<br />
Sensor Surface by Controlling Microfluidic Environments<br />
Chia-Che Wu 1 *, Ling-Hsuan Hung 2 , Ching-Hsiu Tsai 3 , Yao-Lung Liu 4<br />
1,2 Department of Mechanical Engineering<br />
3 Graduate Institute of Biotechnology<br />
1,2,3 National Chung Hsing University, 250, Kuo Kuang Road, Taichung, Taiwan, 402<br />
4 Division of Nephrology, Department of Internal Medicine, China Medical University Hospital, 91, Hsueh-Shih Road,<br />
Taichung, Taiwan 404 1<br />
Tel:+886-4-22840433 x 419<br />
Email: josephwu@dragon.nchu.edu.tw<br />
Abstract- Antibody-antigen reactions are widely used in biological<br />
detection. Researchers typically use either optical or electrical<br />
measurements to examine the sensing surface for specific antigens.<br />
Poor antibody-antigen density results in poor sensitivity of<br />
biological detection. The purpose of this research is to enhance<br />
antibody-antigen density on sensing surfaces by controlling the<br />
microenvironment. Vortices of samples were produced according<br />
to the structure and conditions of the microenvironment. Finite<br />
element analysis was used to compute the velocity field, streamline,<br />
and vorticity of samples in the microenvironment. Fluorescent<br />
particles were used to show streamline of samples experimentally.<br />
Experimental results were compared to simulated ones. Finally,<br />
Turnip Yellow Mosaic Virus (TYMV) was used as the specific<br />
antigen in the experiment. Experiment results showed that the<br />
density of TYMV detected by vortex microenvironment was 16.5<br />
times greater than the density detected by a dipping method. The<br />
duration of experiment by vortex microenvironment was 2.3×10 -4<br />
times less than the duration by dipping method. This research<br />
offers a simple and efficient design that benefits rapid and<br />
real-time detection.<br />
Keywords: Antibody-Antigen Reactions, , Turnip Yellow Mosaic<br />
Virus , Vortex, Microenvironments, Dipping Method<br />
I. INTRODUCTION<br />
Due to increased biological and chemical weapon terrorist<br />
attacks, food contamination, and other medical concerns, much<br />
research attention has focused on rapid and reliable real-time<br />
detection of microbes. This research has included considerable<br />
development of micro-detectors for use in electrochemistry,<br />
optics, thermology, and acoustic applications. Immunoassay is<br />
the most widely used of all present methods for the detection,<br />
diagnosis, and quantification of pathogens. In immunoassay,<br />
we frequently make use of specifically identifying and<br />
distinguishing between antibodies and antigens. In this way,<br />
antibodies can detect the amount of antigens (pathogens, such<br />
as germs or viruses) in the solution. Enzyme-linked<br />
immunosorbent assay (ELISA) [1-2] is one of the most basic<br />
immunoassay methods. In ELISA, the colorimetric method is<br />
used to detect the concentration of microbes. Although high<br />
sensitivity is achieved throughout the process, this method<br />
demands multiple times of conjugating antibody, several hours<br />
to several days of conjugation, and more experimental<br />
procedures. Western blot, or immunoblot [3-4], is another<br />
common detection method that specifically identifies and<br />
distinguishes between antibodies and antigens to color the<br />
samples. This method, however, has some drawbacks:<br />
electrophoretic separation is time-consuming, sensitivity is<br />
poor, a certain amount of antigen in the solution is required for<br />
detection, and expensive equipment is required. For these two<br />
common methods of pathogen detection, in addition to the cost<br />
of the equipment, they have some shortcomings such as the<br />
complexity and time-consuming nature of the analysis process,<br />
and the fact that trained personnel and laboratories are required<br />
for operation and analysis.<br />
Optical or electrical pathogen diagnostic tests are also<br />
widely used in the development of biochips because of their<br />
excellent selectivity and sensitivity [5-6]. The other advantages<br />
of these methods are the ease with which they can be integrated<br />
with microchannels and, thus, form a pathogen diagnosis and<br />
analysis system [7-8]. Li et al. (2006) applied traditional ELISA<br />
theory with the microchannels formed by silicon and glass to<br />
detect Escherichia coli 0157:H7. The absorption band was 402<br />
nm, and the detected concentration range was 10–105 cells/ml<br />
after the integration of E. coli 0157:H7 with the antibody. Gao<br />
et al. [9] utilized electrokinetical driven actuators combined<br />
with the indirect ELISA theory to detect different antigens from<br />
helicobacter pylori and lactobacillus rhamnosus (control group).<br />
The detected lowest concentration was 1ng/m1, and the higher<br />
the concentration of H. pylori, the stronger was the detected<br />
fluorescence signal. Regardless of whether optical or electrical<br />
measurements are used, when the amount of microbes on the<br />
sensing field is too low, the measured signals may be too weak<br />
to distinguish. Thus, an increase in the amount of pathogens or<br />
microbes fixed in the target area would be quite helpful for the<br />
signal measurement of the microbes. The traditional method of<br />
bio-assay normally applies the dipping method with the theory<br />
of antibody and antigen reaction to fix microbes on the sensing<br />
field. Due to Brownian motion, it requires a long time for the<br />
microorganisms in the static solution to adhere to the sensing<br />
field by diffusion and, in fact, few of the microorganisms<br />
actually successfully adhere.<br />
To improve the poor adhesion efficiency of the traditional<br />
366
dipping method, hydrodynamic transmission of the fluid in a<br />
microenvironment is used to drive the motion of the<br />
biomolecules. Tabeling et al. [10] mentioned that when a fluid<br />
flows through a square-shaped downward concave slot, the<br />
fluid causes a degree of swirling flow in the corner of the slot,<br />
and the side length ratio of the square-shaped slot influences the<br />
flow state of the fluid in the slot. Bruus et al. [11] also observed<br />
that the fluid created a swirling flow in the corner of the<br />
square-shaped slot when the Reynolds number of the<br />
microchannel increased. In this study, we first applied the<br />
special structure present in the microenvironment to cause the<br />
internal fluid to swirl in multiple directions, to increase the<br />
chaos of the fluids in the microenvironment. Next, by the<br />
traction force of the chaotic fluid flow, we drove the motion of<br />
the biomolecules, raised the evenness and the coverage rate of<br />
adhesion of the biomolecules to the sensing field, reducing the<br />
time for adhesion of the biomolecules to the sensing field, and<br />
improving the efficiency and sensibility of the microbial sensor.<br />
Finally, we used the plant virus TYMV to test the effect of the<br />
adhesion of TYMV on the sensing surface of the<br />
microenvironment.<br />
II.<br />
2.1 Sensing Principle<br />
SENSING PRINCIPLE AND SIMULATION<br />
The virus, TYMV, [12] is a tymovirus of the family<br />
Tymoviridae. TYMV are propagated in cabbage leaves and<br />
stored at -20 ℃ . The structure of TYMV is simulated by<br />
RasMol, a computer program written for molecular graphics<br />
(Fig. 1) and 400 amino groups on the TYMV surface is found.<br />
These amino groups can be easily linked with antibodies and<br />
quantum dots. The TYMV we used were obtained from the<br />
Graduate Institute of Biotechnology, National Chung Hsing<br />
University, Taiwan. This TYMV had Alexa Fluor 594 applied<br />
to it to examine the virus distribution under a confocal<br />
microscope. The other type of fluorescent we used was NCD-4<br />
(N-Cyclohexyl-N-4-Dime-thylamino naphthyl carbodiimide).<br />
This can also be linked with TYMV by the thiol groups on the<br />
TYMV. We used the self-assembled monolayers (SAMs) and<br />
the linker to achieve the goal of affixing the virus to the sensing<br />
surface in the microfluidics channel. The kind of SAMs used in<br />
this paper were MUA (11-mercaptoundecanoic acid). At one<br />
end of the chemical structure of MUA are thiol groups which<br />
can be attached on the gold film of the sensing surface. At the<br />
other end of the MUA is a carboxyl functional group. This<br />
functional group can form a connection between the linker and<br />
MUA. The linker layer we used was EDC<br />
(1-Ethyl-3-[3-dimethylaminopropyl]<br />
carbodiimide<br />
Hydrochloride) and NHS (N-hydroxysuccinimide). After EDC<br />
interacts with MUA, TYMVs will be captured by designed<br />
SAMs. The compound NHS is used to prevent EDC layer<br />
hydrolysis. Therefore, we can achieve a connection between the<br />
virus and the sensing surface. After the virus with quantum dots<br />
was anchored to the sensing surface of the microfluidics<br />
channel, a confocal microscope was used to monitor the surface<br />
of the silicon membrane. A laser beam with a wavelength of<br />
615 nm was passed through an aperture and was focused by an<br />
objective lens onto a small focal volume on the sensor surface.<br />
A mixture of emitted fluorescent lights (594 nm) and reflected<br />
11-13 <br />
May, 2011, Aix-en-Provence, France<br />
<br />
laser lights from the quantum dots were then reobtained by the<br />
objective lens. A photodetection device was used to transforme<br />
the reflected light signal. The results of the confocal<br />
microscopy measurement are shown in Fig. 10(a). Only a few<br />
TYMV with quantum dots were observed on the silicon<br />
membrane and the coverage rate of TYMV was poor. Also,<br />
some large molecules or particles were stuck on the surface<br />
which might result from the aggregation of MUA, EDC, or<br />
NHS by the dipping method. This result is unfavorable if<br />
researchers want to detect viruses at extremely low<br />
concentrations in the solution.<br />
(a)<br />
(b)<br />
Fig. 1 (a) Structure of TYMV (b) amine functional groups of TYMV<br />
2.2 Microfluidics Theory<br />
When TYMV is in a microenvironment, because it cannot<br />
infect any cells and parasitize on them, the virus does not exist<br />
in a living state but in a form of chemical compound. We<br />
therefore regard it as a small particle without life. When the<br />
liquid in the microenvironment is static, the virus particles can<br />
only move by diffusion of Brownian motion caused by the<br />
continuous collision of the media molecules in the<br />
microenvironment. The coefficient of diffusion signifies the<br />
movement capability of the virus by diffusion. Through the<br />
coefficient of diffusion we can understand the movement speed<br />
of the virus in static liquid. Table 1 [13] shows the coefficients<br />
of diffusion of the different viruses in various sizes in water<br />
under room temperature.<br />
Table 1 Different virus diffusion coefficient<br />
Diameter of<br />
Virus<br />
virus (nm)<br />
Diffusion<br />
coefficients (m 2 /s)<br />
Poliovirus [13] 25 1.72×10 –11<br />
Turnip Yellow Mosaic Virus [12] 31.8 1.35×10 –11<br />
Hepatitis B virus [13] 42 1.02×10 –11<br />
Adenovirus [13] 75 5.72×10 –12<br />
Human Immunodeficiency Virus [13] 120 3.58×10 –12<br />
When the virus is in a microenvironment where the medium<br />
is water, assuming the average displacement from the starting<br />
point to the terminal point is 5 mm, and the temperature is 293<br />
K, the time required for adhesion of each virus is shown in Fig.<br />
2. From the figure, we see that 230 hours is required for the<br />
movement of the virus by diffusion. Therefore, if we apply the<br />
diffusion effect as the mechanism of TYMV for adhering to the<br />
sensing surface, the action time is extremely long and the<br />
adhesion efficiency is poor. To solve the aforementioned<br />
obstacles, our study utilizes a microenvironment to drive the<br />
movement of the TYMV, and thus raise the adhesion efficiency<br />
of the TYMV.<br />
367
Time of diffusion (hr)<br />
1200<br />
1000<br />
800<br />
600<br />
400<br />
200<br />
0<br />
11-13 <br />
May, 2011, Aix-en-Provence, France<br />
<br />
Poliovirus<br />
TYMV<br />
Hepatitis B virus<br />
Adenovirus<br />
HIV<br />
50 100 150<br />
Diameter of virus (nm)<br />
are selected to 1mm, 3mm, and 475m, respectively. The<br />
height of the cylindrical structure will be discussed in the<br />
following paragraph. First, the fluid rotates in direction Y due<br />
to flow difference caused by the level difference between the<br />
channels as the fluid enters the concaved-down surface from<br />
the inlet. Second, rotation in direction Z is created by the flow<br />
rate difference with compression of the fluid when the fluid<br />
flow passes the sides of the cylindrical structure. This structure<br />
facilitates the chaotic effect generated by the rotating flow of<br />
the fluid, enhancing the mobility of the TYMV with the chaotic<br />
flow. The chaotic streamline, might present more opportunities<br />
for sensing surface attachments.<br />
inlet<br />
Fig. 2 Time to diffuse 5 mm for viruses<br />
In micro-fluidics, usually the behavior of small particles<br />
like microbes or even virus particles can be understood through<br />
the Reynolds number and Péclet number. The Reynolds<br />
number of a particle is shown in Equation (1).<br />
av<br />
Re (1)<br />
<br />
a is the diameter of a particle and v is the velocity of flow.<br />
and are density and viscosity of fluid, respectively.<br />
The Reynolds number for TYMV in water with a speed of<br />
order 10 m/s is calculated where the diameter of TYMV is 31.8<br />
3<br />
nm. Density and viscosity of water are 1000kg m and η<br />
=10 –3 Pa-s, respectively. The Reynolds number for the TYMV<br />
(3.18×10 – 7 ) is negligibly small. A small Reynolds number<br />
means that the movement of molecules was dominated by<br />
viscosity force and that molecules stop moving immediately<br />
when the drag force is removed. The movements of viruses or<br />
molecules depend on diffusion or Brownian motion. However,<br />
regardless of whether the movement relies on diffusion or<br />
Brownian motion, the motion is rather slow in terms of virus<br />
movement. This is attributed to poor efficiency in the reaction<br />
zone attached to the microchannel.<br />
Additionally, the Péclet number is shown in Equation (2).<br />
<br />
Ul<br />
d<br />
Pe<br />
(2)<br />
<br />
a<br />
D<br />
<br />
d<br />
is molecular diffusion time and<br />
a<br />
is typical hydrodynamic<br />
transport time. U, D and l are flow velocity, diffusion<br />
coefficient of molecules and depth of microfluidic channel,<br />
respectively. Taking TYMV as an example, the diffusion coefficient<br />
in water is 1.35×10 –11 m 2 /s. The Péclet number is equal to 74.07 when<br />
the fluid flow is 10 m/s while depth of microfluidic channel is<br />
100m. This shows that hydrodynamic transmission is much<br />
more effective than molecular diffusion effect.<br />
III. MICROFLUIDICS SIMULATION & EXPERIMENT<br />
3.1 Simulation<br />
Figure 3 shows the type of microenvironment adopted for<br />
this study. The design added a cylindrical structure onto a<br />
microchannel with a cross-section area of 300m×3000m.<br />
The diameter and the height of cylinder (Fig. 3) are W c and h c ,<br />
respectively. The design used W s mm ×W s mm flat sensing<br />
surface into a h s distance that is concave down. W c , W s , and h s<br />
wc<br />
ws<br />
hc<br />
hs<br />
Au<br />
PDMS<br />
Fig. 3 Microfluidic devices to produce vortex<br />
outlet<br />
This study simulated the height of the cylindrical structure,<br />
and the vorticity of the sensing field was adopted as an indicator<br />
to determine the degree of fluid rotation. The height of the<br />
cylindrical structure, hc, was divided into heights of 300 m,<br />
400 m, 500 m, and 600 m. COMSOL Multiphysics is used<br />
to predict the performance of microfluidic device.during<br />
simulation. The vorticity simulation result is shown in Fig. 4<br />
revealing a positive relationship between the overall size of<br />
vorticity and the height of the cylinder in the sensing field. The<br />
results most significantly reveal the vorticity variation on two<br />
sides of the cylinder. It is noteworthy that the vorticity in the<br />
bottom area of the 600m cylinder actually decreases rather<br />
than increases due to high flow resistance. The aforementioned<br />
easons show that a cylinder structure with a height of 500 m<br />
generates a vorticity with a wider range and higher strength and<br />
allows the fluid to generate a rotation flow effect more easily.<br />
After the height of the cylindrical structure is determined as<br />
500 m, the flow rate of fluid into the microenvironment is then<br />
discussed. The study aims to determine the minimum rate of<br />
vortex, allowing the microenvironment to generate rotational<br />
flow in multiple directions. Hydrodynamics explains that<br />
vortices are most likely to be revealed in a higher rate of flow.<br />
Moreover, the microenvironment is combined by the<br />
microchannel with the cylindrical structure and the substrate of<br />
the concaved-down sensing surface. When flow rate increases<br />
