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CRITERI DI SCELTA DEL MATERIALE
PER L’ALLEGGERIMENTO DI VETTURE
SPORTIVE AD ALTE PRESTAZIONI
P. Veronesi, A. Pivetti, A. Baldini, M. Loiacono, G. Poli
Per le vetture di lusso ad alte prestazioni, il primo elemento di competitività è dato dalle prestazioni dinamiche
della vettura, ragion per cui il fattore peso sta assumendo nel tempo una rilevanza crescente. L’introduzione
di normative sempre più severe in ambito di emissioni, strettamente correlate al consumo della
vettura ha inoltre indotto i progettisti a spostare la propria attenzione non più sulla potenza pura ma sul
rapporto potenza/peso. In questo ambito, è stato studiato l’alleggerimento di una vettura sportiva ad
alte prestazioni, in particolare di una Lamborghini Murciélago, andando a proporre nuovi materiali per
la realizzazione di particolari del gruppo sospensivo. Ottimizzando la scelta del materiale, è possibile ridurre
il peso, rispetto alla soluzione attuale, del 30-35% relativamente alla molla sospensione anteriore, del 50-70%
relativamente alla barra antirollio posteriore, nonché alleggerire il braccio anteriore inferiore dal 3 al
30%, dipendentemente dallo stato tensionale del componente, impiegando opportunamente acciai basso legati,
leghe di alluminio e leghe di titanio.
PAROLE CHIAVE: acciaio, alluminio e leghe, mat. compositi, titanio e leghe, selezione materiali
INTRODUZIONE
Negli ultimi 20 anni la progettazione e realizzazione di autoveicoli
ha subito notevoli cambiamenti per poter venire incontro
a nuove e differenti esigenze da parte degli utenti.
Innanzitutto sono cambiati gli standard per quanto riguarda
confort e abitabilità: lo spazio a disposizione dei passeggeri
è aumentato ad ogni nuova generazione di veicoli e con esso le
dimensioni delle vetture.
Parallelamente a questo, la richiesta di vetture sempre più sicure
per gli occupanti ma anche per i pedoni ha reso necessario
un aumento delle dimensioni medie delle vetture e di conseguenza
un aumento di peso delle stesse [1].
Negli ultimi anni ci si è però resi conto che questo trend
non poteva proseguire poiché il miglioramento tecnologico
di propulsori e combustibili, per quanto notevole, non sarebbe
stato in grado di compensare l’aumento di peso delle vetture,
soprattutto considerando le richieste sempre maggiori del mercato
in termini di prestazioni dinamiche (migliore accelera-
P. Veronesi, G. Poli
Dipartimento di Ingegneria dei Materiali e dell’Ambiente,
Via Vignolese 905, 4110 Modena - Italy
A. Baldini, M. Loiacono
Dipartimento di Ingegneria Meccanica,
Via Vignolese 905, 4110 Modena - Italy
A. Pivetti
Lamborghini, Via Modena 12, S. Agata Bolognese - Bologna - Italy
zione, ripresa, maneggevolezza, minori consumi/emissioni). Il
peso di una vettura infatti può influenzare numerosi parametri
della progettazione oltre che le prestazioni del veicolo stesso:
- veicoli più pesanti richiedono maggiori potenze per ottenere
prestazioni analoghe a mezzi più leggeri; questo
comporta solitamente un maggiore consumo di carburante
e quindi maggiori emissioni di CO 2 .
- Un peso maggiore significa anche maggiori inerzie e quindi
minor prontezza di risposta ai comandi e minor piacere di guida.
- In caso di incidente una maggiore massa implica una maggiore
energia cinetica da dissipare e quindi richiede delle prestazioni
di resistenza strutturale maggiori da parte del veicolo
In questo ambito è opportuno inoltre ricordare i nuovi limiti
di emissioni di CO 2 (130 g/km) imposti dalla Comunità Europea
che entreranno in vigore nel 2012: per riuscire a rientrare
in tali limiti sarà fondamentale lavorare su una riduzione dei
pesi delle vetture, perché la sola efficienza dei motori non sarà
assolutamente sufficiente. I problemi maggiori li incontreranno
sicuramente le vetture di dimensione medio-grande, nonché
quelle ad alte prestazioni, che sono l’oggetto di studio del
presente lavoro.
Nel settore delle vetture di lusso ad alte prestazioni il primo
elemento di competitività è dato dalle prestazioni dinamiche
della vettura; in questo settore, il fattore peso sta assumendo nel
tempo una rilevanza sempre maggiore. L’introduzione di normative
sempre più severe in ambito di emissioni, strettamente
correlate al consumo della vettura ha infatti indotto i progettisti
a spostare la propria attenzione non più sulla potenza pura
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a
b
s
Fig. 2
diagrammi di scelta per tirante rigido di massa
minima e per trave rigida di massa minima, con indici
dei materiali rapportati al materiale attualmente in uso
per il componente.
Selection diagrams for a stiff tie having minimum mass
and for stiff beam having minimum mass; materials
indexes are normalized to the properties of the material
currently used for the component.
Materiale
AISI 94B30 (G94301), temprato e rinvenuto a 205°C
AISI 4042 (DIN 42MnMo7), temprato e rinv. a 205°C
AISI 4340 (UNI 40NiCrMo7), normalizzato
Al-Lega da getto S520.0
Al-2026 T3
Al-2024 T3
Al-6082 T4 (anche forgiato)
Al-7020 T5
Al-7033 T6
Selezione materiali
0 M /M 1 1
1.02
1.00
1.10
1.06
1.00
0.99
0.99
0.96
0.96
s
Fig. 3
Costo per unità di massa dei materiali – database
CES-Edupack 2005.