in the microenvironment, the pressure generated on the<br />
sidewall also increases. To prevent the micro channel and<br />
substrate from collapsing due to extremely high fluid pressure,<br />
this study aims to determine the minimum rate of flow for<br />
generating chaotic flow.<br />
The vorticity component and flow chart of the sensing field<br />
of X, Y, and Z directions when the inlet flow equals 20 ml / hr are<br />
small and have no obvious variation. The flow chart also<br />
reveals that the chaotic flow effect was not created in the<br />
sensing field. When the inlet flow was constantly increased to<br />
1500 ml / hr , the fluid in the cylinder front displayed significant<br />
368
coupling rotation movement in the Y and X directions as shown<br />
in the X, Y, and Z components in Fig. 5. The area on the back of<br />
the cylinder also flows rotationally in the Z direction. Rotation<br />
in the Y direction was also generated at the end of the sensing<br />
field. The flow chart also reveals the chaotic flow effect that<br />
was generated in the microenvironment at this stage<br />
(a)<br />
(c)<br />
(d)<br />
Fig. 4 Vortices when (a) hc =300 m (b) hc = 400 m (c) hc = 500 m (d) hc =<br />
600 m<br />
(a)<br />
(b)<br />
(b)<br />
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May, 2011, Aix-en-Provence, France<br />
<br />
PDMS solidifies after evacuation and heating. PDMS can be<br />
removed from the inserts. Additionally, the atmospheric<br />
pressure plasma can modify the convection channel and the<br />
PDMS surface of the substrate. The parameter 0.5 Torr, 29.6<br />
Watt, and 15 min~20 min represents pressure, plasma<br />
efficiency, and surface modification time. The microchannel<br />
was then connected with the substrate after PDMS was used to<br />
implement the surface modification procedure.<br />
3.2.2 Fluorescence streamline<br />
For this part of the study, we utilized fluorescent particles<br />
to examine the actual streamline condition of the microfluidics.<br />
The fluorescent particles used had a grain size of 10m,<br />
filtering band was between 460 and 490 nm. An Olympus IX71<br />
fluorescent microscope was used in a dark environment to<br />
provide the optical band that excites the particles, allowing the<br />
fluorescent particles to generate fluorescence automatically<br />
after photo-excitation in a specific band. When the fluorescent<br />
particles move with the streamline in the channel, the trajectory<br />
of the fluorescent particles reveals the actual streamline<br />
condition of the microfluidics. Fig. 6 (a) shows the<br />
experimental results of the fluorescent streamline. The fluid in<br />
the figure rotates in direction Y when entering the sensing field<br />
from the left. This is due to the height difference between the<br />
bottom face of the sensing field and the inlet. When the fluidics<br />
passes the cylindrical structure, it reveals an identical result to<br />
the estimation implemented by the finite element simulation<br />
software Fig. 6(b). The flow rate difference causes the fluid to<br />
rotate toward direction Z at the back of the cylindrical structure.<br />
(c)<br />
Fig. 5 Simulation result when flow rate is 0.614 μm/s(a)Vorticity, x component<br />
(b)Vorticity, y component(c)Vorticity, z component(d)Streamline<br />
3.2 Microenvironment design<br />
After analysis simulation by finite element software, a 1<br />
mm cylindrical structure was added to the microenvironment.<br />
The height of the structure was 500 m while the rate of flow in<br />
the inlet area was 1500 ml / hr . The simulation result showed that<br />
the vortices were generated when the structure had a large<br />
speed difference. This phenomenon creates more opportunities<br />
for viruses to collide with the sensing field, enhancing the<br />
efficiency of the virus attachment to the field. We also<br />
discovered that a relationship might exist between the strength<br />
distribution of curl at the back of the cylinder and the<br />
fluorescence strength distribution of the attachment experiment.<br />
We discuss this further in Section 5.<br />
3.3 Fluorescence streamline experiment<br />
3.3.1 Fabrication of microfluidics environment<br />
Firstly, stainless steel or brass was processed by a CNC<br />
milling machine to produce the inserts designed in the study.<br />
After the processed metal inserts were cleaned, PDMS and<br />
hardener were fully mixed in a ratio of 10:1, and then PDMS<br />
was poured into the channel and substrate inserts respectively.<br />
(d)<br />
(a)<br />
Fig. 6 Streamlines by(a)fluorescent particles (b)simulation<br />
IV. FABRICATION OF RECOGNITION LAYERS<br />
First, the substrate and the microfluidic channel were<br />
fabricated by PDMS casting. A layer of Au film was deposited<br />
on the substrate by sputtering. Then, 20 mM 11-MUA<br />
(11-mercaptoundecanoic acid) solution was injected on the<br />
sensor surface by the microfluidics system. This links the<br />
11-MUA with gold films as self-assembled monolayers. EDC<br />
and NHS solutions (molar ratio, 2:1) were then injected on the<br />
sensor surface by the microfluidics system. In the final step,<br />
TYMV particles with fluorescent particles (NCD4) in the<br />
buffer were introduced over the sensor surface at a different<br />
flow rate. In the meantime, the sensor surface was rinsed and<br />
dried in FPLC (Fast Protein Liquid Chromatography) to avoid<br />
nonspecific adhesion. A confocal microscope was then used to<br />
monitor the sensor surface.<br />
To avoid the low-quality MUA molecular layer from<br />
affecting the adhesion effect of TYMV, the study adopted<br />
NCD-4 fluorescent particles as the sample and carried out tests<br />
on the MUA molecular layers. By adopting a cross-linker<br />
mechanism, NCD-4 and MUA can respond and bond more<br />
(b)<br />
369
11-13 <br />
May, 2011, Aix-en-Provence, France<br />
easily. Under a low temperature (4℃) environment, we utilized <br />
an injection system to inject an NCD-4 solution with a molar<br />
concentration of 854.99 nM into the microenvironment. This<br />
allowed the NCD-4 molecules to connect with the MUA<br />
molecules in the sensing field. Then, a confocal microscope<br />
was used to examine the fluorescence signal of NCD-4 to<br />
determine the coverage condition of the MUA molecular layer.<br />
The wavelength of excitation light for NCD4 quantum dots is<br />
405 nm and the wavelength of emitted light ranges from 400–<br />
420 nm.<br />
(a)<br />
V. EXPERIMENT RESULTS<br />
This study used a confocal microscope to examine the<br />
adhesion condition of TYMV on sensing surfaces. The results<br />
were further analyzed with the image analysis software ImageJ.<br />
The ImageJ software was developed by the National Institutes<br />
of Health (NIH). The study also adopted ImageJ to analyze the<br />
surface area of the sensing surface. To prevent the MUA<br />
molecular layer coverage rate affecting the final TYMV<br />
adhesion effect, the MUA self-fabricated molecular layer was<br />
examined using NCD-4 fluorescent molecules as the samples.<br />
Utilizing observation by a confocal microscope and analysis of<br />
the results, Image J software was used to calculate the<br />
fluorescent coverage rate of the sensing field. In this manner,<br />
we located the MUA molecule layer with optimal coverage rate<br />
and uniformity. After the coverage rate and uniformity of the<br />
MUA molecule layer was confirmed, a TYMV adhesion<br />
experiment was conducted. Again, a confocal microscope was<br />
used to observe and analyze the adhesive condition.<br />
To ensure the accuracy of the experimental data,<br />
establishing a database for the control group was necessary.<br />
Fluorescent strength was defined as 100 A.U. Background<br />
fluorescent signals under 100 A.U. were ignored. The Au/MUA<br />
fluorescence testing revealed an absorption peak when the<br />
wavelength equaled to 415 nm. However, the peak value was<br />
4A; hence, it did not affect the experiment when the wavelength<br />
was equal to 415 nm and can, therefore, be ignored. The degree<br />
of adhesion effect generated between Au and NCD-4 was then<br />
tested. When the wavelength equaled 410 mm, an absorption<br />
value of 260 A.U. was observed. This indicated that the<br />
autofluorescence of the two substances could influence the<br />
experiment result. Hence, future experiment statistics should<br />
filter the bright dots that have fluorescence strength of under<br />
260 A.U. to avoid errors. After the control group data was<br />
established, we formed the MUA molecular layer by the<br />
dipping method, and then observed and analyzed by NCD-4.<br />
Fig. 7(a) shows that the MUA self-assembled monolayer (SAM)<br />
matured by the dipping method attributed to an MUA<br />
molecular layer that was unevenly formed and had low<br />
coverage. This is mainly because the molecules rely only on<br />
diffusion movement of Brownian motion as well as the effect of<br />
molecular aggregation. The result showed that fluorescence in<br />
the sensing field only accounts for 33.37 % of the total sensing<br />
area. Figure 7 (b) indicates that the average coverage rate of the<br />
MUA SAMs is approximately 86 % in the microenvironment<br />
when the rate of flow equals 1500 ml/hr.<br />
Fig. 7(a) MUA by dipping method (best viewed in color) (b) MUA<br />
when infusion rate of pump is 1500 ml/hr (best viewed in color)<br />
To overcome the problem of low coverage when<br />
implementing TYMV adhesion with the dipping method, the<br />
study adopted a microenvironment to improve the weakness of<br />
the dipping method. The autofluorescence of<br />
Au/MUA/EDC/NHS/MES is then examined and found that it<br />
cannot be ignored. Fluorescence strength under 180 A.U.<br />
should be filtered out to avoid errors in future experimental<br />
statistics. The results of confocal microscopy measurement are<br />
shown in the dipping method. In these results, the concentration<br />
of TYMV is 10ng/ml, and the adhesion duration is ten hours.<br />
Only a few TYMV with quantum dots are apparent on the<br />
sensor surface. The coverage rate of TYMV is poor with the<br />
dipping method. The average fluorescence strength is<br />
approximately 342.46 A.U. whereas the average fluorescence<br />
coverage is approximately 4.23 %. Additionally, some large<br />
molecules or particles are stuck on the surface which might<br />
have resulted from the aggregation of MUA, EDC, or NHS by<br />
the dipping method. This result is unfavorable for researchers<br />
that wish to detect viruses at extremely low concentrations in<br />
the solution.<br />
Figure 8 indicates the results seen after utilization of a<br />
microenvironment to enable TYMV to adhere to the sensing<br />
surface. We can see here that TYMV with modified NCD-4 has<br />
a concentration of 10ng/ml, and pump infusion rate was<br />
1500ml/hr. The adhesion duration of TYMV was 1.4 minutes.<br />
The coverage was 70.1 % after analysis. When the fluorescence<br />
strength of X=2.5mm in Fig. 8 (the area directly behind the<br />
cylindrical structure) was analyzed, the area directly behind the<br />
cylindrical structure had a lower fluorescence strength while<br />
the sides of the structure revealed much higher fluorescence<br />
strength. The distribution is coincidentally identical with the<br />
curl size distribution that was simulated in this study. This<br />
phenomenon shows that the generation of curl contributes to<br />
the enhancement of the probability of viruses adhering to the<br />
sensing surface. The fluorescence strength is directly<br />
proportional to the size of curl in the position, as shown in Fig.<br />
9. Figure 10 reveals that a microenvironment with vortex could<br />
certainly enhance the adhesion effect of TYMV to the sensing<br />
surface. The fluorescence coverage rate and unit duration of<br />
adhesion efficiency in the microenvironment with chaotic flow<br />
is respectively 16.5 times and 7,102 times greater than adhesion<br />
efficiency with the dipping method. The adhesion duration of<br />
TYMV in the microenvironment with chaotic flow is 0.0023<br />
times less than that of dipping method.<br />
(b)<br />
370
VI. CONCLUSION<br />
This paper discusses how to increase adhesive density of<br />
linkers and viruses on a sensor surface in microfluidic channels.<br />
We designed a flow movement in a microenvironment to<br />
control the adhesive density of MUA and TYMV. Adhesive<br />
density of a linker (MUA) and viruses (TYMV) with specific<br />
fluorescent dyes were measured by a confocal microscope. Our<br />
results show that TYMV and MUA layers disperse randomly<br />
by the dipping method. Infusion rate, flow rate, and vortex flow<br />
affect the adhesive density of the recognition layer on a sensor<br />
surface. An adhesion density of MUA was 86 % when the<br />
infusion rate was 1500ml/hr in the microenvironment. This was<br />
2.57 times larger than the density detected by the dipping<br />
method. The virus, TYMV, could attain 70 % of adhesion<br />
densities when the infusion rate was 1500ml/hr in the<br />
microenvironment. The adhesion density was 16.5 times larger<br />
than the density detected by the dipping method. The duration<br />
of the experiment by vortex flow was 2.3×10 –4 times less than<br />
the duration by the dipping method. An interesting<br />
phenomenon was observed in that the fluorescence intensity<br />
distribution was similar to the vorticity distribution of<br />
simulation. Experimental results show that vortex flow method<br />
is able to increases the adhesive density of antigen-antibody<br />
reaction and it contributes to rapid and real-time detection.<br />
VII. ACKNOWLEDGMENT<br />
This paper is supported by the National Science Council,<br />
Taiwan.<br />
X<br />
Intensity (A.u.)<br />
in<br />
Y<br />
11-13 <br />
May, 2011, Aix-en-Provence, France<br />
1200<br />
out<br />
Fig. 8 TYMV by vortex flow (best viewed in color)<br />
4000<br />
3500<br />
3000<br />
2500<br />
2000<br />
1500<br />
1000<br />
500<br />
0<br />
0 500 1000 1500 2000 2500<br />
4000<br />
3500<br />
3000<br />
2500<br />
2000<br />
1500<br />
1000<br />
500<br />
0<br />
Y position (m)<br />
Fig. 9 Vorticity and fluorescence intensity by vortex flow<br />
Vorticity (1/s)<br />
Average fluorescent intensity (A.U.)<br />
1000<br />
800<br />
600<br />
400<br />
200<br />
0<br />
600 min<br />
Dipping method<br />
1.7 min<br />
Vortex flow<br />
100<br />
Fig. 10 Average fluorescent intensity and coverage by dipping method<br />
and vortex flow<br />
REFERENCES<br />
[1] E. Engvall, P. Perlman, Enzyme-linked immunosorbent<br />
assay (ELISA). quantitative assay of immunoglobulin G,<br />
Immunochemistry, 8 (1971) 871–874<br />
[2] R. M. Lequin, Enzyme immunoassay (EIA)/enzyme-linked<br />
immunosorbent assay (ELISA), Clin. Chem., 51 (2005)<br />
2415–2418 (2005)<br />
[3] W. N. Burnette, Western blotting: electrophoretic transfer<br />
of proteins from sodium dodecyl sulfate - polyacrylamide gels<br />
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(1981) 195–203 (1981)<br />
[4] H. Towbin, T. Staehelin, J. Gordon, Electrophoretic transfer<br />
of proteins from polyacrylamide gels to nitrocellulose sheets:<br />
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(1979) 4350–4354<br />
[5] D. B. Holt, P. R. Gauger, A. W. Kusterbeck, F. S. Ligler,<br />
Fabrication of a capillary immunosensor in polymethyl<br />
methacrylate, Biosens. and Bioelectron., 17 (2002) 95–103<br />
[6] I. S. Park, N. Kim, Thiolated Salmonella antibody<br />
immobilization onto the gold surface of piezoelectric quartz<br />
crystal, Biosens. and Bioelectron., 13 (1998) 1091–1097<br />
[7] C. W. Huang, G. B. Lee, A microfluidic system for<br />
automatic cell culture, J. Micromech. Microeng., 17 (2007)<br />
1266–1274<br />
[8] B. Steinhaus, M. L. Garcia, A. Q. Shen, L. T. Angenent, A<br />
portable anaerobic microbioreactor reveals optimum growth<br />
conditions for the methanogen, Appl. Environ. Microb., 73<br />
(2007) 1653–1658<br />
[9] Y. Gao, F. Y.H. Lin, G. Hu, P. M. Sherman, D. Li,<br />
Development of a novel electrokinetically driven microfluidic<br />
immunoassay for the detection of Helicobacter pylori, Anal.<br />
Chim. Acta., 543 (2005) 109–116<br />
[10] P. Tabeling, Introduction to Microfluidics, Oxford<br />
University Press, New York, 2005, pp95-97<br />
[11] H. Bruus, Theoretical Microfluidics, Oxford University<br />
Press, New York, 2008, pp79-81<br />
[12] K. L. Bransom, J.J. Weiland, C.H. Tasi, T.W. Dreher,<br />
Coding density of the Turnip Yellow Mosaic Virus genome:<br />
roles of the overlapping coat protein and p206-readthrough<br />
coding regions, Virology 206 (1995) 403–412<br />
[13] I. N. Serdyuk, N. R. Zaccai, J. Zaccai, Methods in<br />
Molecular Biophysics Cambridge University Press, New York,<br />
2007, pp318-335<br />
80<br />
60<br />
40<br />
20<br />
0<br />
Average fluorescent coverage (%)<br />
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<br />
Fabrication and application of iron-oxide<br />
nanoparticle/PDMS cone in lab on a chip<br />
Cheng-Chun Huang, Ming-Dao Wu, Yu-Chi Wang, Wen-Pin Shih<br />
Department of Mechanical Engineering, National Taiwan University, Taipei, Taiwan<br />
Abstract- This paper presents the fabrication and application<br />
of an iron-oxide nanoparticle/polydimethylsiloxane (PDMS)<br />
cone as a component integrated in lab on a chip. The two main<br />
functions of this component are to capture magnetic<br />
microbeads in the microfluid and to mix two laminar fluids by<br />
generating asymmetric turbulence. The iron-oxide<br />
nanoparticle/PDMS cone is fabricated by a simple method<br />
without using any mold. The uncured iron-oxide<br />
nanoparticle/PDMS is dropped on the chip by an automatic<br />
dispenser and forms the cone shape by applying the magnetic<br />
field above the top of the drop. Finally, the cone is cured at 70 o C<br />
in the microchannel of the chip.<br />
I. INTRODUCTION<br />
Microfluidic device applications in chemical and<br />
biological process have drawn more and more attention in<br />
research due to versatile advantages such as low fabrication<br />
cost, short reaction time, and low reagent consumption.<br />
However, the laminar characteristics of the flow in the<br />
microchannel make it difficult to mix two separate flow<br />
streams together. Many mixing methodologies have been<br />
reported to enhance the mixing efficiency. One of the<br />
popular mixing methods is to generate chaotic advection<br />
utilizing particular geometric structures in the microchannel<br />
[1-5].<br />
Magnetic beads are popular carriers for biological<br />
manipulation in microfluidic system under the application of<br />
magnetic field. Particularly, bio-analytic processes such as<br />
mixing, separation, capture, and recognition could be<br />
conducted and integrated utilizing magnetic beads [6-8]. In<br />
previous studies [9, 10], the magnetic fields were applied<br />
from the outside of device. If the magnetic fields could be<br />
generated in specified locations in a microchannel, the<br />
external magnet would not be needed and hence the device<br />
implementation would be made easier.<br />
In this paper, we fabricate a magnetic nanoparticle/<br />
polydimethylsiloxane (PDMS) micro-cone for mixing fluids<br />
due to its three dimensional asymmetric shape in a<br />
microfluidic chip. The fabricated micro-cone could also be<br />
used to capture magnetic beads with its magnetism.<br />
II. DESIGN<br />
Fig. 1 depicts the design of the microfluidic chip which is<br />
composed of three polymethylmethacrylate (PMMA) layers.<br />
The microfluidic chip possesses two inlet holes which allow<br />
two different fluids to enter the microchannel and then<br />
encounter the micro-cone. These two fluids can be collected<br />
in the buffer tank and then flow out through the outlet hole.<br />
There are four alignment holes on each PMMA layer for<br />
accurately assembling the microfluidic chip.<br />
Fig. 2 illustrates the functionalities of the proposed<br />
micro-cone in the microfluidic chip. Let the first fluid in the<br />
microchannel contain particles A without magnetism and<br />
particles C coated with magnetic beads while the second<br />
fluid contains only particles B. The micro-cone is proposed<br />
to separate particles A from particles C and then to mix<br />
particles A with particles B. Before the fluids passing<br />
through the micro-cone, they are laminar flows in the<br />
microchannel. The particles C would be captured by the<br />
micro-cone due to the magnetic force when the first fluid<br />
passes through the micro-cone. Beyond the micro-cone, the<br />
first and the second fluids could be mixed due to the<br />
turbulence caused by the cone shape. Therefore, particles A<br />
and B could interact. The particles C captured on the<br />
micro-cone and the mixture of particles A and B can then be<br />
analyzed for versatile applications of lab on a chip. It is<br />
worthy to mention that the position of the proposed<br />
micro-cone has to be deviated from the central line of the<br />
microchannel for enhancing the mixing effect. In our design,<br />
the deviated distance is 10% of the microchannel width.<br />
Alignment holes<br />
Inlet holes<br />
Cone<br />
Outlet hole<br />
Channel<br />
Buffer tank<br />
Central line of<br />
the microchannel<br />
PMMA<br />
Fig. 1. Schematics of the proposed micro-cone in a microfluidic chip.<br />
372
Second<br />
fluid<br />
Particle A<br />
Particle B<br />
First fluid<br />
Captured particles C<br />
Particle C coated with magnetic beads<br />
III.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Particles A and B are mixed.<br />
FABRICATION<br />
Micro-cone<br />
Fig. 2. Illustrative function of the proposed micro-cone in a<br />
microfluidic chip.<br />
Fig. 3 details the process for fabricating the microfluidic<br />
chip which is composed of three PMMA layers. All the<br />
layers are machined using a CO 2 laser. Each layer is made of<br />
a bulk PMMA plate of 1 mm in thickness. To assemble the<br />
microfluidic chip, both sides of the middle layer are coated<br />
with adhesive of 50 μm in thickness (Fig.3 (a)). Therefore,<br />
the total depth of the microchannel is 1.1 mm. The channel<br />
width is 2 mm.<br />
To fabricate the micro-cone, the PDMS prepolymer and<br />
curing agent with 10:1 weight ratio are prepared, Then the<br />
iron(III)-oxide nanoparticles are dispersed and mixed<br />
thoroughly in the PDMS. The mixture is then put in a<br />
vacuum chamber for degassing. The iron(III)-oxide<br />
nanoparticles (SIGMA-ALDRICH) are comprised of<br />
primarily the gamma-form Fe 2 O 3 which exhibits<br />
superparamagnetic behavior [11, 12]. The average diameter<br />
of the nanoparticles is 50 nm. After degassing, the<br />
nanoparticle/PDMS composite is poured into a syringe and<br />
then another degassing process is conducted. An automatic<br />
dispenser is exploited to apply the nanoparticle/PDMS<br />
composite into the microchannel. In the beginning of the<br />
dispensing process, the initial portion of the<br />
nanoparticle/PDMS composite is disregarded in order to<br />
obtain constant volume of every single drop. A reference<br />
plate with a defined area is put onto the sample stage for<br />
aligning the dispenser (see Figs. 4 and 5). Therefore, the<br />
syringe needle can be precisely placed above the designated<br />
area. After finishing the alignment, the reference plate is<br />
replaced by bottom layer of the microfluidic chip. Then a<br />
drop of nanoparticle/PDMS composite is dispensed onto the<br />
bottom layer of the microfluidic chip (Fig. 3(b)). The<br />
dispensing pressure is 0.25 MPa. After dispensing the<br />
nanoparticle/PDMS composite, a permanent magnet is<br />
placed above the composite (Fig. 3(c)). The composite forms<br />
a cone shape due to the external magnetic field. Then the<br />
micro-cone is cured at 70 o C for 30 minutes. The fabrication<br />
environment is at 23 o C and 71% relative humility. After the<br />
micro-cone is cured, the three layers of the microfluidic chip<br />
are assembled together (Fig. 3(d)).<br />
There are two setups to facilitate the fabrication process. One<br />
is the automatic dispensing system (Fig. 4), and the other is<br />
the magnetic platform for generating the cone shape of the<br />
composite (Fig. 5). The automatic dispensing system<br />
consists of a dispenser (SR-330D), two CCD cameras, and a<br />
sample stage. The dispenser is used to apply iron-oxide<br />
nanoparticle/PDMS composites. It features three-axis<br />
movement with 50 μm resolution and a controller for<br />
adjusting the dispensing pressure and duration. The two<br />
CCD cameras assist the alignment of the syringe needle on<br />
the substrate and monitoring the dispensing process. The<br />
sample stage is used to move the substrate. The magnetic<br />
platform includes a vertical manipulation stage, a permanent<br />
magnet, a CCD camera, a light source, a sample stage, and a<br />
hot plate. The permanent magnet is connected to the vertical<br />
manipulation stage of 10 nm resolution. The high resolution<br />
of the manipulation stage is necessary to finely adjust the<br />
distance between the bottom of the magnet and surface of the<br />
substrate. The CCD camera is used to monitor the<br />
deformation of the uncured nanoparticle/PDMS composite.<br />
There is also a sample alignment stages which is placed on<br />
the hot plate so that the alignment would not be deviated in<br />
the curing process.<br />
The composites with different weight ratios of the<br />
nanoparticles which are 2.22%, 6.38%, and 10.20%,<br />
respectively are added into the PDMS for evaluating the<br />
fabrication parameters. Different magnetic fields which are<br />
controlled by the distance (D) from the bottom of the<br />
magnetic to the substrate surface are applied and measured<br />
using a Tesla meter (TM-701, KANETEC CO., LTD). The<br />
results are summarized in Table 1.<br />
(a)<br />
(c)<br />
Top layer<br />
Middle layer<br />
Bottom layer<br />
Magnet<br />
Cone<br />
(b)<br />
(d)<br />
Syringe<br />
Drop<br />
Nanoparticle/PDMS<br />
Fig. 3. Process for fabricating the nanoparticle/PDMS micro-cone. (a) The<br />
chip is composed of three bulk PMMA plates. (b) The nanoparticle/PDMS<br />
composite is applied through the automatic dispenser. (c) The micro-cone is<br />
formed under the applied external magnetic field. (d) Scheme of the fabricated<br />
microfluidic chip.<br />
373
Chip<br />
Magnet<br />
Light<br />
source<br />
Hot plate<br />
Fig. 2. The application mechanism of the chip.<br />
Reference<br />
plate with<br />
alignment<br />
ring<br />
Sample<br />
stage<br />
Syringe<br />
CCD Camera<br />
Dispenser<br />
Fig. 4. Configuration of the automatic dispensing system.<br />
Vertical<br />
manipulation<br />
stage<br />
CCD Camera<br />
Chip plate<br />
Fig. 5. Configuration of the magnetic platform.<br />
Table 1<br />
Weight ratios of nanoparticles and magnet height for parametric study<br />
Weight ratio of nanoparticles<br />
2.22 % 6.38 % 10.20 %<br />
Distance 1.0 1.25 1.5<br />
between<br />
1.25 1.5 1.75<br />
substrate and<br />
magnet (mm) 1.5 1.75 2.0<br />
Table 2<br />
The parameters of weight ratio of the nanoparticles and the magnetic filed.<br />
wt.<br />
Height increase, Δh<br />
% 30 s 40 s 50 s 60 s<br />
Distance<br />
2.22 0.033 0.040 0.048 0.054<br />
1.5<br />
between<br />
6.38 0.086 0.102 0.123 0.145<br />
substrate and<br />
6.38 0.032 0.035 0.004 0.042<br />
1.75<br />
magnet (mm)<br />
10.2 0.032 0.038 0.043 0.051<br />
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May 2011, Aix-en-Provence, France<br />
<br />
IV.<br />
Sample<br />
stage<br />
RESULT<br />
Fig. 6 shows the images of PDMS droplets without<br />
nanoparticles at different dispensing duration t= 2.5, 5, 7.5<br />
10 s. The dispensing pressure is fixed at 0.25 MPa. The<br />
variation of the droplet size at the same dispensing<br />
parameters is not obvious if the nanoparticles are added into<br />
the PDMS. This parametric study is conducted for obtaining<br />
the desirable droplet diameter (d) and height (h) by<br />
controlling the dispensing duration. Since the height of the<br />
microchannel is 1.1 mm, the initial height of the PDMS<br />
droplet should be relatively smaller. All the initial heights of<br />
the PDMS droplets using different dispensing durations in<br />
our tests are smaller than 1.1 mm, as shown in Fig. 7(a). The<br />
average height is h=0.09, 0.14, 0.17, 0.19 mm for t=2.5, 5,<br />
7.5, 10 s, respectively. Since the channel width is 2 mm and<br />
the deviated distance from the cone center to the central line<br />
of the microchannel is 0.2 mm, the diameter of the PDMS<br />
droplet should be smaller than 1.6 mm. The obtained cone<br />
diameter is shown in Fig. 7(b). The average diameter is<br />
d=1.28, 1.64, 1.79 mm for t=0.25, 5, 7.5 s, respectively.<br />
Therefore, the optimum dispensing duration is 5 s in this<br />
test.<br />
For each parameter in Table 1, the image of the cone<br />
formation is captured every 5 s for 60 s. For the 2.22%<br />
weight ratio with 1.0 mm magnetic height, the 6.38% weight<br />
ratio with 1.25 mm magnetic height, and the 10.2% weight<br />
ratio with 1.5 mm magnetic height, the nanoparticle/PDMS<br />
droplets are pulled up and touch the magnet immediately.<br />
Therefore, the images of the cone formation for these three<br />
particular samples are captured every 1 s. Fig. 8(a) shows<br />
the height increase (Δh) of the droplet deformation as a<br />
function of process time. All the data in Fig. 8(a) are from<br />
the samples of the 10.20% nanoparticles with the magnetic<br />
height of 1.5 mm, 1.75 mm and 2.0 mm, respectively. For<br />
the same process time, the height increase is larger for the<br />
larger magnetic height. The error bars in Fig. 8(a) stands for<br />
the standard deviation from five samples for each fabrication<br />
parameter. The deviations are attributed to the vertical and<br />
horizontal alignment between the magnet and the center of<br />
the droplet. The non-dispersed nanoparticles in the uncured<br />
PDMS might also affect the repeatability of the cone<br />
formation.<br />
The inset of Fig. 8(a) shows the height increase of the<br />
droplet deformation for the magnetic height of 1.5 mm. The<br />
five samples are pulled up and reach the magnet at the<br />
process time of 11, 12, 19, 20 and 23 s, respectively. Because<br />
the magnet is very close to the droplet, the magnetic force is<br />
strong and sensitive to the magnetic height. The variation of<br />
the initial droplet might significantly deviate the process time<br />
for the droplet to reach the magnet. The height increase of the<br />
droplet with 1.5 mm magnetic height and the nanoparticle<br />
weight ratio of 2.22%, 6.38% and 10.20%, respectively, is<br />
shown in Fig. 8(b). The height increase is larger for the<br />
higher weight ratio of nanoparticles. Table 2 shows the<br />
height increase at the process time of 30, 40, 50 and 60 s,<br />
respectively. The evolution of the height increase with the<br />
increasing process time for the samples with 1.5 mm<br />
magnetic height and 2.22% nanoparticles is close to that with<br />
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11-13 <br />
May 2011, Aix-en-Provence, France<br />
1.75 mm magnetic height and 10.2% nanoparticles.<br />
<br />
For the 1.75 mm magnetic height in Fig. 8(a) and the (a)<br />
0.08<br />
6.83% nanoparticles in Fig. 8(b), the slope of the curve<br />
1.5<br />
changes in a piecewise manner with the process time. In the<br />
1.0<br />
beginning of applying the magnetic field, the slop is steep for<br />
0.06<br />
the process time from 0 s to 5 s in Fig. 8(a), while the slope<br />
0.5<br />
becomes smaller for the process time of 5~25 s. Fig. 9(a)<br />
0<br />
shows the fabricated microfluidic chip, and Fig. 9(b) shows<br />
0.04<br />
the cross-sectional image of the microchannel. The<br />
nanoparticle/PDMS cone is successfully fabricated in the<br />
microchannel. The height of the micro-cone is 0.3 mm, and<br />
0.02<br />
the deviated distance from the central line is 0.3 mm.<br />
(a) 0.5 mm t = 2.5 (b)<br />
t = 5 s<br />
Height difference of the drops deformation (mm)<br />
0.00<br />
0 5 10 15 20 25<br />
D=1.5 mm<br />
0 10 20 30 40 50 60<br />
Time (s)<br />
D=1.75 mm<br />
D=2.0 mm<br />
Δh<br />
Fig. 6. Images of the PDMS droplets at different dispensing durations<br />
(t). The droplet height (h) and diameter (d) are measured. (a) t=2.5 s. (b)<br />
t=5 s. (c) t=7.5 s. (d) t=10 s.<br />
(a)<br />
0.22<br />
1.26<br />
0.09 mm<br />
(c) t =7.5 s (d)<br />
t =10<br />
1.68<br />
0.18 mm<br />
1.54<br />
1.90<br />
0.14 mm<br />
0.17 mm<br />
(b)<br />
Height difference of the drops deformation (mm)<br />
0.25<br />
0.20<br />
0.15<br />
0.10<br />
0.05<br />
0.00<br />
1.5<br />
Δh<br />
1.0<br />
0.5<br />
0<br />
0 5 10 15 20 25<br />
10.20 %<br />
6.38 %<br />
2.22 %<br />
0 10 20 30 40 50 60<br />
Time (s)<br />
Droplet height (mm)<br />
(b)<br />
Droplet diameter (mm)<br />
0.20<br />
0.18<br />
0.16<br />
0.14<br />
0.12<br />
0.10<br />
0.08<br />
1.9<br />
1.8<br />
1.7<br />
1.6<br />
1.5<br />
1.4<br />
1.3<br />
h<br />
h<br />
d<br />
0.06<br />
0 2.5 5.0 7.5 10.0 12.5<br />
Dispensing time (s)<br />
2.0<br />
d<br />
1.2<br />
0 2.5 5.0 7.5 10.0 12.5<br />
Dispensing time (s)<br />
Fig. 7. Dispensing test for controlling the droplet height and<br />
diameter at the constant dispensing pressure of 0.25 MPa. (a) The<br />
relation between the droplet height and dispensing duration. (b) The<br />
relation between the droplet diameter and dispensing duration.<br />
Fig. 8. The height different, Δ h, of the nanoparticle/PDMS drops<br />
deformation. (a) the wt. % is 10.2%, the D=1.5mm, 1.75 mm and 2.0 mm. (b)<br />
D=1.5 mm, the wt. %=2.22 %, 6.38 % and 10.2%.<br />
(a)<br />
(b)<br />
0.3 mm<br />
Central line<br />
2 mm<br />
0.3 mm<br />
1.1 mm<br />
Fig. 9 The images of the fabricated microfluidic chip. (a) The fabricated chip<br />
with fluidic interconnects. (b) The cross-sectional view of the<br />
nanoparticle/PDMS cone in the microchannel.<br />
V. DISCUSSION<br />
The method to form the cone shape of the<br />
nanoparticle/PDMS composite is to exploit the uncured<br />
PDMS deformation as result of being dragged by the<br />
nanoparticles. The iron-oxide nanoparticles can be attracted<br />
by the permanent magnet. They are mixed in the PDMS and<br />
enveloped by the entangling polymer chains. As the<br />
magnetic nanoparticles are attracted by the magnetic field<br />