Suspension system: front lower arm.
ni alternative in grado di conferire migliori prestazioni al componente,
relativamente all’obiettivo individuato e ai vincoli agenti.
Un fattore inizialmente non considerato, ma che può giocare un
ruolo determinante nella scelta del materiale, è il costo. In Fig. 3
è rappresentato il diagramma relativo al costo per unità di massa
di differenti tipologie di materiali, con indicate le barre relative
ad acciai, leghe di alluminio e titanio.
Dall’analisi dei grafici di Fig. 2 e 3, si può notare come dal punto
di vista della rigidezza specifica (M1), si ha una famiglia di materiali
in grado di avere prestazioni superiori all’acciaio: le leghe di
titanio; tuttavia i miglioramenti sono possibili ma in misura
ridotta (pochi punti percentuali) e a costi nettamente maggiori.
Non sembra quindi opportuno proseguire verso una
scelta del genere, considerando i limitati vantaggi in rapporti
ai costi. Dal punto di vista della resistenza specifica (M2),
invece, si possono ottenere miglioramenti considerevoli.
Relativamente alle prestazioni di rigidezza flessionale, i grafici
mostrano che vi sono soluzioni in grado di fornire buoni miglioramenti:
in particolare le leghe di alluminio sembrano
offrire le migliori soluzioni.
La rosa di candidati per l’applicazione, si restringe pertanto ai
materiali riportati in Tab. 1, con i relativi indici di prestazione
riferiti al materiale attualmente in uso (apice “0”).
Informazioni di supporto e scelta del materiale
Le informazioni ricavate dagli indici di prestazione hanno
0 M /M 2 2
5.50
3.59
2.60
1.92
2.77
2.50
1.53
2.15
3.07
0 M /M 3 3
1.00
1.00
1.00
1.79
1.68
1.67
1.69
1.64
1.67
0 M /M 4 4
2.3
2.3
1.9
2.2
2.8
2.6
1.9
2.3
3.0
s
Tab. 1
indici
di prestazione
dei materiali
candidati.
Materials
indexes for
candidate
materials.
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molla in seguito a variazioni della quota z del contatto ruotasuolo.
In ogni caso, per qualsiasi scuotimento della sospensione
la molla rimane sempre sollecitata a compressione. Il filo della molla
è avvolto a spirale attorno ad un cilindro fittizio di diametro
costante, denominato diametro di avvolgimento. Come già
accennato, il componente deve rispettare determinati ingombri
per consentirne il montaggio: in particolare un parametro che non
può essere modificato è il diametro interno della molla. Tale quota
è infatti limitata dalla necessità di montare all’interno della molla
l’ammortizzatore. Un’altra specifica tecnica che non deve essere
modificata è la rigidezza; tale parametro infatti influenza il comportamento
dinamico della vettura e viene quindi stabilito in fase
progettuale per ottenere le prestazioni dinamiche volute. Anche
questo parametro verrà quindi mantenuto costante. Per quanto
riguarda la lunghezza della molla, si cercherà di mantenerla il più
possibile simile a quella di origine, in modo da ridurre al minimo
eventuali modifiche su componenti necessari al fissaggio
della molla, come per esempio i piattelli; ovviamente esistono
limitazioni legate all’ingombro longitudinale ma una variazione
di qualche millimetro può essere concessa. Va osservato che si
tratta di una molla con passo e diametro medio di avvolgimento
costanti; la sua caratteristica forza-spostamento sarà quindi
lineare. Il dimensionamento ha portato ai valori riportati in
Tab. 2, espressi in termini di variazione percentuale per ragioni
di riservatezza.
ANALISI ED ALLEGGERIMENTO BARRA ANTIROLLIO
La barra antirollio svolge un compito di ausilio al sistema sospensivo
solamente in situazioni in cui si ha trasferimento di carico.
Gli estremi della barra sono fissati ai braccio superiore del
sistema sospensivo: in tal modo quando si verifica un trasferimento
di carico tra le due ruote dell’assale, gli scuotimenti verticali di
verso opposto tra le due ruote generano un momento torcente sulla
barra, che quindi viene sollecitata a torsione. Il componente è mostrato
in Fig. 5.
Un parametro fondamentale che sintetizza il comportamento
della barra stessa è la sua rigidezza torsionale: più la barra è
rigida maggiore sarà la resistenza che essa opporrà e di conseguenza
minori saranno gli scuotimenti del sistema sospensivo
dell’assale.
Nel dimensionare il componente quindi l’aspetto principale da
s
Tab. 2
Dimensionamento della molla in lega di titanio e
relativa riduzione di peso.
Dimensioning of titanium alloy spring and consequent
weight reduction.
Selezione materiali
s
Fig. 5
Barra antirollio.
Anti-roll bar.
tenere i considerazione sarà la rigidezza torsionale che si vorrà
ottenere; come si vedrà in seguito tale proprietà dipende sia da
quote dimensionali (diametro della barra, lunghezza del braccio di
torsione) che dalle proprietà del materiale.
Traduzione dei requisiti di progetto
Oltre alle funzioni esposte al paragrafo precedente, un parametro
importante per la selezione del materiale è la massima
temperatura di esercizio: la barra infatti è situata in prossimità del
motore e dell’impianto di scarico della vettura, e si troverà quindi
a lavorare ad una temperatura leggermente superiore a quella ambiente.
Come limite inferiore è stata imposta una temperatura di
lavoro di 70° C.