375
above, they would move to upward and carry the PDMS<br />
together. Because the distribution of the magnetic is<br />
hyperbolic (Fig. 10(a)), the center of the nanoparticle/PDMS<br />
composite is subjected to a relatively stronger magnetic<br />
force. Conversely, the rim of the composite is subjected to a<br />
weaker magnetic force. As a result, the nanoparticle/PDMS<br />
cone can be obtained.<br />
When the magnetic field is applied, the nanoparticles in<br />
the composite start to aggregate and to form radiated lines on<br />
the cone surface (Fig. 10(b)). There are more nanoparticles<br />
aggregating on the tip of the cone in comparison to the rim<br />
region. Therefore, a darker color around the cone tip is<br />
observed. When the nanoparticle/PDMS composite is in its<br />
uncured state, the nanoparticles are allowed to move in the<br />
PDMS. As soon as the PDMS is cured, the nanoparticles<br />
cannot move anymore. This results in the gradient<br />
magnetism due to the non-uniform distribution of the<br />
nanoparticles. Originally, the nanoparticles distribute<br />
uniformly in the PDMS. Then the magnetic field on the<br />
droplet attracts the nanoparticles which are closer to the<br />
droplet surface. These nanoparticles move upward, causing<br />
the PDMS to deform as illustrated in Fig. 11(a). The<br />
momentum of the nanoparticles causes the PDMS to deform<br />
rapidly, and hence the initial slope in Fig. 8(a) is steep. When<br />
the magnetic force and the restoration force of the polymer<br />
chains reach transient equilibrium, the speed of deformation<br />
becomes slower, as illustrated in Fig. 11(b). Although it is in<br />
equilibrium at 5 s in Fig. 8(a), the height increase reduces the<br />
distance between the magnet and the top of the droplet.<br />
Therefore, the magnetic force exerting on the nanoparticles<br />
which are close to the droplet surface increases.<br />
Nevertheless, the nanoparticles are insufficient to cause rapid<br />
deformation of the PDMS. Hence the slope becomes smaller<br />
for the process time of 5~25 s. Since the top of the droplet is<br />
subjected to a relatively stronger magnetic force, the<br />
nanoparticles in the internal portion of the droplet will move<br />
to the surface gradually and then accumulate on top of the<br />
droplet, in the center of the drop top, as illustrated in Fig.<br />
11(c). When the nanoparticles are sufficient and the droplet<br />
reaches the appropriate height, the droplet deformation<br />
becomes more rapid for the process time of 25~30 s in Fig.<br />
8(a). The nanoparticles keep moving to the droplet top and<br />
hence the PDMS deformation concentrates at the top of the<br />
droplet until the cone shape is obtained. The more<br />
nanoparticles in the PDMS, the stronger magnetism the<br />
cured cone has. However, the shorter magnetic height would<br />
cause the cone to touch the magnet. The 6.38% nanoparticles<br />
with the 1.5 mm magnetic height are the optimum parameters<br />
in this study and hence are used in the following fabrication<br />
of the microfluidic chip.<br />
Fig. 12 shows the pictures of the deformation process for<br />
the 10.2% nanoparticles at 0, 10 and 20 s in Fig. 8(a). The<br />
first set of the pictures (Figs. 12(a)-(c)) indicate that the cone<br />
for the 1.5 mm magnetic height is formed quickly at the<br />
process time of 20 s but is unstable. For the process time<br />
beyond 20 s in Fig. 8(a), the micro-cone will touch the<br />
magnet and then break into two parts, as shown in (Fig. 12(j)).<br />
The smaller portion of the composite is attached to the<br />
magnet while the large portion remains on the substrate, as<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
shown in Fig. 12(k). The process images for the magnetic<br />
height of 1.75 mm are shown in Fig. 12(d)~(f). The cone is<br />
formed gradually but cannot be clearly observed for the<br />
process time shorter than 20 s. The process images for the<br />
magnetic height of 2.0 mm are shown in Fig. 12(g)~(i). The<br />
composite barely deforms at this magnetic height because the<br />
magnetic force is not strong enough to overcome the<br />
restoration force of the polymer chains.<br />
(a)<br />
(a)<br />
(b)<br />
(c)<br />
S<br />
N<br />
S<br />
N<br />
Magnetic field<br />
S<br />
N<br />
(b)<br />
Top view<br />
Fig. 10. Illustration of the magnetism and shape-formation of the cured<br />
micro-cone. (a) Schematic of the magnetic field and nanoparticles in PDMS.<br />
(b) Top view of the fabricated micro-cone with nanoparticle aggregation.<br />
The composite is subjected to the<br />
magnetic field.<br />
Nanoparticle/PDMS<br />
composite<br />
PMMA substrate<br />
Nanoparticles move upward and reach<br />
the transient equilibrium<br />
Moving direction of<br />
nanoparticles<br />
Magnetic force<br />
Restoration force of<br />
polymer chains<br />
Nanoparticles move toward the<br />
cone tip and aggregate.<br />
Fig. 11. Illustration of the mechanism of the composite deformation.<br />
376
(a)<br />
D=1.5 mm<br />
t=0 s<br />
0.5 mm<br />
(d)<br />
D=1.75 mm<br />
t=0 s<br />
(g)<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
the arrangement of the nanoparticles. The fabricated cone in<br />
D=2.0 mm<br />
t=0 s the microchannel could be applied in lab on a chip in the<br />
future.<br />
(b)<br />
(c)<br />
(j)<br />
Breaking point<br />
D=1.5 mm<br />
t=10 s<br />
D=1.5 mm<br />
t=20 s<br />
(e)<br />
(f)<br />
D=1.75 mm<br />
t=10 s<br />
D=1.75 mm<br />
t=20 s<br />
(k)<br />
(h)<br />
(i)<br />
Smaller portion<br />
D=2.0 mm<br />
t=10 s<br />
D=2.0 mm<br />
t=20 s<br />
Larger portion<br />
Fig. 12. The process of the nanoparticle/PDMS droplet becomes a cone shape.<br />
The weight ratio of the nanoparticles is 10.2%. (a) D=1.5 mm, t=0 s. (b) D=1.5<br />
mm, t=10 s. (c) D=1.5 mm, t=20 s. (d) D=1.75 mm, t=0 s. (e) D=1.75 mm, t=10<br />
s. (f) D=1.75 mm, t=20 s. (g) D=2.0 mm, t=0 s. (h) D=2.0 mm, t=10 s. (i) D=2.0<br />
mm, t=20 s. (j) D=1.5 mm. The composite which expands between the substrate<br />
and the magnetic is about to break, (k) D=1.5 mm. The composite breaks into<br />
two parts.<br />
REFERENCES<br />
[1] H. J. Sheen, C. J. Hsu, T. H. Wu, H. C. Chu, C. C. Chang, and U. Lei,<br />
“Experimental study of flow characteristics and mixing<br />
performance in a PZT self-pumping micromixer,” Sens. Actuators, A,<br />
vol. 139, pp. 237-244, March 2007.<br />
[2] D. Stroock, S. K. W. Dertinger, A. Ajdari, I. Mezic, H. A. Stone, and<br />
G. M. Whitesides, “Chaotic mixer for microchannels,” SCIENCE,<br />
vol. 295, pp. 647-651, January 2002.<br />
[3] Y. Z. Liu, B. J. Kim, and H. J. Sung, “Two-fluid mixing in<br />
microchannel,” Int. J. Heat Fluid Flow, vol. 25, pp. 986-995, July<br />
2004.<br />
[4] A. P. Sudarsan and V. M. Ugaz, “Fluid mixing in planar spiral<br />
microchannels,” Lab Chip, vol. 6, pp. 74-82, January 2006.<br />
[5] M. Zhang, J. Wu, L. Wang, K. Xiao, and W. Wen, “A simple method<br />
for fabricating multi-layer PDMS structures for 3D microfluidic<br />
chips,” Lab Chip, vol. 10, pp. 1199-1203, May 2010.<br />
[6] S. G. Grancharov, H. Zeng, S. Sun, S. X. Wang, S. O’Brien, C. B.<br />
Murray, J. R. Kirtley, and G. A. Held, “Bio-functionalization of<br />
monodisperse magnetic tunnel junction based sensor,” J. Phys.<br />
Chem. B, vol. 109, pp. 13030-13035, April 2005.<br />
[7] M. A. M. Gijs, “Magnetic bead handling on-chip: new opportunities<br />
for analytical applications,” Microfluid Nanofluid, vol. 1, pp. 22-40,<br />
October 2004.<br />
[8] A. Sandhu, H. Handa, and M. Abe, “Synthesis and applications of<br />
magnetic nanoparticles for biorecognition and point of care medical<br />
diagnostics,” Nanotechnology, vol. 21, 442001-23, September 2010.<br />
[9] S. A. Peyman, A. Iles, and N. Pamme, “Mobile magnetic particles as<br />
solid-supports for rapid surface-based bioanalysis in continous<br />
flow,” Lab Chip, vol. 9, pp. 3110-3117, November 2009.<br />
[10] N. Pamme and C. Wilhelm, “Continuous sorting of magnetic cells<br />
via on-chip free-flow magnetophoresis,” Lab Chip, vol. 6, pp.<br />
974-980, August 2006.<br />
[11] C. Pascal, J. L. Pascal, F. Favier, M. L. E. Moubtassim, and C. Payen,<br />
“Electrochemical synthesis for the control of γ-Fe 2O 3 nanoparticle<br />
size, morphology, microstructure, and magnetic behavior,” Chem.<br />
Mater., vol. 11, pp. 141-147, January 1999.<br />
[12] L. C. A. Oliveira, R. V. R. A. Rios, J. D. Fabris, K. Sapag, V. K.<br />
Garg, and R. M. Lago, ”Clay-iron oxide magnetic composites for the<br />
adsorption of contaminants in water,” Appl. Clay Sci., vol. 22, pp.<br />
169-177, February 2003.<br />
VI. CONCLUSION<br />
An approach to fabricate the 3-D iron-oxide<br />
nanoparticle/PDMS cone in a microchannel has been<br />
demonstrated. The 6.38% weight ratio of the nanoparticles in<br />
the PDMS was used. An automatic dispenser was used to<br />
align the composite cone in the microchannel and to control<br />
the diameter of the composite droplet. It was then used to<br />
apply the uncured droplet. The dispensing pressure is 0.25<br />
MPa, and the dispensing time is 5 s. A permanent magnet<br />
connected to the vertical manipulation stage was used to<br />
attract the droplet to form the cone shape. The distance<br />
between the magnet and the PMMA substrate is 1.5 mm. The<br />
cone is cured by heating at 70 o C for 30 mins.<br />
The cone has been successfully fabricated in the<br />
microchannel. The height of the cone is 0.3 mm, and the<br />
deviated location of the cone from the central line of the<br />
microchannel is 0.3 mm. The cone has the magnetism due to<br />
377
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Diamond-based technology dedicated to Micro<br />
Electrode Arrays for Neuronal Prostheses<br />
A. Bongrain 1 , A. Bendali 2 , G. Lissorgues 3 , Lionel Rousseau 3 , B. Yvert 4 , E. Scorsone 1 , P.Bergonzo 1 , S. Picaud 2<br />
1 CEA, LIST, Diamond Sensor Laboratory, CEA/Saclay, Gif-sur-Yvette, France<br />
2 Institut de la Vision, INSERM UMRS-968, UPMC, Paris, France<br />
3 ESYCOM - ESIEE, University Paris-Est, 93162, Noisy le grand, France, lissorgg@esiee.fr<br />
4 INCIA, University Bordeaux, France<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
<br />
I. INTRODUCTION<br />
Recent advances in nanotechnology have opened new routes<br />
for the fabrication of MicroElectrode Arrays (MEAs), offering<br />
an elegant way to probe the neuronal activity distributed over<br />
large populations of neurons either in vitro or in vivo.<br />
Specific electrical stimulations can be delivered to neuronal<br />
networks when using MEAs as stimulating electrodes.<br />
Conversely, MEAs can provide a mean to record the activity<br />
of many cells simultaneously over large neuronal networks [1<br />
- 4]. These MEAs now are an increasingly common approach<br />
for neurologic pathologies treatment strategies[5, 6]: they can<br />
be used to build neural prostheses to balance function losses<br />
due to lesions or degeneration of part of the Central Nervous<br />
System (CNS) such as for Parkinson disease treatment, or for<br />
cochlear or retinal implants.<br />
The contact to the cells for such MEAs commonly uses gold,<br />
platinum, black platinum or iridium oxide as the electrode<br />
materials. Any non-optimal contact can induce reactive gliosis<br />
(Muller cells) in the vicinity of the micro electrodes producing<br />
an insulating surface between the MEA and the neuron.<br />
This paper describes an alternative approach based on the use<br />
of a non conventional material, namely Boron Doped<br />
Diamond (B-NCD), to fabricate different kind of MEAs.<br />
Indeed diamond is now considered as a promising material for<br />
micromechanical or microelectronics device applications [7].<br />
The challenge is then to build new electrodes that exhibit both<br />
a high potential window with respect to water electrolysis, and<br />
possess a high electrode reactivity which is important to obtain<br />
high signal to noise ratios. We show that B-NCD, as fabricated<br />
using nano-processing coupled with chemical vapour<br />
deposition (CVD), leads to semiconducting electrode<br />
properties with bio-inert capabilities adapted to efficient<br />
neuronal stimulation and recording.<br />
This paper is divided into three parts. We first present the<br />
specific technology developed to fabricate diamond based<br />
MEA, then second the fabrication of the MEAs and some<br />
characterisation results. We finish introducing an application<br />
of such MEAs to retinal implants.<br />
II.<br />
DIAMOND BASED TECHNOLOGY<br />
We developed a novel technology enabling the fabrication of<br />
diamond based microelectrode arrays either on silicon, glass<br />
or soft substrates.<br />
Today, the Microwave Plasma Chemical Vapour Deposition<br />
(MPCVD) technique allows the growth of polycrystalline<br />
diamond films over large areas (2 to 4 inches on a variety of<br />
substrate materials, including silicon) but the innovation<br />
remains to accurately pattern the diamond layers to define in<br />
our case the electrodes of the MEAs. Various approaches to<br />
pattern diamond layers have been investigated, the most<br />
common approach being selective etching in oxygen/argon [8]<br />
or oxygen/CF 4 [9, 10] plasma. However, the chemical<br />
resilience of diamond renders the etching step time consuming<br />
and often unsuitable for mass production. Another technique<br />
relies on the patterning of a diamond nano powder layer from<br />
which the diamond film is selectively grown [11].<br />
As we are using Boron Doped Nanocrystalline Diamond<br />
378
(BNCD) as the active microelectrode material, our approach is<br />
based on a bottom-up patterning process [12], fully compatible<br />
with standard clean room fabrication methods.<br />
In this process, see Figure 1, the wafer (typically oxidised<br />
silicon) undergoes at first a nano-seeding operation on its<br />
entire surface by spreading a colloidal solution of diamond<br />
nano-particles in suspension in ethanol (example of nanoparticle<br />
solution: provided by SP3 SEKI Technotron). Then, a<br />
protective sacrificial metal layer is used as hard mask to<br />
protect the nano-particles while the unprotected regions are<br />
etched away using an oxygen/argon plasma RIE step.<br />
This step required technological optimization and results are<br />
reported in figure 2. One can see that after 10 minutes of<br />
plasma exposure, the crystal density decreases below 10 6 cm -2 ,<br />
and after 20 minutes duration of the plasma exposure, the<br />
changes in crystal density becomes non significant, which<br />
corresponds to the end of this etching phase. Indeed all<br />
diamond nano-particles originating from the nano-seeding<br />
have been etched away and the residual density is mainly due<br />
to surface defects, the nano-seeding density remaining above<br />
10 10 cm -2 on non etched surfaces while reaching only 5.10 5<br />
cm -2 on etched surfaces.<br />
Finally, the protective metal layer is completely removed<br />
considering that it is not affecting significantly the nanopowders<br />
immobilized onto the substrate surface. One<br />
difficulty here was the control of the uniformity of the nanoparticle<br />
distribution.<br />
The final step is the local growth of diamond on pre-defined<br />
patterned electrodes.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
Residual density (cm -2 )<br />
1,0E+09<br />
1,0E+08<br />
1,0E+07<br />
1,0E+06<br />
1,0E+05<br />
1,0E+04<br />
0 5 10 15 20 25<br />
Plasma etching duration (min)<br />
Figure 2. Residual nucleation density of diamond nanoparticles<br />
versus plasma etching duration.<br />
III.<br />
MEA FABRICATION AND CHARACTERISATIONS<br />
The technological fabrication route for MEAs implies a<br />
similar process as the one described in part II. MEAs can be<br />
done either with (i) the local growth of diamond on existing<br />
metal electrodes (Titanium) and using Si 3 N 4 for insulation, or<br />
(ii) with the homogeneous growth of diamond followed by the<br />
definition of annular metallic contacts. The top side<br />
passivation is ensured using SU8. Both solutions have been<br />
tested, see Figure 3.<br />
i)<br />
ii)<br />
Figure 3. Diamond based MEA using the process i) or ii).<br />
Figure 1. Diamond based Microelectrode fabrication process.<br />
Fabricated MEAs are characterised using electrochemical<br />
measurements in ferri/ferrocyanide 1mM (cyclic voltametry<br />
and Electrochemistry Impedance Spectrocscopy (EIS)) and<br />
their performances are compared with that of Pt identical<br />
devices.<br />
In-vitro B-NCD MEAs were tested with retina of rat, i.e. (i)<br />
cultures of ganglion cells (CGC) and spinal cords, i.e. (ii)<br />
organotypic cultures of mouse spinal cords. The tests<br />
demonstrated that no difference could be observed with<br />
respect to glass control, see Figure 4. Also, no proteinic<br />
coating was found to be necessary to ensure cell growth.<br />
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May 2011, Aix-en-Provence, France<br />
We can conclude that B-NCD offers several advantages when<br />
<br />
compared to metallic materials. Its carbon surface offers high<br />
biocompatibility, and the B-NCD potential window is about<br />
twice that of Pt.<br />