Un altro aspetto da tenere in considerazione è il comportamento
del materiale in caso di frattura: la caratteristica desiderata, ovviamente,
è che in caso di cedimenti il materiale non ceda di schianto
ma resista il più possibile alla propagazione della cricca. Il parametro
solitamente utilizzato per definire il comportamento di un
materiale in caso di cedimento è la tenacità a frattura. La
traduzione dei requisiti di progetto è la seguente:
- Funzione: barra di torsione
- Vincoli: rigidezza torsionale, resistenza a rottura, resistenza a fatica
e ad agenti atmosferici, vincoli geometrici di ingombro; temperatura
di impiego superiore a 70°C; tenacità a frattura superiore a
15 MPa m 1/2 .
- Obiettivo: ottenere rigidezza di progetto con minimo peso
- Variabile libera: materiale da utilizzare, dimensione della sezione
resistente
Screening e classificazione: calcolo indici di prestazione e individuazione
materiali alternativi
L’indice di prestazione per la rigidezza torsionale e per la resistenza
torsionale sono i seguenti [5]:
(7)
(8)
Il relativo diagramma di scelta, nella zona di interesse, è riportato
in Fig. 6.
In base alle considerazioni effettuate in precedenza relativamente
all’utilizzo di componenti in materiale composito, la scelta più
ragionevole sembra quella delle leghe di alluminio da trattamento
termico e/o deformazione plastica. In base all’analisi del grafico le
leghe migliori sono quelle delle serie 5000 e 2000. In particolare
le leghe migliori sono quelle della serie 5000, che presentano
caratteristiche di resistenza sufficienti e una densità leggermente
inferiore rispetto alle altre serie, massimizzando così l’indice di
prestazione. Una scelta opportuna potrebbero essere le leghe 5052
e 5086. Tuttavia, semplificando la geometria del componente,
la metallurgia italiana >> marzo 2009 5

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Metalli leggeri
INNOVATIVE TECHNOLOGIES IN
MOULD RELEASE AGENTS
G.Natesh, A. Colori
The drive for improved fuel efficiencies in the automobile industry has led to continuing growth in aluminium
die casting as manufacturers strive to reduce the weight of automobiles by replacing steel with light metal components.
Larger and more complex parts are being cast and this has set new challenges to die casters in their
quest for improved quality and productivity. The paper examines the impact of these trends on die lubrication
and discusses an innovative lubrication technology that has evolved to satisfy these requirements.
KEYWORDS: : mould release agent, die-casting, high temperature, automotive, die lubricant, solder protection,
leidenfrost effect
High Pressure die-casting is a very popular process for making
complex mechanical parts out of light metals like aluminium
and magnesium alloys. It is capable of rapidly producing parts
with high dimensional accuracy. High pressure die-casting
grew along with the growth of the automobile industry, where
the demands of assembly line manufacture spurred the
demand for a quick reliable way to make components. With
the growth of JIT manufacturing, the automobile industry still
continues to be the dominant user of high-pressure die cast
parts. Other end uses for die casting include recreational vehicles,
power tools, electrical machinery, electronic components
and house-ware. The rapid growth of the world economy has
spurred a demand for all of these products and the European
die cast industry is gearing up to meet this demand.
The rising cost of fuel and increasingly stringent environment
and fuel performance regulations are forcing the auto industry
to seek novel ways to achieve these goals. Weight reduction
of vehicles is a key step to reducing fuel consumption, so
the industry is actively looking at replacing steel components
with aluminium and magnesium castings. With constant innovation
in aluminium alloys and casting technology, improved
strength and other properties are being engineered, that allows
bigger and more complex parts to be die cast. Engine blocks,
instrument panels and complete door frames are just some of
the examples of aluminium components now being produced
by die-casting. This has led to a trend towards bigger die-cast
machines and larger shot weights.
The complexity of these large parts makes it difficult to design
internal cooling to adequately cool all parts of the die uniformly.
A natural consequence of this is that die surface temperatures
have increased. Previously, the die surface temperatures
before spray used to range between 250°C to 350°C. With the
estimenti di allumina, corrosi
one, leghe di alluminio, LEIS, EIS, ENA
mina, corrosi
one, leghe di alluminio, LEI
mina, corrosi
one, leghe di alluminio, LEI
large components, the maximum temperature can be as high
as 400°C while the cooler portions of the die may be as low as
200°C.
This leads to the development of localized hot spots which, in
turn, create solder problems. This places a greater dependence
on the die lubricant to provide cooling for the die surface. Yet
the higher temperatures encountered before the spray make
this difficult to do because of the Leidenfrost effect. This requires
greater quantities of die lubricant to be sprayed, which
increase cycle times and costs.
The Leidenfrost phenomenon is well known to die casters.
When water is sprayed on to a hot surface, which is at a temperature
well above the boiling point of water, it is unable to
make contact with the metal surface. Instead, the drops of water
float on a cushion of water vapor and thus are unable to
wet the surface (Fig. 1). Die lubricant active materials are therefore,
unable to be laid down on the die surface. The highest
temperature at which water, or a water based die lubricant,
can contact the metal surface is known as the Leidenfrost temperature.
Our research focused on two separate approaches to develop
high performance die lubricants. The first was to try and in-
s
Fig. 1
Schematic rendering of the Leidenfrost phenomenon.
Rappresentazione grafica del fenomeno di Leidenfrost.
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a
b
s
Fig. 5
a) Old product (after 8 hours); b) New Product
(after 8 hours).
a) Vecchia tecnologia (dopo 8 ore); b) Nuova tecnologia
(dopo 8 ore).
The first case is from a North American die caster making
engine blocks with steel cylinder inserts on a 3500 T Ube®
machine with a total cycle time of less than 120 seconds. They
were getting good solder protection, with low overspray and
in-cavity buildup with a conventional die lubricant. When
they started casting a new engine design, they noticed solder
formation near the water jacket area on the part. This required
them to do die polishing for about 30 minutes once every
8 hours. Running at richer concentrations did not help, giving
buildup that also needed to be polished. Increasing the
spray duration also had a major impact on productivity.