We can also see on Figure 5 that retinal ganglion cells can<br />
grow their neurites onto electrodes. Cells can be seen to<br />
organize along lines (Fig. 5B) whereas they do not show any<br />
organization in the absence of similar patterns (Fig. 5A).<br />
Such fabricated B-NCD MEAs were also tested to record the<br />
spontaneous neural activity on mouse spinal cord, and they<br />
were compared with standard Pt MEAs in term of noise level<br />
found to be around 10µV peak to peak, Figure 6.<br />
Different procedures are still under investigation to obtain<br />
other significant biological measurable signals.<br />
Figure 6. Illustration of B-NCD MEAs used to record<br />
spontaneous neural activity.<br />
IV.<br />
Recorded spontaneous activity<br />
Rms noise: 10µV peak to peak<br />
MEA APPLICATION TO RETINAL IMPLANTS<br />
10 µV<br />
A<br />
Figure 4. Retinal Ganglion Cell survival on different<br />
substrates.<br />
B<br />
At present time, in the case of retinal diseases, affecting<br />
around 12 Million people in both Europe and US, among the<br />
most frequent pathologies, photoreceptor degeneration is<br />
causing blindness in both hereditary diseases like retinitis<br />
pigmentosa and non-hereditary diseases like age-related<br />
macular degeneration (AMD).<br />
In such cases of retinal dystrophies, the number of<br />
photoreceptor cells is significantly reduced, causing a<br />
progressive reduction of visual acuity and, in worst cases,<br />
complete blindness. It can cause photoreceptor degeneration<br />
following which other layers of the retina, including bipolar<br />
and ganglion cells, partially remain [13]. Amongst several<br />
strategies, the concept of retinal prostheses was developed to<br />
restore useful vision in blind patients by activating this<br />
remaining inner retinal network using subretinal stimulation,<br />
as illustrated on Figure 7 [14].<br />
This concept was also validated in several clinical trials<br />
showing that patients were able for instance to follow moving<br />
light targets and were able to identify specific known<br />
contrasted objects [15, 16]. The first existing system (from<br />
Second Sight, US) is composed by an implanted stimulating<br />
device, using a micro-electrode array (MEA), coupled with an<br />
external camera and a coding device [17].<br />
Figure 5. Retinal ganglion cell cultures on multi-electrode<br />
arrays. Note in (B) the presence of cells organized along lines<br />
when a square pattern is aligned with electrodes.<br />
Based on the patterning of nanodiamond seeds prior to growth<br />
as described in part II, it comes possible to process structured<br />
MEAs where the active area is diamond (B-NCD). Here an<br />
additional challenge is to propose a fabrication of diamond<br />
electrodes compatible with a soft substrate material, which is<br />
achieved using a sacrificial substrate lift-off technique,<br />
resulting in retinal implants embedded into polyimide.<br />
Such implants are based on 16, 32 or 64 electrodes arrays and<br />
are tested in-vivo on rats, as one can see on Figure 8.<br />
Biologists developed a surgical technique to introduce the<br />
implants in the subretinal space of blind P23H rats. After 14<br />
weeks in vivo, the eye can be fixed and histological sections of<br />
the eye are obtained to visualize the retinal tissue with respect<br />
to the implant. These preliminary results are very encouraging<br />
because no major reactive gliosis is detected in contact with<br />
the implant. The significant reduction of glial cells appearance<br />
380
for diamond electrodes when compared to metallic electrodes<br />
are demonstrating that B-NCD MEAs provide a promising<br />
solution for the design of robust neural prostheses for long<br />
term interfacing of complex nervous systems, including retinal<br />
implants.<br />
Normal retina<br />
Degenerated retina<br />
Sub retinal position<br />
Epi retinal position<br />
Figure 7. Illustration of the MEA positioning in retinal implant<br />
strategies.<br />
11-13 <br />
May 2011, Aix-en-Provence, France<br />
<br />
[4] Ensell G., Banks DJ, Ewins DL & al, “Silicon based Microelectrodes for<br />
Neurophysiology Fabrication using a gold metallization/nitride passivation<br />
system.”, J. Microelectromech.Syst, 5(2), 117-121, 1996.<br />
[5] Kipke DR, Shain W, Buzsaki G, Fetz E, Henderson JM, Hetke JF, Schalk<br />
G, “Advanced Neurotechnologies for Chronic Neural Interfaces: New<br />
Horizons and Clinical Opportunities”. J Neurosci 28:11830-11838, 2008.<br />
[6] Lebedev, M. A. and M. A. Nicolelis , “Brain-machine interfaces: past,<br />
present and future”.Trends Neuroscience 29(9): 536-546, 2006.<br />
[7] Y. Gurbuz, O. Esame, I. Tekin, W. P. Kang, J. L. Davidson, “Diamond<br />
semiconductor technology for RF device applications”, Solid-State<br />
Electronics, NO 49, p. 1055-1070, 2005.<br />
[8] J. Enlund, J. Isberg, M. Karlsson, F. Nikolajeff, J. Olsson, D. J. Twitchen,<br />
“Anisotropic dry etching of boron doped single crystal CVD diamond”,<br />
Carbon 43 , p. 1839–1842, 2005.<br />
[9] H. Uetsuka, T. Yamada, S. Shikata, “ICP etching of polycrystalline<br />
diamonds: Fabrication of diamond nano-tips for AFM cantilevers”, Diamond<br />
and Related Materials 17, p. 728–731, 2008.<br />
[10] J. Zhang, J. W. Zimmer, R. T. Howe, R. Maboudian, “Characterisation of<br />
boron-doped micro- and nanocrystalline diamond films deposited by waferscale<br />
hot filament chemical vapor deposition for MEMS applications”,<br />
Diamond and related Materials 17, p. 23-28, 2008.<br />
[11] Y. Fu, H. Du, J. Miao, “Patterning of diamond microstructures on Si<br />
substrate by bulk and surface micromachining”, Material Processing<br />
Technology, NO 132, p 73-81, 2003.<br />
[12] E. Scorsone, A. Bongrain, C. Gesset, G. Lissorgues, L. Rousseau,<br />
«Process for microstructuring a diamond film », PCT/EP2009/059575, n° WO<br />
2010/010176 A1.<br />
[13] Humayun MS, de Juan E, Jr., Weiland JD, Dagnelie G, et al. 1999.<br />
« Pattern electrical stimulation of the human retina”, Vision Res 39:2569-76.<br />
[14] J. Salzmann, O.P.Linderholm, J-L Guyomard and al, ”Subretinal<br />
electrode implantation in the P23H rat for chronic stimulations”, Br. J.<br />
Ophthalmol.; 90; 1183-1187, 2006.<br />
[15] Humayun MS, Weiland JD, Fujii GY, Greenberg R, Williamson R, Little<br />
J, et al. “ Visual perception in a blind subject with a chronic microelectronic<br />
retinal prosthesis ”, Vision Res;43:2573–81, 2003.<br />
[16] Veraart C, Wanet-Defalque MC, Gerard B, Vanlierde A, Delbeke J.<br />
“Pattern recognition with the optic nerve visual prosthesis”, Artificial Organs;<br />
27:996–1004, 2003.<br />
[17] Weiland JD, Liu W, Humayun MS. “Retinal prosthesis”, Annual Rev<br />
Biomed Eng 7:361-401, 2005.<br />
a) b)<br />
Figure 8. Example of B-NCD MEA used as retinal implants<br />
a) the MEA on a soft substrate b) a prototype in the subretinal<br />
space with retinal blood vessels clearly seen above.<br />
ACKNOWLEDGMENT<br />
The authors want to acknowledge the French ANR for<br />
granting this project referred as MEDINAS, ANR-07<br />
TecSan-014.<br />
REFERENCES<br />
[1] Blum, R.A.; Ross, J.D.; Brown, E.A.; DeWeerth, S.P., “An Integrated<br />
System for Simultaneous, Multichannel Neuronal Stimulation and<br />
Recording”, IEEE Transactions on Circuits and Systems I, vol. 54, No. 12, p.<br />
2608 – 2618, 2007.<br />
[2] Chen C., Yao D-J., Tseng S.H., Lu S-W., Chiao C-C., Yeh S-R. Micromulti-probe<br />
electrode array to measure neural signals, Biosensors and<br />
Bioelectronics 24, 1911–1917, 2009.<br />
[3] Charvet G, Rousseau L, Billoint O, Gharbi S, Rostaing J-P, Joucla S,<br />
Trevisiol M, Chauvet P, Moulin C, Goy F Mercier B, Colin M, Fanet H,<br />
Meyrand P, Guillemaud R, Yvert B. “BioMEA: A versatile high-density 3D<br />
microelectrode array system using integrated electronics”. Biosens<br />
Bioelectron, 2010.<br />
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<br />
Measurement of Diffusivity in Nanochannels<br />
Yu-Tze Tsai 1 1, 2*<br />
and Gou-Jen Wang 1 Department of Mechanical Engineering<br />
2 Graduate Institute of Biomedical Engineering<br />
National Chung-Hsing University, Taichung 40227, Taiwan<br />
Tel:+886-4-22840725 x 320<br />
Email: gjwang@dragon.nchu.edu.tw<br />
Abstract- Diffusion is the ruling manner of the migration of ions<br />
through a nanochannel. Fick’s law and its derivatives are used as<br />
the basis for diffusion mathematical modeling. In this study, a<br />
simple principle for the detection of the diffusivity of<br />
nanoparticles in a nanochannel based on the Fick’s first law is<br />
proposed. The diffusivity in a nanochannel can be estimated by<br />
simply plotting the natural logarithmic value of the electrolyte<br />
conductance difference across the nanochannel versus time and<br />
calculating its slope. Experimental results demonstrate the<br />
feasibility of the proposed nanochannel diffusivity measuring<br />
scheme.<br />
I. INTRODUCTION<br />
Most of the physiological reactions in a cell are owing to<br />
the small changes of the surrounding environment. The small<br />
variations of the ambient environment are carried out in terms<br />
of the migrations of ions through the ion channels of the cell<br />
membrane, resulting in slight molecular variations inside the<br />
cell. The migrations of calcium ions, potassium ions, and<br />
sodium ions are the well-known examples. The slight<br />
molecular variations hence induce syntheses of corresponding<br />
macromolecules such as proteins to counter the variations. The<br />
in-vivo detection of the physiological reaction induced<br />
molecular variations can provide a very useful tool for better<br />
understanding of the physiological reaction. Hence the trend of<br />
nanopore research has been pushed forward by the recent<br />
progress in nanobiotechnology. The applications of nanopore in<br />
biotechnology include ion-pumping [1], ion-channel biosensors<br />
[2], DNA sequencing [3-4], polymers moving counting [5],<br />
biosensor [6], artificial cell membrane [7, 8], and nucleic acid<br />
detection [9-12].<br />
Diffusion is the ruling manner of the migration of ions<br />
through a nanochannel. Diffusion due to concentration gradient<br />
allows particles to travel from a higher concentration region to a<br />
lower concentration region. Diffuser, mixer, reactor, and<br />
doping of semiconductor are the commonly seen applications in<br />
our daily life [13-15]. Fick’s law [16] and its derivatives are<br />
used as the basis for diffusion mathematical modeling [17-19].<br />
If effective methods for the on-line sensing of the diffusion<br />
coefficient and concentration gradient can be developed, it will<br />
provide a useful tool for the modeling and investigation of the<br />
dynamic behavior of ions in a nanochannel. Resultantly the<br />
physiological reactions of a cell due to small variations of the<br />
ambient environment can be further explored.<br />
For the measurement of diffusivity in microchannel<br />
devices, many approaches such as the on-the-flyby<br />
-electrophoresis [20], stopped flow [21], and the E-field method<br />
[22] have been proposed. Culbertson et al. measured the<br />
diffusion coefficient of microfluidic devices using a static<br />
imaging method and three dynamic methods--stopped flow,<br />
E-field method, and length method [23]. Wu et al. [24]<br />
observed that the etching rate of oxide in a nanochannel is much<br />
slow than that in a microchannel. It was presumed that the cause<br />
is the low diffusivity of the etchant molecules in a nanochannel.<br />
If the diffusivity in a microchannel was multiplied by 6.5×10 -2<br />
as the nanochannel diffusivity, the resulting etching rate could<br />
match the experimental results. However, the presumption for<br />
the nanochannel diffusivity was not further verified by a real<br />
measurement.<br />
Besides the ion concentration gradient across the<br />
nanochannel, the migration of charged nanoparticles in a<br />
nanochannel was also influenced by the electric double-layer<br />
and the Zeta potential on the channel wall [25-27]. A feasible<br />
method for the measurement of the diffusivity in a nannochanel<br />
is thus desired. In this study, a simple principle for the detection<br />
of the diffusivity of nanoparticles in a nanochannel based on the<br />
Fick’s first law is proposed. Anodic aluminum oxide (AAO)<br />
membranes are used to replace membranes with single<br />
nanochannel for the measurement of the diffusivity. A<br />
home-made electrochemical bath that can hold an AAO<br />
membrane to separate vessels with different ion concentrations<br />
is built. The across channel ionic concentration difference is<br />
estimated in terms of the conductance difference across the<br />
AAO membrane using a Wheatstone bridge circuit.<br />
II. MATERIALS AND METHODS<br />
Assuming ideal diffusive behavior, for a sufficiently low<br />
diffusivity membrane and sufficiently large vessels, the<br />
concentration profile across the membrane should become<br />
practically linear after some initial induction period. At this<br />
point, the flux of iodide would be constant across the membrane,<br />
and the corresponding concentration gradient would also be<br />
constant. This behavior is known as the constant gradient<br />
approximation (CGA) [28], and has been used elsewhere to<br />
analyze diffusion data. The schematic shown in Figure 1<br />
depicts a system in the constant gradient state. The thin line<br />
depicts the concentration of iodide throughout the system;<br />
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May, 2011, Aix-en-Provence, France<br />
constant in each vessel and a straight line with constant slope <br />
1<br />
across the membrane. The membrane apparent diffusivity is D,<br />
cs<br />
(6)<br />
<br />
M<br />
Rs<br />
the thickness is l, the area is A, and the volume of each vessel is<br />
v 1 and v 2 , each with iodide concentration c 1 and c 2 , respectively.<br />
j<br />
DA 1 1<br />
( ) j<br />
t l v1 v2<br />
(1)<br />
Substituting Fick’s law into the above equation, the time<br />
dependent concentration difference c(t) (c(t)=c 1 (t)- c 2 (t)) can<br />
be expressed as:<br />
DA 1 1<br />
ct ( ) c(0)exp( ( ) t)<br />
l v1 v2<br />
(2)<br />
Figure 2. Wheatstone bridge circuit for electrolyte conductance measurement<br />
= c(0)exp( t/ )<br />
Substituting Equation (6) into Equation (4), the diffusivity<br />
The above equation can be used to estimate the variation of of the nanochannel can be calculated as follows.<br />
concentration difference across the membrane when the<br />
l Mi<br />
l<br />
D ( ) ( ) K<br />
diffusion coefficient is known.<br />
DM<br />
Akq (1/ Rs11/ Rs2)<br />
Akq<br />
(7)<br />
R s1 and R s2 denote the resistances of the electrolytes in<br />
vessel 1 and vessel 2 , respectively. Assuming the diffusivity of the<br />
nanochannelsis fixed, i/(1/ Rs 1/ R 1 s<br />
) K<br />
2 D<br />
is a constant.<br />
Substituting Equation (7) into Equation (2),<br />
Rt<br />
() KD<br />
1 1<br />
exp( M<br />
( ) t)<br />
R(0)<br />
kq v1 v2<br />
(8)<br />
where Rt () 1/ Rs<br />
1() t 1/ Rs2()<br />
t is the conductance difference<br />
between vessel 1 and vessel 2. The above Equation can by<br />
Figure 1: Schematic of the constant gradient approximation (CGA) rewritten as Equation (9) under the natural logarithmic<br />
operation.<br />
2.1 Diffusivity measurement principle<br />
ln( Rt<br />
( ) / R(0)) KD<br />
1 1<br />
[ ( )] M<br />
K<br />
(9)<br />
M<br />
Recall the CGA system illustrated in Figure 1, the current<br />
t kq v1 v2<br />
through the membrane due to diffusion of ions can be estimated<br />
1 ln( Rt<br />
( ) / R(0))<br />
as shown in Figure 1 based on the Fick’s law.<br />
M<br />
<br />
(10)<br />
c<br />
K t<br />
i jAkq D Akq<br />
(3) Substituting Equation (10) into Equation (7), the estimation of<br />
x<br />
the diffusivity can be further simplified using Equation (11).<br />
Where k is the number of charges each ion carries, q is the<br />
l 1 ln( R( t) / R(0))<br />
magnitude of electronic charge =1.610 -19 C. The diffusion D <br />
(11)<br />
A(1/ v<br />
coefficient of the membrane can be estimated using the formula<br />
11/ v2)<br />
t<br />
shown below.<br />
When a nanochannel with knowing cross-sectional area and<br />
i x l i<br />
length is used as the passageway for the diffusion of<br />
D <br />
(4) nanoparticles, the diffusivity in the nanochannel can be<br />
Akq c Akq ( c1c<br />
2)<br />
estimated by simply plotting the natural logarithmic value of<br />
the electrolyte conductance difference across the nanochannel<br />
versus time and calculating its slope.<br />
For a given nanochannel with predesigned depth and<br />
cross-sectional area, its diffusivity can be estimated by<br />
measuring the ratio of the ionic diffusion induced current and<br />
the concentration difference across the nanochannel. The<br />
problem in hands is how to precisely measure the concentration<br />
difference.<br />
Since the conductance of an electrolyte is proportional to its<br />
ionic concentration [29-32], the across pore ionic concentration<br />
difference can be estimated by the conductance difference. The<br />
conductance of an electrolyte can be measured using the<br />
Wheatstone bridge circuit shown in Figure 2. Where R s denotes<br />
the resistance of the electrolyte and can be determined<br />
according to Equation (5). The ionic concentration can be<br />
estimated using Equation (6). M represents the Moore<br />
conductance of the electrolyte.<br />
Vref<br />
R1Vo<br />
( R1R3)<br />
Rs<br />
R2[ ]<br />
(5)<br />
V R V ( R R )<br />
ref<br />
3 o 1 3<br />
2.2 Nanochannel fabrication<br />
In this study, anodic aluminum oxide (AAO) membranes<br />
are used to replace membranes with single nanochannel for the<br />
measurement of the diffusivity. The feasibility study is shown<br />
below.<br />
According to the definition of flux, Fick’s law can be rewritten<br />
as,<br />
c<br />
ni<br />
ji<br />
- Di<br />
<br />
(12)<br />
x A<br />
i<br />
t<br />
Where n i is the total number of particles diffusing through<br />
nanochannel i of area A i within time interval t.<br />
c<br />
ni - Di Ait<br />
(13)<br />
x<br />
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May, 2011, Aix-en-Provence, France<br />
Assume there are N nanochannels in the applied AAO <br />
membrane. The total number of particles passing through the<br />
N<br />
AAO membrane during time interval t is n . The flux passing<br />
through the AAO membrane can be estimated as,<br />
N N N<br />
c<br />
c<br />
ni (- Di At<br />
i<br />
) ( DiAi)<br />
c <br />
i1 i1 x x<br />
i1<br />
(14)<br />
j - D <br />
x<br />
N N N<br />
<br />
( A) t ( A)<br />
t A<br />
<br />
i1<br />
<br />
i i i<br />
i1 i1 i1<br />
The diffusion coefficient of the AAO membrane can be<br />
calculated to be,<br />
N<br />
D ( DA)/<br />
A<br />
N<br />
<br />
(15)<br />
i i i<br />
i1 i1<br />
For those uniformly distributed nanochannels in an AAO<br />
membrane, the diffusion coefficient for each nanochannels can<br />
be assumed to be similar. It can be derived from Eq. (15) that<br />
the diffusion coefficient of the AAO membrane is close to that<br />
of the individual nanochannel. It is thus feasible using an AAO<br />
membrane to replace a nanochannel for the diffusion<br />
coefficient measurement.<br />
The AAO templates were prepared using the well known<br />
anodizing process. Aluminum foils (99.9995% pure; 175 m<br />
thick) were cleansed using ethanol and then acetone, followed<br />
by annealing at 400C for 3 hours in a vacuum.<br />
Electropolishing was then carried out using a 1:4 perchloric<br />
acid and anhydrous ethanol solution as the electrolyte, under a<br />
constant voltage of 20 V at 40 C for 2 minutes to further polish<br />
the surfaces of the foil. A =10 mm home-made Teflon fixture<br />
was used for the anodization. The anodization process was<br />
conducted using a 0.3 M oxalic acid solution as the etchant<br />
under 90 V of applied voltage at 0C for 2 hours. The remaining<br />
aluminum beneath the barrier layer was dissolved in an aqueous<br />
CuCl 2 HCl solution that was prepared by dissolving 13.45 g of<br />
CuCl 2 powder in 100 ml of 35 wt% hydrochloric acid solution.<br />
The sample was then immersed in a 30 wt% phosphoric acid<br />
solution at room temperature for 80 min to process the barrier<br />
layer by purging and pore widening.<br />
2.3 Experimental apparatus<br />
An electrochemical bath that could hold an AAO membrane<br />
to separate vessels with different ion concentrations was built<br />
(Figure 3). The size of each vessel is 2 2 2 cm 3 . A CH263A<br />
electrochemical analyzer (CH Instruments) integrated with a<br />
Wheatstone bridge circuit having R 1 =R 2 =R 3 =200 was used<br />
for the conductance measurement. The electrolyte used was<br />
potassium chloride (KCl) with initial concentrations of 1M and<br />
0.25 M in vessel 1 and vessel 2, respectively.<br />
i<br />
Figure 3. Electrochemical bath that could hold an AAO thin film to separate<br />
vessels with different ion concentrations<br />
III. RESULTS AND DISCUSSIONS<br />
Figure 4 is a SEM image of an AAO membrane. The<br />
nanopore diameter is around 80 nm and the thickness is 60 m.<br />
The rate of coverage of the nanopores is estimated around<br />
84.2%. Since the diameter of the AAO membrane is 1 cm, the<br />
pore area can be calculated to be 0.6613cm 2 . Three membranes<br />
were fabricated.<br />
Figure 4. SEM image of an AAO membrane<br />
Figure 5 shows the diffusion induced i-t curves for various<br />
AAO membranes. The numerals denote the sample number.<br />
The ion diffusion induced currents reach their steady-state<br />
conditions in less than 200 sec. In general, 80-nm diameter<br />
pores should allow both the cation (K + , 0.137 nm) and anion<br />
(Cl - , 0.181 nm) to penetrate simultaneously. Therefore, the KCl<br />
electrolysis may not be suitable for the measurement<br />
experiments. However, it was reported that there are 4.3 K + ions<br />
and 0.067 Cl - ions respectively on average flow in a nano<br />
channel due to the electric osmosis inside the nanochannel [33].<br />
It is reasonable to assume that K + ions contribute most of the<br />
diffusion current.<br />
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May, 2011, Aix-en-Provence, France<br />
<br />
y 0.0062x0.932<br />
ln( R(t)/ R(0))<br />
y 0.0069x1.292<br />
y 0.007x1.973<br />
Figure 5. Diffusion induced i-t curves for various AAO membranes<br />
Figure 6 displays the trajectories of the ratio of conductance<br />
difference ( Rt<br />
( ) / R(0)<br />
) with respect to the i-t curves shown in<br />
Figure 5. The conductance difference data were measured using<br />
the Wheatstone bridge circuit shown in Figure 2.<br />
Figure 6. Trajectories of the ratio of conductance difference ( Rt<br />
( ) / R(0)<br />
)<br />
The ( ln( Rt<br />
( ) / R(0))<br />
, t) curves are plotted in Figure 7.<br />
Since the ion diffusion induced currents as shown in Figure 5<br />
almost reached their steady-state conditions in less than 200 sec,<br />
the ( ln( Rt<br />
( ) / R(0))<br />
, t) curves were linearly approximated for<br />
the period from 0 to 180 sec. The approximation equations are<br />
listed adjacent to each curve. The slope of each<br />
( ln( Rt<br />
( ) / R(0))<br />
, t) curve can thus be calculated as depicted in<br />
Table 1. The diffusivity for each sample is calculated using<br />
Equation (11) and shown in Table 1. It can be observed that the<br />
diffusivities are close.<br />
R()<br />
t<br />
ln( )<br />
R(0)<br />
t<br />
D<br />
(m 2 /sec)<br />
Table 1. Diffusivity for various AAO membrane<br />
#1 #2 #3 mean<br />
-0.0070 -0.0062 -0.0069<br />
-0.0067<br />
0.00036<br />
2.5410 -8 2.2510 -8 2.5010 -8 2.430.13<br />
10 -8<br />
Figure 7. ( ln( Rt<br />
( ) / R(0))<br />
, t) curves for different samples<br />
IV. CONCLUSION<br />
Most of the physiological reactions in a cell are owing to<br />
the small changes of the surrounding environment. The in-vivo<br />
detection of the physiological reaction induced molecular<br />
variations can provide a very useful tool for better<br />
understanding of the physiological reaction. The small<br />
variations of the ambient environment are carried out by way of<br />
the diffusion of ions through the ion channels of the cell<br />
membrane. In this study, a simple principle for the detection of<br />
the diffusivity of nanoparticles in a nanochannel based on the<br />
Fick’s first law is proposed. Anodic aluminum oxide (AAO)<br />
membranes are used to replace membranes with single<br />
nanochannel for the measurement of the diffusivity. An<br />
electrochemical bath that can hold an AAO membrane to<br />
separate vessels with different ion concentrations is built. The<br />
across channel ionic concentration difference can be estimated<br />
in terms of the conductance difference that is measured using a<br />
Wheatstone bridge circuit. The diffusivity in the nanochannel<br />
can be estimated by simply plotting the natural logarithmic<br />
value of the electrolyte conductance difference across the<br />
nanochannel versus time and calculating its slope. The average<br />
diffusivity in an AAO membrane with nanopore diameter being<br />
around 80 nm and the thickness being 60 m was measured to<br />
be 2.430.1310 -8 m 2 /sec.<br />
ACKNOWLEDGEMENTS<br />
The authors would like to address their thanks to the<br />
National Science Council of Taiwan for their financial support<br />
of this work under grant NSC-98-2212-E-005-072- MY3.<br />
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[10] A. Meller, L. Nivon, and D. Branton, Phys. Rev. Lett. 86, 3435, 2001.<br />
[11] R. Gaspaac, D. T. Mitchell, C. R. Martin, Electrochim. Acta 49, 847, 2004.<br />
[12] J. Mathé, A. Aksimentiev, D. R. Nelson, K. Schulten, and A. Meller, Proc.<br />
Nat.Acad. Sci. USA 102, 2377, 2005.<br />
[13] G. T. A. Kovacs, Micromachined Transducers Sourcebook (McGraw-Hill:<br />
NewYork 2000) p.706.<br />
[14] M. Sato, A. Yamada, and R. Aogaki, Jpn. J. Appl. Phys. 42, 4520, 2003.<br />
[15] D. Strock, S. K. W. Dertinger, A. Ajdari, I. Mezic, H. A. Stone, and G.<br />
M.Whitesides,Science 295, 647, 2002.<br />
[16] J. Philibert, Diffusion Fundamentals 2, 1.1-1.10, 2005.<br />
[17] W. E. Alley and B. J. Alder, Phy. Rev. Lett. 43, 653, 1979.<br />
[18] M. H. Lee, Phy. Rev. Lett. 85, 2422, 2000.<br />
[19] B. Ph. van Milligen, P. D. Bons, B. A. Carreras, and R. Sánchez, Eur. J.<br />
Phys. 26,913, 2005.<br />
[20] C. T. Culbertson, S. C. Jacobson, M. Ramsey, Talanta 56, 365, 2002.<br />
[21] Y. Walbroehl, J. W. Jorgenson, J. Microcolumn Separations 1, 41, 1989.<br />
[22] Y. J. Yao, S. F. Y. Li, J. Chromatograhic Science 32, 117, 1994<br />
[23] C. T. Culbertson, S. C. Jacobson and J. M. Ramsey, Talanta, 56 (2),<br />
365-373, 2002<br />
[24] C. Wu, Z. Jin, H. Ma, S. Lin, Y. Wang, J. Micromechanics and<br />
Microengineering 16, 2323, 2006.<br />
[25] G. L. Fain: Molecular and Cellular Physiology of Neurons (Harvard<br />
UniversityPress, New York 1999) p.63<br />
[26] R. B. Schoch, H. V. Lintel, and P. Renaud, Physics of fluids 17, 100604,<br />
2005<br />
[27] J. F. Smalley, M. D. Newton, and S. W. Feldberg, Electrochem. Commun.<br />
2, 832, 2000.<br />
[28] K. A. Snyder, Concrete Science and Engineering 3, 216, 2001<br />
[29] D. Q. Li, Microfluid Nanofluid, 1, 1, 2004.<br />
[30] D. Monk, Controlled structure release for silicon surface micromachining,<br />
Ph.D. thesis, University of California, Berkeley 164-180, 1993.<br />
[31] G. A. Bozhikov, G. D. Bontchev, P. I. Ivanov, A. N. Priemyshev, O. D.<br />
Maslov, M. V. Milanov, and S. N. Dmitriev, J. of Radioanalytical and<br />
Nuclear Chemistry 258, 645, 2003<br />
[32] T. Shedlovsky, A. S. Brown, D. A. Macinnes, Trans. Electrochem. Soc. 66,<br />
1934.<br />
[33] C. L. Gardner, W. Nonner, and R. S. Eisenberg, J. Computational<br />
Electronics 3, 25-31, 2004.<br />
Dr. Gou-Jen Wang received the B.S. degree on 1981<br />
from National Taiwan University and the M.S. and<br />
Ph.D. degrees on 1986 and 1991 from the University<br />
of California, Los Angeles, all in Mechanical<br />
Engineering. Following graduation, he joined the<br />
Dowty Aerospace Los Angeles as a system engineer<br />
from 1991 to 1992. Dr. Wang joined the Mechanical<br />
Engineering Department at the National Chung-Hsing<br />
University, Taiwan on 1992 as an Associate Professor<br />
and has become a Professor on 1999. From<br />
2003-2006, he served as the Division Director of<br />
Curriculum of the Center of Nanoscience and Nanotechnology. Since 2007, he<br />
has been the joint Professor and Chairman of the Graduate Institute of<br />
Biomedical Engineering, National Chung-Hsing University, Taiwan. On 2008,<br />
he served as the Conference Chair of the Microfabrication, Integration and<br />
Packaging Conference (April/2008, Nice, France). From 2009, he is a<br />
Committee member of the Micro- and Nanosystem Division of the American<br />
Society of Mechanical Engineers. His research interests include MEMS,<br />
biomedical micro/nano devices, nano fabrication, and dye-sensitized solar<br />
cells.<br />
386
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May 2011, Aix-en-Provence, France<br />
<br />
Energy Harvesting System for Cardiac Implant<br />
Applications<br />
Martin Deterre (1,2,3) , Bertrand Boutaud (1) , Renzo Dalmolin (1) , Sébastien Boisseau (4) , Jean-Jacques Chaillout (4) ,<br />
Elie Lefeuvre (2,3) , Elisabeth Dufour-Gergam (2,3)<br />
(1) Sorin CRM SAS, Clamart 92143, France<br />
(2) Univ Paris- Sud, Laboratoire IEF, UMR 8622, Orsay 91405, France<br />
(3) CNRS, Orsay 91405, France<br />
(4) CEA-LETI, MINATEC, Grenoble 38054, France<br />
Abstract- The miniaturization process in active medical<br />
implantable devices is driving the development of novel energy<br />
sources such as small volume, high longevity energy harvesting<br />
systems. In this study, we present an approach for the design<br />
of an inertial energy scavenger powering cardiac implants<br />
from heart generated vibrational energy. The heart<br />
acceleration spectrum has been measured and analyzed.<br />
Achievable power level and design parameters are determined<br />
from a spectral analysis to about 100µW before electronics<br />
efficiencies for a 0.5 cm 3 volume.<br />
I. INTRODUCTION<br />
Advances in microfabrication and bio/chemical<br />
engineering techniques are now enabling a large variety of<br />
miniaturized implantable systems for sensing, health<br />
monitoring or deficiency treatments. This progress is<br />
driving physicians and patients to express an increasing<br />
need for miniaturized implantable devices as they are<br />
offering less invasive implantation procedures, greater<br />
comfort for the patient, improved performance, and often<br />
provide innovative measurements and treatments [1]. Fig. 1<br />
illustrates the recent remarkable expansion of the<br />
application field of these devices.<br />
Fig. 1. Broadening diversity of the implantable medical devices<br />
applications [1].<br />
These devices most often need to include an energy<br />
source to power their active elements, such as sensing<br />
components or transmission modules, while keeping the<br />
size at the smallest level. Some progress has been made in<br />
battery technology, but batteries have more and more<br />
difficulties to follow the size reduction rhythm of the active<br />
components without significantly shortening the device<br />
lifetime [2, 3]. An alternative approach is to harvest the<br />
energy available from the surrounding environment. But<br />
traditionally energy harvesting devices can produce only a<br />
limited amount of power as the quantity of wasted energy to<br />
be harvested is small. Hence, first energy harvesting<br />
applications were limited to very low duty cycle systems.<br />
But progress in electronics power management in<br />
conjunction with the above-mentioned miniaturization<br />
process is now increasingly reducing sensors and<br />
miniaturized devices power requirements. In the meantime,<br />
performances and efficiencies of harvesting devices are<br />
improving [4, 5, 6]. This opens energy harvesting power to<br />
an always greater number of applications including medical<br />
implants. Furthermore, the substantial amount of energy<br />
produced by the human body motivates the development of<br />
an element that could extract a part of it. This humangenerated<br />
energy is available at various locations in the<br />
body and can take different forms: dissipated heat, inertia,<br />
muscle contraction, joint movement, heel strike, etc…<br />
Numerous types of human body energy sources are<br />
presented in a study by Starner [7, 8]. For instance,<br />
consumed power levels are calculated to be in the order of<br />
several watts from body heat, about one watt from breathing<br />
and one watt from blood pressure. The latter energy source<br />
has been exploited by Clark and Mo [9] where a<br />
piezoelectric membrane for blood pressure variation energy<br />
harvesting has been studied. Some commercial applications<br />
of human-powered devices have already been developed,<br />
such as shake-driven flashlights, thermal or inertia driven<br />
wristwatches or heel-strike powered LEDs to name a few<br />
[10]. A more extensive review of human body energy<br />
scavenging microsystems has been published by Romero et<br />
al. [11].<br />
387
This phenomenon is now reaching the particular case of<br />
pacemakers, where power consumption and theoretical<br />
generated power from a reasonably sized energy harvester<br />
are both reaching a value of several tens of microwatts.<br />
Goto et al. [12] have already proven the feasibility of using<br />
an energy harvesting system to power a mongrel dog’s<br />
pacemaker. In their work, they removed the powergenerating<br />
mechanism from a SEIKO kinetic watch and<br />
encapsulated it in a polyvinyl case. SEIKO’s energy<br />
harvesting system is based on a rotating eccentric mass<br />
transmitting its energy through a gear train to a rotor that<br />
generates a voltage electromagnetically. This device is then<br />
placed on the right atrioventricular wall of the dog’s heart.<br />
Although the extracted energy (13 µJ/heartbeat) was lower<br />
than the pacemaker consumption (50 µJ/heartbeat), the<br />
feasibility of an energy harvester powered pacemaker is<br />
envisioned. Tashiro et al. also addressed this subject in [13]<br />
where they present an experiment of an electrostatic system<br />
harvesting enough energy from the motion of a canine heart<br />