We took thermal images of the die before and after spray to
monitor temperature profiles and spray distribution. The temperature
ranged from 450°F to 750°F (232°C to 399°C) before
spray on the ejector die. We also observed that the previous
product was not covering the problem area adequately particularly
at the high temperature zones. The new Safety-Lube®
product could wet the hot surface earlier providing better co-
Metalli leggeri
s
Fig. 6
Thermal image of die before spray.
Immagine termica dello stampo prima della spruzzatura.
s
Fig. 7
Thermal image of die after spray (old product).
Immagine termica dello stampo dopo la spruzzatura (vecchia
tecnologia).
verage and could rapidly form an adequate lubricating film
at the high temperatures seen on the die. Fig. 5 shows the
dramatic reduction in solder. The improved performance
from the Safety-Lube® product eliminated the need to polish
every shift and reduced the cleaning time by 50%.
The second example is from a European die caster making
automotive components, who were extremely concerned
about the long cycle times needed to make a particular casting.
They also had problems with porosity, soldering and
in-cavity build-up which led to poor yields and productivity.
Investigation of the problem revealed a clear pattern. Fig. 6
shows that the die temperatures before spray ranged from
417°C to 230°C across the face of the die.
The incumbent product was unable to form adequate protective
film against solder, therefore a long spray time was needed.
However, this caused some areas of the die to be cooled
excessively, leading to in-cavity build and porosity. Reducing
the spray time gave solder, and in both cases downtime was
needed to do die polishing. As can be seen in Fig. 7, the typical
die temperatures after spray with the conventional product
ranged from 250°C to 160°C.
From this analysis, it was clear that we needed to form good
la metallurgia italiana >> marzo 2009 3

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Memorie >>
Trattamenti superficiali
MODERN THERMAL ELECTRON BEAM
PROCESSES – RESEARCH RESULTS
AND INDUSTRIAL APPLICATION
R. Zenker
Thermal electron beam (EB) technologies are becoming more and more attractive especially
because they are ecologically friendly and energy saving on the one hand and highly precise, excellently
controllable and highly productive on the other hand.
Using three-dimensional energy transfer fields, the interaction conditions between the EB and the surface
of the material, the conditions of the heat conduction in the material, the geometry of the part, and the load
conditions of the component can be taken into account. High flexibility, precision, and reproducibility are
typical characteristics of EB technologies and facilities. High productivity is achieved by new technological
solutions like simultaneous interaction of the EB in several processing areas (spots) or by carrying out several
processes simultaneously in modern EB facilities and systems (such as multi-chamber, lock-type and other
concepts). The influence of beam parameters and energy transfer conditions on the microstructure of the
materials and its properties will be discussed for different EB technologies. Information on ideal treatment
conditions will be given. The paper deals with the current development state regarding beam deflection
techniques, technological processes and some facility concepts, and with the state of industrial application.
KEYWORDS: electron beam processes, surface treatment, combined technologies, welding, engraving, profiling,
beam deflection techniques, materials (structure-property relation), applications
INTRODUCTION
By using the advantages of EB, modern EB technologies differ
from other technologies in their advantageous characteristics
(Tab. 1).
These characteristics are typical for all EB technologies, i.e.
welding, surface treatment, surface ablation or perforation,
which are carried out as one-spot techniques. If a multi-spot
technique or multi-process technology is applied, the effects of
these characteristics are much more effectively.
The present paper will exemplarily demonstrate the state of
the art of development and application of EB technologies.
MULTI-TOOL ELECTRON BEAM
Beam deflection techniques
The development of the two-dimensional high-frequency
Rolf Zenker
TU Bergakademie Freiberg, IWT,
Zenker Consult, Mittweida, Germany
Paper presented at the European Conference „Innovation in heat
treatment for industrial competitiveness”, Verona, 7-9 May,
organised by AIM
beam deflection technique was the beginning of a new area
of thermal electron beam (EB) technologies. The high speed
scanning (HSS) technique has been available since 1986 [1]-[3].
In 2000 a high frequency 3D beam deflection technique was
created with new possibilities for load and couture specific EB
technologies, not only for surface treatment [4]-[10] but also
for welding [7][10]-[13] and engraving [14]-[15]. These beam
deflection techniques are based on the fact that the EB can act
simultaneously in several spots [10]-[13]. In this case the same
task is realised in every spot.
A defined and exact positioning of the (mostly oscillation)
Electron beam (EB)
excellent formability and deflect ability
good beam profile
high efficiency
large penetration depth
high beam stability
EB technologies
high productivity
excellent flexibility
good process safety
high reproducibility
ecologically friendly
s
Tab. 1
Characteristics of EB and EB technologies (behind
others) .
Principali caratteristiche del fascio elettronico e delle tecnologie EB.
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Combined surface treatment (tools, automotive components)
Combination EBH/Nitriding
With regard to complex load conditions for most tools and
components, especially close to the surface, the properties attainable
by single treatments (mechanical, thermal, thermochemical
and coating technologies), in particular, often are
insufficient. Therefore, combined processes (duplex or hybrid
process) (Fig. 6a) came into the focus of examinations and
meanwhile of industrial application [17]-[20].
In case of the sequence of the combined surface treatment
combination EBH+nitriding (N) or nitrocarburising (NC) the
level of the processing temperature of N or NC in relation to
the tempering temperature of the bulk material determines the
success of this treatment combination.
The better the tempering stability of the steel the smaller is the
hardness reduction in the previously produced EBH layer as a
result of the subsequent nitriding process.