wall to power a pacemaker. However, this system is so<br />
cumbersome that it would be impossible to implant and it<br />
had to stay on a simulation table for this proof of concept<br />
experiment.<br />
II. INERTIAL ENERGY HARVESTERS<br />
The vast majority of up-to-date energy harvesters are<br />
based on inertial power generation [6, 14, 15]. This focus is<br />
due to two main reasons: vibrations are widespread in our<br />
environment and acceleration is inherently transferred<br />
through packaging, which greatly helps sealing and<br />
integration. These considerations apply for harvesting the<br />
vibrational energy near the heart. While some of the<br />
conventional energy harvesting technologies are not<br />
applicable such as photovoltaic or thermoelectric<br />
conversions as the body is mostly opaque and<br />
thermoregulated, heart beats are providing a continuous<br />
source of vibrational energy. Additionally, the inertial<br />
harvesting device can be properly encapsulated in a rigid<br />
package which helps biocompatibility and integration.<br />
Under the assumption that the transducing force<br />
(electromagnetic, electrostatic or piezoelectric) [6, 15] is<br />
acting as a viscous damper, a typical inertial energy<br />
harvesting system can be modeled as in Fig. 2, following<br />
the analysis of [16].<br />
Fig. 2. Mechanical system of an inertial energy harvester with viscous<br />
damping transduction.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
In this figure, m represent the proof mass, k represents<br />
the stiffness of the spring attaching the proof mass to the<br />
frame, b m and b e respectively the mechanical and transducer<br />
(electrical) viscous damping constants, y(t) the<br />
displacement of the frame and z(t) the relative displacement<br />
of the proof mass. The equation of motion can be written as:<br />
. (1)<br />
If the excitation is harmonic at an angular frequency ω, we<br />
can analytically find an expression for the displacement and<br />
the mean power of the transducing force [14] is expressed<br />
as:<br />
<br />
<br />
<br />
<br />
<br />
<br />
, (2)<br />
<br />
where ω n represents the resonant angular frequency ⁄ ,<br />
and ζ e and ζ m represent the normalized electrical and<br />
mechanical damping ratios , ⁄ 2√, and Y 0 is the<br />
frame motion amplitude. This expression shows that in<br />
order to maximize the output power, the system resonant<br />
frequency should match the excitation frequency as closely<br />
as possible. Additionally, the electrical damping ratio<br />
should be equal to the mechanical damping ratio and they<br />
need to be as small as possible. However, one should be<br />
careful of the displacement range that will greatly increase<br />
when the damping decreases.<br />
As the excitation is rarely purely harmonic, the response<br />
to the whole excitation spectrum has to be analyzed. If the<br />
spectrum has a narrow bandwidth and is not subjected to<br />
shift, then a high quality factor harvester centered on the<br />
same frequency can generate a high mechanical<br />
amplification hence a high power as expressed in (2). The<br />
limitation comes then from the travel range, as the<br />
amplitude of the mass movement is largely amplified by the<br />
same quality factor. Hence, high quality factor inertial<br />
harvesting systems are best suited for narrow and stable<br />
spectrum, low amplitude excitation. This is typically<br />
interesting for industrial applications or for machines that<br />
vibrate at specific known frequencies, such as the electrical<br />
grid frequency. When the excitation spectrum is wide or has<br />
an unsteady peak, energy harvesters should be damped<br />
further in order to provide mechanical amplification for a<br />
broader range of frequency, even though the amplification<br />
magnitude is lower. This principle has been applied by<br />
Despesse et al. in [17] where highly damped electrostatic<br />
harvesters able to harvest wide spectrum vibrations such as<br />
cars, drill or metallic stairs vibrations are presented.<br />
III.<br />
HEART ACCELERATION ENERGY HARVESTER<br />
A. Heart acceleration spectrum<br />
To predict the amount of energy that can be harvested<br />
from the heart acceleration and to determine the harvesting<br />
device and transducer characteristics, the acceleration<br />
spectrum of the heart has to be measured. Therefore we<br />
have implanted different types of accelerometers (one- or<br />
388
three-dimensional sensors) inside several heart cavities and<br />
recorded the acceleration. We then transferred these<br />
measures into the frequency domain to analyze the<br />
spectrum.<br />
It has been found that the main excitation component is<br />
concentrated on the heartbeat frequency. This frequency lies<br />
generally in the 1-1.5 Hz range, but is subjected to<br />
continuous change depending on the individual’s activity<br />
and can go up to three hertz during physical exercise.<br />
Additionally, it has been found that the spectrum shows an<br />
interesting plateau in the 10-30 Hz frequency range, which<br />
corresponds to the width of the main acceleration impulse<br />
in time domain (tens of microseconds). The amplitude of<br />
the acceleration is found to be approximately three times<br />
more important in the ventricle than in the atrium. One of<br />
the principle challenges in the design of a heart inertial<br />
harvester consists in the considerable variations of the<br />
acceleration spectrum shape and amplitude from a patient to<br />
the next, and also for different conditions on a single patient<br />
(heart activity, heart health or artificial stimulation to name<br />
a few). The magnitude of these variations can go up to 50%<br />
depending on the cases. For this reason, it has been chosen<br />
to illustrate the acceleration spectrum with a typical shape<br />
of what can be found in the right atrium (Fig. 3), keeping in<br />
mind that the variance is very large.<br />
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<br />
one gram and can be easily extrapolated for higher masses.<br />
Simulation results are shown in Fig. 4 where a mechanical<br />
damping coefficient of ζ m =0.005 has been chosen (typical<br />
parameter for silicon based microsystems). The power<br />
profile depends greatly on the electrical damping factor ζ e ,<br />
as the latter determines the broadness and magnitude of the<br />
mechanical amplification spectrum. In an electrostatic<br />
harvesting system, this electrical damping can be tuned to<br />
optimize the output power by playing with the electrical<br />
voltage, the electrode pattern or the downstream electronics.<br />
The same considerations apply for electromagnetic systems,<br />
but these typically require a permanent magnet or are very<br />
sensitive to magnetic field. As implants are likely to require<br />
compatibility with MRI imaging systems, the latter<br />
transducing method is discarded in this study. In the<br />
piezoelectric transduction case, the electrical damping<br />
depends mainly on the piezoelectric material properties and<br />
geometry.<br />
Fig. 4. Simulated output power per gram of proof mass as a function of the<br />
harvesting system resonant frequency for different electrical damping<br />
factors. The inset shows a close-up on the 20-30 Hz plateau.<br />
Fig. 3. Typical shape of the acceleration spectrum in the right atrium.<br />
The frequency plateau around 20 Hz could<br />
advantageously be targeted for energy harvesting<br />
applications, as high frequencies generally provide better<br />
power performances for a lower volume and shorter travel<br />
range, as well as being more suitable for MEMS<br />
fabrication. In accordance to the above-mentioned<br />
discussion, this type of spectrum (wide and prone to shift) is<br />
appropriate for a low quality factor, significantly damped<br />
energy harvesting system.<br />
B. Generated power<br />
By applying the same method as in Despesse et al. [17]<br />
and integrating (2) where Y 0 is deduced from the<br />
acceleration (a=Y 0 ω 2 ) for our typical input spectrum<br />
depicted in Fig. 3, we can theoretically determine the<br />
recoverable power in function of the energy harvester<br />
resonant angular frequency ω n . As the output power is<br />
proportional to the proof mass (2), simulations are run for<br />
As expected from the above discussion, the highest<br />
output power is generated around the heartbeat frequency<br />
for low damping systems. Also, the plateau in the low tens<br />
of hertz frequency range can be identified and designing a<br />
system for this frequency range could provide up to about<br />
30 µW/g of power. The fact that the power level is higher<br />
for ζ e =0.1 than for ζ e =0.01 and for ζ e =1 confirms that the<br />
electrical damping has to be carefully chosen in order to<br />
collect a wide part of the spectrum without reducing<br />
excessively the displacement amplitude. In the present case,<br />
ζ e =0.1 seems to be a reasonable choice. Although power<br />
levels are about ten times higher at very low frequency (1-3<br />
Hz), the system cannot be reasonably designed for those as<br />
it would require extremely compliant springs when coupled<br />
to the required mass. Low resonant frequency<br />
microfabricated energy harvesters can implement high<br />
aspect ratio parylene springs to further reduce the stiffness<br />
[18, 19]. Alternatively, more complex systems include a<br />
frequency up-conversion system that transfers the energy<br />
collected by a low frequency element to a high frequency<br />
high efficiency transducer [20-22]. However, it often<br />
includes non-MRI compatible magnetic components or<br />
suffers from mechanical impacts.<br />
389
C. Proof mass displacement<br />
Moreover, the proof mass displacement is a critical<br />
parameter to determine as stated previously. From the<br />
displacement expression derived from (1) and integrated for<br />
the acceleration spectrum we can simulate the amplitude of<br />
the proof mass displacement as a function of the system<br />
resonant frequency (Fig. 5). This simulation confirms the<br />
fact that a freely resonating system below approximately<br />
five hertz is inconceivable as it would induce a<br />
displacement of several centimeters. Such a system would<br />
be space-constraint and would require mechanical stops<br />
against which the proof mass will bump into regularly,<br />
hence reducing the mechanical reliability of the system.<br />
Furthermore, the electrical damping factor ζ e has a<br />
significant effect on the displacement amplitude as shown<br />
in Fig. 5. A high damping factor is needed to limit the travel<br />
range, keeping in mind that it could reduce the output<br />
power as shown in Fig. 4.<br />
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May 2011, Aix-en-Provence, France<br />
<br />
factor and fatigue resistance). Typical dimensions of such<br />
flexible springs made in the bulk of a silicon wafer are in<br />
the order of a few tens of micrometers for the width and a<br />
couple of millimeters for the length. Overall, the size of a<br />
complete harvester system for this application has to be in<br />
the order of 15 x 7 x 5 mm 3 to fit all the components as well<br />
as accommodate for the proof mass travel range.<br />
IV.<br />
CONCLUSION<br />
Through in-situ measurements, a typical shape of the<br />
heart acceleration spectrum has been determined. We<br />
presented a preliminary design study of an inertial energy<br />
scavenger able to provide 100 µW of power before<br />
application of the transducer and the downstream power<br />
management electronics efficiency coefficients. It consists<br />
of a mass of 3.5 g of tungsten with millimeter long, tens of<br />
microns wide connecting arms in the bulk of a silicon<br />
wafer. The volume of the whole system is expected to be in<br />
the order of 500 mm 3 . This energy harvester module could<br />
be implanted in a pacemaker or other implant on the heart<br />
and provide enough energy for battery-less autonomous<br />
operation.<br />
ACKNOWLEDGMENT<br />
Heart acceleration measurements were conducted by<br />
Sorin CRM Clinical Research and Advanced Research<br />
departments through the help of Alaa Makdissi.<br />
Fig. 5. Simulated displacement of the proof mass as a function of the<br />
harvesting system resonant frequency for different electrical damping<br />
factors.<br />
D. System design<br />
The best compromise between power output, travel range<br />
and frequency shift tolerance seems to be for medium<br />
electrical damping (ζ e = 0.1) and a resonant frequency<br />
around 25 Hz. For these parameters, we obtain a smooth<br />
harvesting spectrum, a displacement of a few millimeters<br />
for approximately 30 µW per gram output power.<br />
Considering the electrical consumption of a pacemaker<br />
and the efficiencies of the transduction and the downstream<br />
power management electronics, an approximate power of<br />
100 µW is required. This corresponds to about 3.5 g of<br />
proof mass for our system. In order to limit the volume to a<br />
fraction of a cubic centimeter, the proof mass has to be<br />
made of a high density material. The choice of a tungsten<br />
alloy seems natural due to its very high density (ρ ≈ 17.5<br />
g/cm 3 ) and its reasonable price and manufacturability.<br />
Hence, the proof mass has a volume of 200 mm 3 . Then, we<br />
can determine the system stiffness k. For a 25 Hz resonant<br />
system, this corresponds to approximately k = 100 N/m.<br />
The springs that connect the proof mass to the frame can be<br />
made in microstructured silicon for ease of fabrication as<br />
well as mechanical performances (high mechanical quality<br />
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[20] H. Kulah and K. Najafi, “An electromagnetic micro power<br />
generator for low-frequency environmental vibrations”, 17th IEEE<br />
International Conference on Micro Electro Mechanical Systems,<br />
2004, pp. 237 - 240<br />
[21] T. Galchev, H. Kim, and K. Najafi, “A Parametric Frequency<br />
Increased Power Generator for Scavenging Low Frequency<br />
Ambient Vibrations”, Procedia Chemistry 1, 2009, 1439–1442<br />
[22] L. Gu and C. Livermore, “Impact-driven, frequency up-converting<br />
coupled vibration energy harvesting device for low frequency<br />
operation”, Smart Mater. Struct. 20, 2011, 045004 (10pp)<br />
doi:10.1088/0964-1726/20/4/045004<br />
391
11-13 May 2011, Aix-en-Provence, France<br />
Behavioural Modelling of MEMS oscillators<br />
and Phase noise simulation<br />
G. Papin 1 , R. Levy 1 , G. Lissorgues 2 , P. Poulichet 2<br />
1<br />
ONERA-DMPH, 29 av. de la division Leclerc, 92322, Chatillon, France, Guillaume.Papin@onera.fr, Raphael.Levy@onera.fr<br />
2<br />
ESIEE, University Paris-Est, 93162, Noisy le grand, France, lissorgg@esiee.fr, p.poulichet@esiee.fr<br />
Abstract- MEMS oscillators offer interesting prospects in<br />
terms of performance. Behavioural modelling in Verilog-A and<br />
phase noise simulation can help improving the current<br />
performances. The proposed macro-model permits multiphysic<br />
simulations including mechanical, piezoelectric and<br />
electrical analytic descriptions. Phase noise analysis are then<br />
performed with this model and compared to the standard<br />
Leeson phase noise calculation.<br />
Key words- MEMS, resonator, oscillator, Verilog-A model,<br />
phase noise<br />
I. INTRODUCTION<br />
Vibrating MEMS sensors such as accelerometers [1], [2] or<br />
quartz crystal microbalances [3] use a resonator driven at its<br />
resonance by an oscillator circuit. The output of such sensors<br />
is the frequency of the output signal representing the physical<br />
measured data. Their performances are limited by frequency<br />
stability and phase noise.<br />
In order to predict the behaviour and all phenomenons<br />
within MEMS oscillators, a multiphysics model including the<br />
mechanical behaviour of the resonator, the transduction type<br />
and the electronics is under development. The Cadence<br />
software suite enables the model development in Verilog-A –a<br />
modelling language adapted for multiphysics design– and<br />
phase noise analysis with Virtuoso SpectreRF Simulator.<br />
Firstly the multiphysics verilog-A model is described, then<br />
simulations are run and compared to experimental<br />
measurements, and finally phase noise analysis are performed<br />
and discussed.<br />
II. MODEL DESCRIPTION<br />
The MEMS oscillator model is divided into two blocks<br />
(figure 1): the resonator including its mechanical behaviour<br />
and piezoelectric transduction, and the oscillator circuit. This<br />
later block is split up into two sub-blocks: the<br />
transconductance amplifier and the feedback amplifier<br />
(Automatic Gain Control). Each block contains the equations<br />
that describe its own physical behaviour.<br />
Automatic Gain<br />
Control<br />
Fig. 1. MEMS oscillator model<br />
A. The resonator<br />
The resonator mechanical behaviour in one dimension is<br />
described by the spring-mass-damper system shown in figure<br />
2.<br />
Fx<br />
Fig. 2. The spring-mass-damper system.<br />
The equation is given by (1):<br />
m x<br />
+ ρ x<br />
+ k x = F<br />
(1)<br />
x<br />