It is true that a subsequent EBH after N (NC) transforms the
compound layer partially (wider seam of pores), but the hardness
of the diffusion layer is higher than after EBH or N (NC)
[21]. In case of the component shown in Fig. 6b hardness rises
by ~ 200HV0.3 (Fig. 6c).
It has been shown that in the case of optimised process parameters
the advantages of this combined treatment complement
Trattamenti superficiali
a b c
s
Fig. 4
EB hardening of a power train component (two-process EBH technology) - a) Controlling component; b) Twoprocess
technology; c) EB hardening depths.
Indurimento superficiale mediante EB di un componente di un sistema di potenza (tecnologia EBH).
a b c
s
Fig. 5
EB Hardening of a spherical surface (flash technique) - a) Calotte carrier; b) Energy transfer field; c) Process
thermocycle.
Indurimento superficiale mediante EBH di una superficie sferica (tecnica flash).
each other and the disadvantages of the single processes cancel
each other out at least partially [20][21].
Combination of EBH and HC
Hard coatings based on titanium, aluminium or chromium
carbides are successfully applied as hard wear resistant layers
for tools and components. These hard but also brittle coatings
are often unable to bring their excellent properties fully to bear
on relatively soft base materials. Therefore the base materials
are usually subjected to additional heat treatment before or after
hard coating [22]-[24]. It is possible to limit the heat treatment
to the highest loaded areas and up to the depth where a
martensitic transformation is necessary. The thermal loading
of the overall component is minimised.
With regard to a subsequent heat treatment of hard coated
steels it allows for prevention of undesirable changes of composition,
structure and properties of the hard coating [25][26].
A very short interaction time and the process-related vacuum
support these effects. Moreover, the electron beam hardening
technology is well known to cause small changes in size and
shape which means that distortion is also reduced in that way
also. A combined EBH+HC is successful only if the treating
temperature of the hard coating process is lower than the tempering
temperature of the bulk material [26].
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Regarding the application of the HC+EBH combination the
surface deformation due to the martensitic transformation
must be taken into consideration (Fig. 7c).
EB WELDING
Welding of steel (powertrain components)
A typical powertrain welding unit is shown in Fig. 8a). The
conventional procedure used for the EB or laser welding of
gear components is a tack welding in a first step and then
the joining of the two welding partners by single-pool welding
in a second step.
By contrast, in multi-pool welding the components are fixed
simultaneously at the time of the first interaction of the EB
with the material at several points (Fig. 8b).
It follows in the same processing step a simultaneous movement
of several melting pools along one and the same welding
seam (Fig. 8b) up to an overlapping zone in the area
which has already been welded. The number of welding
pools depends on the size and geometry of the part and on
the deflection width of the EB. In comparison to the above
mentioned two-step technology the welding time is reduced
up to one third and also the distortion is minimized because
of the lower heating of the parts [7][10][13].
A speciality of multi-pool EB welding is that the welding
seam is not perpendicular to the surface (Fig. 8b). The incidence
angle of the EB and, consequently, the seam depend
on the welding diameter of the welding circle and the distance
between beam source and component (Fig. 8c) [10]
[13].
One important fact for a successful application of multi-pool
welding is that the program must be optimized in relation to
the jump frequency of the beam from one melting pool to the
next and the beam oscillation for an open vapour capillary.
Welding of al alloys (cylinder liner ensembles)
The production of engine blocks as hybrid casting is state of
the art. In the automotive industry cast iron cylinder liners
are usually applied but there are also cylinder liners made of
spray-formed Al materials. The cast engine block either consists
of different Al or Mg alloys.
One of the technical difficulties is the precise positioning of
each cylinder liner in the mould, one after the other. A technologically
more smarter and more profitable technology is the
Trattamenti superficiali
positioning of the liners as so called “liner ensemble” (Fig. 9a).
In this case several (2…6) cylinder liners are assembled by EB
welding and positioned in the mould as an ensemble in one
step. The technical expenditure is much lower [10][15].
Electron beam welding for that application is realised with
two-spot techniques using a welding spot and a smoothing
spot in one run (Fig. 9c, d).
It has to be taken into account that water jackets are integrated
between the cylinder liners which must stay open after welding.
Because of the high penetration depth of the EB, the welding
takes place only from one side across the water channel up
to a depth of 45 mm without closing it (Fig. 9c, d). The diameter
of the hole must be ≥ 3.0 mm, but this is normal standard
design. The fact that the liners are welded from one side contributes
to a very economical production [15].
The application of the two-spot (pool) technique is necessary
because most spray formed alloys cannot be welded easily and
have a very rough welding bead. The task of the second spot
is to smooth the bead.
EB SURFACE ABLATING
Engraving (shaft for force fit with tube)
EB engraving follows the well-known method of producing
lateral surface patterns to improve the sliding conditions [27]
[28], produce reservoirs for colour particles [29], texture the
cold rolls to improve sheet quality [30] and - in this present
case - to increase the friction in force fits [13].
The principle of these different processes is the same (Fig. 10).
At first, the EB with a small diameter remelts the surface in
a small pool (Fig. 10a). Then a vapour capillary is produced
because of the high beam energy. The vaporised material and
some of the liquid material squirt out of the capillary and a
molten shell is formed around it (Fig. 10b). Depending on the
material and the beam parameters dimples and/or protrusions
develop (Fig. 10c).
By applying the EB multi-spot technique, many protrusions
can be generated simultaneously around dimples, as spot lines
(up to 200 spots per line, Fig. 11a) or as patterns (on a plane
surface up to 3.500 spots) during less than 0.15 ms [31].
Large protrusions (Fig. 11b, c) are desired and necessary for
force fits of shaft/tube assemblies. The dimples (Fig. 11c) are
necessary because they prepare the material for the protrusions.
a b c d
s
Fig. 9
Two-spot welding of cylinder liner ensembles - a) Cylinder liner ensemble; b) Single cylinder with water channels;
c) Welding seam (schematic); d) Welding seam.