F x is the excitation force, x the mechanical displacement, m<br />
the mass, k x the spring constant and ρ x the damper constant.<br />
Its electrical equivalent representation including the<br />
mechanical and the piezoelectric domains is an oscillating<br />
RLC circuit (figure 3).<br />
L1<br />
V x<br />
C0<br />
m o<br />
Resonator<br />
Oscillator<br />
circuitry<br />
T1<br />
k x<br />
x<br />
R1<br />
x<br />
C1<br />
Fig. 3. The resonant RLC circuit.<br />
The transformer T1 describes the piezoelectric transduction.<br />
Here, piezoelectricity is used to transform mechanical energy<br />
into electrical energy. The 1 st order piezoelectricity<br />
transduction is described in the Verilog-A model by equations<br />
(2) and (3).<br />
ρ x<br />
Transconductance<br />
amplifier<br />
x<br />
392
11-13 May 2011, Aix-en-Provence, France<br />
V<br />
x<br />
= nx<br />
Fx<br />
(2)<br />
i n x<br />
x<br />
= (3)<br />
n x is the piezoelectric conversion factor, V x , the resonator<br />
excitation voltage and i x , the motional current.<br />
C 0 is the inter-electrode capacitance formed by the<br />
electrodes with quartz as dielectric. Finally, the output current<br />
in the resonator is:<br />
i<br />
tot<br />
x<br />
dVx<br />
= C 0<br />
+ nx<br />
x<br />
(4)<br />
dt<br />
+<br />
V +<br />
VerilogA OUT<br />
-<br />
Rf<br />
V −<br />
V<br />
Fig. 4. transconductance amplifier<br />
The resonator block in Verilog-A is built with equations (1),<br />
(2), (3) and (4). They give the physical description. The code<br />
is:<br />
module resonator(a,b);<br />
input a; output b;<br />
electrical a,b;<br />
kinematic Z,Fx;<br />
kinematic_v Vc;<br />
parameter real k=1,4e3 from [0:inf);<br />
parameter real m=1e-8 from [0:inf);<br />
parameter real ρ=3,1e-7 from [0:inf);<br />
parameter real n=2,8e-3 from [0:inf);<br />
analog<br />
begin<br />
Vel(Vc)
11-13 May 2011, Aix-en-Provence, France<br />
R1<br />
V in<br />
+<br />
V +<br />
VerilogA<br />
OUT<br />
-<br />
V −<br />
V<br />
Transconductance<br />
amplifier output<br />
R2<br />
Rf<br />
Resonator<br />
input<br />
C1<br />
Fig. 6. Integrator block<br />
From now, these blocks stand as the MEMS oscillator model,<br />
which consists of the association of the resonantor and the<br />
oscillator circuit. Then transient simulations can be performed<br />
and compared to experimental measurements to validate the<br />
model.<br />
Fig. 7. Starting oscillator<br />
This accelerometer being modelled by equation (1), the<br />
theoretical resonance frequency f 0 is given by (6) and (7)), and<br />
should be 59,5kHz using values from Table 1:<br />
k x<br />
2<br />
ω<br />
0 = with ω<br />
0<br />
= 2π<br />
f0<br />
(6), (7)<br />
m<br />
III.<br />
MODEL SIMULATIONS<br />
One finds the same value from direct measurement on the VIA<br />
and from temporal simulations.<br />
A. Transient simulation<br />
The Cadence transient simulation allows solving the current<br />
and voltage in each node of the MEMS oscillator model. To<br />
run the simulations, it becomes necessary to input a Dirac<br />
spike on the transconductance amplifier. Indeed, this type of<br />
self-sustained oscillator starts from the white noise flowing in<br />
the loop. As the loop gain is greater than one at the resonator<br />
eigenfrequency f 0 , it is the only frequency amplified.<br />
17µs<br />
Transconductance<br />
amplifier output<br />
In order to validate the model, simulations are performed<br />
taking into account the parameters of the VIA Vibrating Beam<br />
Accelerometer developed at ONERA [4] shown in table 1.<br />
TABLE I<br />
VIA accelerometer parameters<br />
m (kg) ρ x (kg/s) k x (kg/s²)<br />
1.10 -8 3,1.10 -7 1,4.10 3<br />
These parameters have been extracted from measurements on<br />
existing VIA devices.<br />
The starting oscillator is shown in figure 7:<br />
Resonator<br />
input<br />
Fig. 7. Transient simulation<br />
Indeed, the transient simulation shows about a 17µs period i.e.<br />
59kHz frequency. Moreover, the transconductance amplifier<br />
has multiplied the amplitude by 10 and is limited by the<br />
Automatic Gain Control.<br />
B. Phase noise simulations<br />
Once the resonance frequency is found, the Virtuoso<br />
SpectreRF Simulator can simulate the loop phase noise.<br />
The interesting point of this model resides in the control<br />
possibility of the system parameters on each block. Thus, it is<br />
possible to obtain the evolution of the phase noise as a<br />
function of the resonator physical parameters. Phase noise<br />
simulation is performed using the PSS module (Periodic<br />
Steaty-State).<br />
The module advantage is the simulation time. Indeed, it avoids<br />
the long time temporal simulation.<br />
394
The simulation output phase noise has been compared with the<br />
theoretical phase noise from Leeson effect.<br />
11-13 May 2011, Aix-en-Provence, France<br />
One can calculate the phase noise of such an oscillator with<br />
the Leeson effect [5], and examine the phase noise variations<br />
with the resonator physical parameters. The Leeson equation<br />
between the oscillator and the feedback amplifier [6] phase<br />
spectral densities follows relation (8):<br />
Leeson frequency<br />
17Hz<br />
1 ν<br />
0 2<br />
Sϕ<br />
( f ) = [1 + ( ) ] S ( f )<br />
2 φ<br />
(8)<br />
f 2Q<br />
S ϕ<br />
( f ) is the phase spectral density of the oscillator<br />
S φ<br />
( f ) is the phase spectral density of the amplifier<br />
ν<br />
0 is the resonance frequency and Q is the quality factor of<br />
the resonator.<br />
f is the offset frequency from the carrier.<br />
The cut-off frequency, called Leeson’s frequency, is:<br />
ω<br />
f L<br />
= 0<br />
(9)<br />
2Q<br />
Simulations are run taking the physical quantities of the VIA<br />
(table 1) into account. The quality factor of the characteristic<br />
equation (1) is:<br />
Fig. 8. phase noise output<br />
This Leeson's model theoretical value matches the verilog-A<br />
model simulation described above. Leeson’s frequency varies<br />
only with the damping coefficient and the mass.<br />
In the next simulations, only the mass of the<br />
resonator is changed. The theoretical Leeson's frequency is<br />
proportional to m -1 . We have previously shown that the<br />
Leeson's frequency was 15.5 Hz with a mass of 1.10 -8 kg. For<br />
1.10 -9 kg and 1.10 -10 kg, we have respectively frequencies of<br />
155Hz and 1.55 kHz, as verified by the following simulation<br />
on figure 9:<br />
ω m<br />
Q = 0<br />
(10)<br />
ρ<br />
x<br />
The Leeson’s frequency is:<br />
ρ<br />
x<br />
f<br />
L<br />
= (11)<br />
2m<br />
f L<br />
= 15. 5Hz<br />
Phase noise analysis is performed with Virtuoso simulator and<br />
the output signal is selected to create the phase noise output. It<br />
becomes possible to read the Leeson’s frequency (figure 8).<br />
Fig. 9. phase noise output depends on the mass<br />
C. Influence of the transconductance amplifier noise<br />
The white noise and flicker noise can be added to this<br />
model and particularly to the transconductance amplifier<br />
block.<br />
1. White noise<br />
White noise function is added to the the transconductance<br />
amplifier plus pin. Therefore, we add a VerilogA block. A<br />
Cadence special function is used : white_noise().<br />
395
The VerilogA block code is :<br />
11-13 May 2011, Aix-en-Provence, France<br />
module generator_noise(a,b)<br />
input a; output b;<br />
electrical a, b;<br />
parameter real noise_generator=1e-10 from [0:inf);<br />
analog<br />
begin<br />
V(a,b)
The oscillator output phase noise calculated from the<br />
Leeson formula is then:<br />
S<br />
2 2<br />
1 ν<br />
0 2 2en<br />
Rm<br />
f ) = 10log([1 ( ) ] )<br />
2<br />
2<br />
(18)<br />
f 2Q<br />
V R<br />
ϕ<br />
( +<br />
2<br />
x f<br />
11-13 May 2011, Aix-en-Provence, France<br />
2. Flicker noise<br />
The flicker_noise function is also added into the<br />
transconductance amplifier plus pin. Flicker noise is directly<br />
visible on the output phase noise. The results are consistent<br />
with the Leeson effect.<br />
IV. CONCLUSION AND FUTURE WORK<br />
As a fundamental result, a MEMS oscillator model has been<br />
developed using the multiphysics Verilog-A language. This<br />
model includes the mechanical and piezoelectric equations,<br />
and the electronic circuitry. This description main asset lies in<br />
allowing the study of critical parameters influencing<br />
performances and especially the phase noise analysis.<br />
This model will be used to compare the influence on phase<br />
noise performances of different types of transduction (optical<br />
and electrostatic…) and other types of oscillator circuits as<br />
PLL oscillator circuit [7]. These values will also be compared<br />
with experimental results obtained with MEMS oscillators.<br />
REFERENCES<br />
[1] O. Le Traon, “The VIA Vibrating Beam Accelerometer: Concept and<br />
Performances”, Proceedings of the PLAN Symposium, 1998.<br />
[2] O. Le Traon, D. Janiaud, M. Pernice, S. Masson, S. Muller, J-Y<br />
Tridera, “A New Quartz Monolithic Differential Vibrating Beam<br />
Accelerometer”, Position, Location, And Navigation Symposium, pp.6-15,<br />
2006<br />
[3] Loreto Rodríguez-Pardo, “Sensitivity, Noise, and Resolution<br />
in QCM Sensors in Liquid Media”, IEEE Sensors Journal, vol.5, n°6, 2005<br />
[4] O. Le Traon, “The VIA Vibrating Beam Accelerometer: Concept and<br />
Performances”, Proceedings of the PLAN Symposium, 1998.<br />
[5] D.B. Leeson “A simple model of feedback oscillator noise spectrum”,<br />
Proceedings of the IEEE, vol. 54, issue 2, p. 329-330, 1966.<br />
[6] E. Rubiola, R. Brendel “A generalization of the Leeson effect”,<br />
arXiv:1004.5539 [physics.ins-det], April 2010.<br />
[7] R. Levy, D. Janiaud, O. Le Traon, S. Muller, JP. Gilles, G. Raynaud,<br />
“A new analog oscillator electronics applied to a piezoelectric<br />
vibrating gyro”, Proceedings of the IEEE Frequency Control Symposium<br />
pp.326-329, 2004<br />
Author Biography:<br />
Guillaume Papin is engineer from the ENSMM since 2010. He is currently a<br />
PhD student at ONERA working on the development of multi-physic models<br />
to improve phase noise performances of vibrating MEMS sensors.<br />
397
Author Index<br />
Abi-Saab D. 81<br />
Aimez Vincent 300<br />
Aini Md Ralib Aliza 85<br />
Akarvardar K. 348<br />
Allen David M. 29<br />
Angelescu D. 81<br />
Anis Nurashikin Nordin 18, 85<br />
Ardila G. 348<br />
Asgari M.B. 3, 103<br />
Ayon Arturo 72, 137, 185<br />
Azaïs F. 14<br />
Ballet Jérôme 249<br />
Bancaud Aurélien 241<br />
Bartolucci Giancarlo 263<br />
Basset P. 81<br />
Begbie M. 253<br />
Bendali A. 378<br />
Bergonzo P. 378<br />
Berthillier Marc 187<br />
Boisseau Sébastien 386<br />
Bongrain A. 378<br />
Bornoff Robin 324<br />
Bosch Robert 1<br />
Bossuyt Remy 249<br />
Bouchaud Jérémie 211<br />
Bourouina T. 81<br />
Boutaud Bertrand 386<br />
Bouteloup G. 348<br />
Brisard Thierry 211<br />
Brown Keith 41<br />
Brusa Eugenio 356<br />
Buser R. 128<br />
Camon Henri 241, 249<br />
Cano Jean-Paul 249<br />
Chabanov Andrey 72, 185<br />
Chaehoi A. 253<br />
Chaillout Jean-Jacques 386<br />
Chang Ho-Hsien 352<br />
Chang Pei Hua 151<br />
Chang Ting-Chou 245, 305<br />
Chang Tung-Yu 294<br />
Chao Ching-Kong 176<br />
Charlot Benoît 134<br />
Charrette Paul G. 300<br />
Chatani Keisuke 338<br />
Chau Lai-Kwan 245, 305<br />
Chen Chien-Hsing 245, 305<br />
Chen Chonglin 72, 185<br />
Chen Chun Huei 170<br />
Chen Guan-Lan 159<br />
Chen Jyh Jian 170<br />
Chen Taco 329<br />
Cheng Ya-Chi 159<br />
Chuang Cheng-Hsin 122<br />
Collins Greg 72<br />
Combette Philippe 134<br />
Conédéra Véronique 249<br />
Corigliano Alberto 53<br />
Costello Suzanne 206, 208<br />
Dalmolin Renzo 386<br />
Dany Maximilien 29<br />
Dauksevicius Rolanas 164<br />
De Angelis Giorgio 263<br />
De Pasquale Giorgio 97, 356<br />
Delabie Christophe 193<br />
Desmulliez Marc P.Y. 41, 110, 206, 207<br />
Deterre Martin 386<br />
Dinglreiter Heinz 277<br />
Dovhij Victor 184<br />
Drysdale D. 35<br />
Dufour-Gergam Elisabeth 386<br />
Dumas Norbert 14, 315, 320<br />
Eftekhar Azam Saeed 53<br />
Elam David 72, 185<br />
Esteves Josué 309<br />
Exertier Anne 193<br />
Fakri Abdenasser 193<br />
Fakri-Bouchet Latifa 193<br />
Flourens F. 81<br />
Flynn David 41<br />
Fujimori Tsukasa 237<br />
Fujimoto Akifumi 333<br />
Fujimoto Jun 227<br />
Garcia Ronald 151<br />
Garraud Alexandra 134<br />
Ghisi Aldo 53<br />
Giani Alain 134<br />
Goto Yasushi 237<br />
Gué Anne-Marie 249<br />
Gyenge Oliver 46<br />
Hacine Souha 320<br />
Hackworth R. 137<br />
Hansen Ulli 46<br />
Hauck Karin 46<br />
398
Heeb P. 128<br />
Heilig Markus 283<br />
Hinchet R. 348<br />
Ho Jing-Yu 23<br />
Hodossy Sandor 324<br />
Hoeffmann Janpeter 200<br />
Holota Victor 184<br />
Hosoi Atsushi 333<br />
Howe R. T. 348<br />
Hsiao Fei-Bin 362<br />
Hsiao Ju-Hsiu 156<br />
Hsu Jen-Sung 294<br />
Hsu Shan-Shan 294<br />
Huang Cheng-Chun 372<br />
Huang Yi-Hsuan 122<br />
Hui Hui 59<br />
Hung Ling-Hsuan 366<br />
Iamoni Sonia 97<br />
Ikehara Tsuyoshi 338<br />
Imamoto Hiroshi 237<br />
Isagawa Kohei 231<br />
Ishida Takao 180<br />
Itao Kiyoshi 212<br />
Itoh Toshihiro 142, 217, 221, 227, 231, 237<br />
Jen Chun-Ping 156, 352, 362<br />
Ju Yang 333<br />
Kaminaga Susumu 288<br />
Karam Jean Michel 211<br />
Kaufmann Jens G. 41<br />
Kay Robert W. 110<br />
Khan Malek Chantal 278<br />
Khan Sheroz 18<br />
Knechtel Roy 106, 344<br />
Knoll Thorsten 273<br />
Kobayashi Takeshi 142, 221, 231<br />
Kogut Igor 184<br />
Köhler Andreas 64<br />
Kolew Alexander 277, 283<br />
Kotha R. 137<br />
Krebs Annabel 273<br />
Kulvietis Genadijus 164<br />
Kurata, Hideaki 237<br />
Lardiès Joseph 187<br />
Latorre Laurent 315, 320<br />
Lee Chai-Yu 245, 305<br />
Lee Chungda 90<br />
Lee Shin-Li 122<br />
Lee Yung-Chun 362<br />
Lefeuvre Elie 258, 386<br />
Leib Juergen 46<br />
Lenczner Michel 59<br />
Leprince-Wang Y. 81<br />
Levy R. 392<br />
Lin Ming-Je 268<br />
Lin Ming-Tzer 159, 329<br />
Lin Tsung-Hung 176<br />
Lissorgues G. 378, 392<br />
Liu Ming 72<br />
Liu Yao-Lung 366<br />
Loisel Pierre 134<br />
Lu Guo-Neng 300<br />
Lu Jian 217<br />
Lucibello Andrea 263<br />
Ma Chunrui 72<br />
Maeda Ryutaro 180, 217, 227, 231, 237, 338<br />
Mailly F. 14<br />
Mailly Frederick 320<br />
Malak M. 81<br />
Marcelli Romolo 263<br />
Marek Jiri 1<br />
Mariani Stefano 53<br />
Markus Heilig 277<br />
Maroufi M. 3<br />
Martel Stéphane 300<br />
Martell Steven R. 206, 210<br />
Martincic Emile 258<br />
Marty F. 81<br />
Masmoudi M. 14<br />
Maus Simon 46<br />
Maxwell R. 137<br />
Mazenq Laurent 249<br />
Megherbi Souhil 258<br />
Metwally Khaled 278<br />
Mezghani B. 14<br />
Mias Solon 241<br />
Michael Steffen 106, 344<br />
Michaelsen J. A. 116<br />
Milasauskaite Ieva 164<br />
Mirabbaszadeh K. 103<br />
Miyake Koji 142<br />
Montès L. 348<br />
Moriera J. R. 137<br />
Moulin Johan 258<br />
MuÅNnch Daniel 283<br />
Nayebi P. 103<br />
Nemashkalo Anastasiia 185<br />
Neylon Sean 211<br />
Nguyen K.N 81<br />
Niessner Martin 8<br />
Nouet Pascal 14, 315, 320<br />
399
Nussbaum Dominic 273<br />
O’Hara T. 35<br />
Okada Hironao 221<br />
Ostasevicius Vytautas 164<br />
Othman Raihan 85<br />
Ouellet Luc 300<br />
Papin G. 392<br />
Parsa R. 348<br />
Paul Steffen 200<br />
Peters-Drolshagen Dagmar 200<br />
Picaud S. 378<br />
Pistor Jonas 200<br />
Pittet Patrick 300<br />
Poulichet Patrick 193, 392<br />
Proietti Emanuela 263<br />
Ramstad J. E. 116<br />
Redondo Roxana 29<br />
Rehder Gustavo 309<br />
Reinert Wolfgang 206, 209<br />
Reitz Sven 64<br />
Rekik A.A. 14<br />
Rencz Marta 324<br />
Ress Sandor 324<br />
Rezaie A.H. 3<br />
Richalot E. 81<br />
Richard Charles 300<br />
Robert Laurent 278<br />
Rousseau Lionel 193, 378<br />
Rufer Libor 309<br />
S. Mousavi M. Mehdi 356<br />
Salleh Hanim 85<br />
Salut Roland 278<br />
Sarkany Zoltan 324<br />
Schäffel Christoph 106, 344<br />
Scherner S. 253<br />
Schneider Marc 277<br />
Schneider Peter 64<br />
Schrag Gabriele 8<br />
Scorsone E. 378<br />
Shahosseini Iman 258<br />
Shamshirsaz Mahnaz 3, 103<br />
Shibayama Nobuhisa 142<br />
Shie Shian Ruei 170<br />
Shiga Shouhei 180<br />
Shih Hsin-Yuan 362<br />
Shih Wei-Hung 245, 305<br />
Shih Wen-Pin 372<br />
Sikora Karsten 283<br />
Silva Emanuel 72<br />
Simoni Barbara 53<br />
Soeraasen O. 116<br />
Somà Aurelio 97, 356<br />
Strzhemechny Yuri 185<br />
Takamatsu Seiichi 142<br />
Tang Jaw-Luen 245, 305<br />
Toepper Michael 46<br />
Tong Chi-Jia 159<br />
Trigona Carlo 315<br />
Tsai Ching-Hsiu 366<br />
Tsai Yu-Tze 381<br />
Tschanun W. 128<br />
Ueda T. 148<br />
Vass-Varnai Andras 324<br />
Velten Thomas 273<br />
Voigt Sebastian 106, 344<br />
Wachutka Gerhard 8<br />
Wang C.H. 35<br />
Wang Dong F. 180, 231, 338<br />
Wang Gou-Jen 23, 381<br />
Wang Jian-Neng 245, 305<br />
Wang Shau-Chun 245, 305<br />
Wang Song Hao 151<br />
Wang Yu-Chi 372<br />
Wang Yu-Ting 159<br />
Weiland D. 253<br />
Weng Feng-Tsai 268<br />
Wilhelm Stefan 110<br />
Wisland D. 116<br />
Wong Philip H.-S. 348<br />
Worgull Matthias 277, 283<br />
Woytasik Marion 258<br />
Wu Chia-Che 366<br />
Wu Ming-Dao 372<br />
Wu Wei-Te 245, 305<br />
Ya Ma Li 18<br />
Yang Hsiharng 176, 268, 294<br />
Yu Jyh-Cheng 90, 197<br />
Yvert B. 378<br />
Zaminpeyma E. 103<br />
Zhang Lan 333<br />
Zhang Yi 217, 221<br />
Zihajehzadeh Sh. 3<br />
Zimin Y. 148<br />
Zoschke Kai 46<br />
400