Saldatura a due fasci di un complesso di elementi cilindrici allineati.
la metallurgia italiana >> aprile 2009 5

Trattamenti superficiali Tm, Casting) and
- metallurgical connection in combination with mechanical interlocking
(Tm, Insert ~ Tm, Casting)
As it is known, further factors such as the casting method and
conditions, temperature control, the position in the mould, etc.
affect the result with regard to the form fit and connection between
insert and casting.
If the insert is made of an Al alloy as well as the casting the contact
between insert and casting is generally good (mechanical
interlocking and metallurgical connection, Fig. 12c). Ultrasonic
measurements support these results. In areas where there is no
contact the ultrasonic waves are reflected at the interface. In
case of metallurgical connection there is no or a weak signal
(Fig. 12e).
To assess the bond strength of castings with profiled inserts,
samples were analysed using static tensile tests. The analyses
showed that a bond strength which was 3…10 times higher
than it was for non-profiled inserts (1…2.5 kN) was achieved
by EB profiling [15] (basis: force up to the point where the insert/casting
connection is separated). This may be attributed
in particular to the fact that the breaking of samples did not
occur at the interface between insert and casting, as it was the
case in non-profiled reference samples, but predominantly in
the casting.
6 aprile 2009

Memorie >>
300
3] PANZER, S.; MÜLLER, M.: Härten von Oberflächen mit
Elektronenstrahlen. In: HTM 43(1980), 2, pp. 103-111
4] ZENKER, R., Electron beam surface treatment: industrial
application and prospects. In: Surface Engineering 12(1996),
4, pp. 296-297
5] ZENKER, R.; WAGNER, E.; FURCHHEIM, B.: Electron
beam – a modern energy source for surface treatment. In:
6th International Seminar of IFHT: Advanced Heat Treatment
Techniques Towards the 21st Century: 15.-18.10.1997,
Kyongju, 1997
6] ZENKER, R.; FRENKLER, N.; PTASZEK, T.: Electron
beam surface treatment of Al, Mg, and Ti alloys. In: Proceedings
of the 7th International Seminar of IFHT: Heat treatment
and surface engineering of light alloys: 15.-17.9.1999,
Budapest, 1999, pp. 18-21
7] ZENKER, R.: Electron beam surface treatment and multipool
welding – state of the art. In: EBEAM 2002, International
Conference on High-Power Electron Beam Technology:
27.-29.10.2002, Hilton Head Island, 2002, pp. 12-1–12-5
8] ZENKER, R.: Structure and properties of electron beam
surface treatment. In: Advanced Engineering Materials
6(2004), 7, pp. 581-588
9] ZENKER, R.: Elektronenstrahl-Randschichtbehandlung,
Innovative Technologie für höchste industrielle Ansprüche.
Monographie, pro-beam AG & Co. KGaA, 2003
10] ZENKER, R.: Elektronenstrahlbearbeitung für Powertrainkomponenten.
In: Kooperationsforum Metalle im Automobilbau,
Innovationsforum in Be- und Verarbeitung,
29.11.2005, Hof, 2005
11] MATTAUSCH, G.; MORGNER, H.; DAENHARDT, J.;
ET. AL.: Survey of electron beam technologies at FEP. In:
Proceedings / EBEAM 2002: International Conference on
Trattamenti superficiali
a b c
s
Fig. 12
Multi track profiling of bearing insert.
Profilo a più tracce di un particolare di cuscinetto.
d e
High-Power Electron Beam Technology, Hilton Head Island,
27.-29.10.2002, pp. 11/1-11/11
12] LOEWER, T.: Analysis, visualisation and accurate description
of an electron beam for high repeatability of industrial
production processes. In: Proceedings of the 7th
International Conference on Electron Beam Technologies,
Varna, 1.-6.6.2003, pp. 45-50
13] ZENKER, R.; BUCHWALDER, A.; FRENKLER, N.;
THIEMER, S.: Moderne Elektronenstrahltechnologien zum
Fügen und zur Randschichtbehandlung. In: Vakuum in der
Praxis, 17(2005), 2, pp. 66-72
14] ZENKER, R.; BUCHWALDER, A.; SPIES, H.-J.: New
electron beam technologies for surface treatment. In: Proceedings
of the 7th International Conference on Electron
Beam Technologies: 1.-6.6.2003, Varna, 2003, pp. 202-209
15] ZENKER, R.; KRUG, P.; BUCHWALDER, A.; DICK-
MANN, T.; FRENKLER, N.; THIEMER, S.: Elektronenstrahlschweißen
und –profilieren von sprühkompaktierten
Zylinderlaufbuchsen aus Al-Si-Werkstoffen. In: Zylinderlaufbahn,
Kolben, Pleuel – Innovative Systeme im Vergleich,
Tagung Böblingen, 7.-8.03.2006, VDI-Verlag GmbH: Düsseldorf,
2006, VDI-Berichte 1906, pp. 259-274
16] BUCHWALDER, A.: Beitrag zur Flüssigphasen-Randschichtbehandlung
von Bauteilen aus Aluminiumwerkstoffen
mittels Elektronenstrahl. Dissertation TU Bergakademie
Freiberg, 2007
17] ZENKER, R.: Kombinierte thermochemisch-thermische
Wärmebehandlung. Neue Hütte 28(1983), 10, pp. 379-385
18] SPIES, H.-J.: Erhöhung des Verschleißschutzes von Eisenwerkstoffen
durch die Duplex-Randschichttechnik.
Stahl und Eisen 117(1997), 6, pp. 45-62
19] KEßLER, O., HOFFMANN, F., MAYR, P.: Combinations
of coating and heat treating processes: establishing a system
for combined processes and examples. Surf. Coat. Technol.,
108-109(1998), pp. 211-216
20] ZENKER, R., SPIES, H.-J.: 15 Jahre industrielle Anwendung
der Elektronenstrahl Randschicht-behandlung. 57. Härtereikolloquium,
Wiesbaden, Germany, Oct 10-12, 2001
la metallurgia italiana >> aprile 2009 7

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Memorie >>
Alluminio e leghe
SQUEEZE CAST AUTOMOTIVE
APPLICATIONS AND DESIGN
CONSIDERATIONS
Z. Brown, C. Barnes, J. Bigelow , P. Dodd
With an increasing emphasis on vehicle weight reduction, the demand for lighter weight automotive components
continues to increase. Squeeze casting is an established shape-casting process that is capable of producing
lightweight, high integrity automotive components that can be used for structural applications.
In recent years the squeeze casting process has been used with various aluminum alloys to produce
near-net shape components requiring high strength, ductility, pressure tightness or high wear resistance
[1]. Squeeze casting has proven to be an economical casting process for high volume applications and
offers design and materials engineers an alternative to conventional casting processes such as gravity
permanent mold (GPM), low pressure die casting (LPDC), sand cast aluminum/ iron, and conventional
high pressure die casting (HPDC).
This paper describes Contech’s squeeze casting technology (P2000 TM ) and provides examples of high
volume automotive components manufactured at Contech. This paper also includes product design
considerations, an overview of process simulation techniques, a comparison of mechanical properties, and case
studies for select automotive components.
KEYWORDS: squeeze casting, aluminum, automotive applications, die casting, safety critical
INTRODUCTION
Conventional HPDC is a well-established process for the
manufacturing of a wide variety of aluminum automotive
components such as engine blocks, pump housings, oil
pans, and transmission components. Conventional HPDC
has many advantages including near-net shape capability,
low manufacturing cost, and excellent dimensional accuracy
and repeatability.
Achievable casting performance is limited however, due to defects
that emerge during the casting process such as gas and
shrink porosity, laminations, and inclusions. In addition,
HPDC components are not considered heat treatable, which
further limits achievable performance.
For applications that require higher component integrity
(high strength and ductility, reduced porosity, uniform
microstructure, and ability to heat treat), alternative cast-
Zach Brown, Chuck Barnes, Joe Bigelow
Contech U.S. LLC
Paul Dodd
Contech UK LLC
Paper presented at the International Conference “High Tech
DieCasting”, Montichiari, 9-10 April 2008, organised by AIM
ing processes such as squeeze casting should be considered.
Squeeze casting is an established process that builds upon conventional
HPDC practices and is used to manufacture various
automotive components that require high strength and
ductility, as well as applications that require high pressure
tightness or wear resistance. Examples include steering
column components, steering knuckles, control arms,
suspension links, pump housings, and various powertain
components [1]. The squeeze casting process is capable
of producing components with dimensional accuracy and
near-net shape capability that is equal to conventional HPDC.
Unlike HPDC however, the squeeze casting process is capable
of producing higher integrity components. As a result,
design engineers are able to further optimize current aluminum
designs or substitute aluminum in place of heavy materials
such as steel and cast iron.
SQUEEZE CASTING TECHNOLOGY (P2000 TM )
Squeeze casting can be divided into two categories; “direct”
and “indirect”. Direct squeeze casting, often termed “liquidmetal
forging”, consists of pouring metal into a lower die contained
within a hydraulic press. The upper die closes over the
lower die and high pressure is applied throughout the entire
solidification process. In contrast, indirect squeeze casting consists
of pouring molten metal into the cold chamber of a die
la metallurgia italiana >> marzo 2009 1

Alluminio e leghe

Memorie >>
s
Fig. 3
Example of an aluminum bearing cap that was
converted from GPM to the P2000TM squeeze cast
process. All secondary machining operations were
eliminated.
Esempio di una calotta in alluminio prodotto mediante
squeeze casting (P2000TM ) anzichè in gravità in
conchiglia. Tutte le operazioni secondarie di lavorazione
sono state eliminate.
Solidification simulations are used mainly to predict shrink
porosity and evaluate directional solidification. Fill simulations
are used to identify potential fill related issues such
as laminations due to merging flow fronts, turbulence, and
improper venting.
Tooling design engineers rely on these tools when optimizing
gating size and location, cooling line placement, cooling media
and temperature, die configuration, and process development.
New process simulation techniques are now being used
to predict the microstructure at various locations throughout
the casting. Since strength and ductility are influenced by
the microstructure, this tool can be used to predict mechanical
properties at various locations throughout the casting. This
information can then be used when interpreting FEA results.
Other new developments allow for the prediction of residual
stresses induced during the casting and heat treating process.
Most commercially available FEA software does not
consider residual stress. High residual stress can result in
lower than expected component performance and dimensional
capability.
Design Recommendations
The squeeze casting process is capable of producing complex
geometries with high dimensional accuracy and repeatability.
This allows designers to create near-net shapes, thus
Alloy-Temper
A356- T6
ADC12-F
ADC12- T5
ADC12- T6
Yield (MPa)
220-260
140-170
230-260
290-320
Tensile (MPa)
290-340
200-270
280-320
344-380
s
Tab. 1
indici di prestazione dei materiali candidati.
Materials indexes for candidate materials.
% Elongation
9-15
2-3.5
1-3
2-5
Alluminio e leghe
s
Fig. 4
Example of P2000TM squeeze cast knuckle.
Esempio di snodo prodotto con la tecnica di squeeze
castingP2000TM .
minimizing secondary machining operations. Fig. 3 shows an
example of an aluminum bearing cap that was converted from
gravity permanent mold to squeeze casting. Due to the near net
shape capability of the squeeze casting process, all secondary
machine operations were eliminated. The use of precision
cores with minimal draft (less than .5º per side) eliminated
the need for a secondary drilling operation. The flatness
and surface finish requirements were achieved in the ascast
condition, eliminating the milling operation.
For applications that require high mechanical stiffness, design
engineers must consider both the modulus of elasticity
and section modulus. Modulus of elasticity is a function of the
stiffness of the alloy itself and is fairly similar for most aluminum
casting alloys. Section modulus is a function of stiffness
from the casting geometry. Increasing the section modulus
through design can offset issues with a lower modulus of
elasticity. Complex geometries such as ribs, pockets, and
u-shaped sections can be used to increase section modulus. It
is recommended to avoid drastic wall thickness changes and
isolated thick sections. By avoiding localized thick sections and
drastic wall thickness changes, the tendency to form shrink
porosity is greatly reduced. Isolated thick sections can also induce
stress concentration points and cause casting defects such
as hot tears and heat sinks.
Material Selection
One important advantage of the squeeze casting process is that
it is can be used with various alloy/ heat treat combinations
that can be tailored to meet design requirements. Primary alloys,
such as A356 (AlSi7Mg) are used in the T6 condition for
applications that require high strength and ductility such as
control arms, steering knuckles, and suspension links. Secondary
alloys such as
ADC12 (AlSi11Cu3Fe) are used in the
as-cast, T5, and T6 conditions for applications
that require high strength,
pressure tightness, and wear resistance.
Typical mechanical properties
are shown in Tab. 1.
P2000TM Hardness (HBN)
85-100
95-105
110-130
120-140
APPLICATIONS
Fig. 4 shows an example of a squeeze
cast front steering knuckle. In this
la metallurgia italiana >> marzo 2009 3

Alluminio e leghe

Memorie >>
Alluminio e leghe
CASTABILITY MEASURES FOR
DIECASTING ALLOYS: FLUIDITY,
HOT TEARING, AND DIE SOLDERING
B. Dewhirst, S. Li, P. Hogan, D. Apelian
Tautologically, castability is a critical requirement in any casting process. Traditionally, castability in sand
and permanent mold applications is thought to depend heavily on fluidity and hot tearing. Given
capital investments in dies, die soldering is a critical parameter to consider for diecasting. We discuss
quantitative and robust methods to insure repeatable metal casting for diecasting applications by investigating
these three areas. Weight reduction initiatives call for progressively thinner sections, which in turn
are dependent on reliable fluidity. Quantitative investigation of hot tearing is revealing how stress develops
and yields as alloys solidify, and this has implications on part distortion even when pressure-casting
methodologies preclude hot tearing failures.
Understanding the underlying mechanism of die soldering presents opportunities to develop methods
to avoid costly downtime and extend die life. Through an understanding of castability parameters,
greater control over the diecasting process can be achieved.
KEYWORD: castability, die soldering, fluidity, hot tearing, part distortion, residual stress
INTRODUCTION
Over the years, castability has been addressed through various
angles and perspectives. However no matter what has
been accomplished, it is fair to state that at the present there
is not a single method that the community can point to as a
means of defining an alloy’s castability in terms of measurable
quantitative parameters. It is critical that means for
controlling the casting process be developed. Without robust
measures, one will not be able to control the casting process.
It is the latter that is the motivating force behind this project.
Hopefully, the investigative techniques being developed in
this research will become standardized so that an accepted
lexicon and methodology is practiced throughout the
casting community.
This paper will focus on three parallel lines of research with
applicability to light metals diecasting: fluidity, hot tearing
(as it relates to stresses developing within solidifying metals
as a function of chemistry and microstructure), and die soldering.
Each of these three areas of research has the potential
to positively benefit the HPDC industry, either directly or as
B. Dewhirst, S. Li, P. Hogan, D. Apelian
Metal Processing Institute - WPI, 100 Institute Road -
Worcester, MA 01609 USA
Paper presented at the International Conference High Tech
DieCasting, Montichiari, 9-10 April 2008, organised by AIM
an accompanying benefit to research conducted for other
purposes.
Vacuum fluidity testing allows for the evaluation of various
alloys and process modifications in a laboratory setting
under rapid solidification conditions, but suffers from
a poor reputation and, as a consequence, has principally been
used for qualitative experimentation. Hot tearing, a consequence
of stresses developing during feeding until the casting
tears itself apart, is not found in alloys used in HPDC, but the
investigative techniques being applied to understand hot
tearing are providing a window into how these stresses
develop. Die soldering is important because, in improperly
designed castings, soldering can be a significant problem that
can severely inhibit productivity.
FLUIDITY
Fluidity is a material’s ability to flow into and fill a given cavity,
as measured by the dimensions of that cavity under specified
experimental conditions, and fluidity is heavily dependent
on heat flow during solidification.
Investigations into the impact of foundry variables such as
mold coatings, alloying additions, head pressure, and especially
superheat have been investigated and correlated with
mechanisms. For sand and permanent mold castings, it is
abundantly clear that increasing solidification range results
in decreasing fluidity (all other factors being equal). Specific
investigations are often alloy or metal/mold/coating
specific in scope, but very subtle influences of minor varia-
la metallurgia italiana >> marzo 2009 37

Alluminio e leghe

Memorie >>
s
Fig. 1
Cast Iron Mold designed to detect the onset of
the hot tearing. Commercial cast alloy 713 and 518
were evaluated; the former is known to be sensitive to
hot tearing, and the latter has good resistance to hot
tearing. The pouring temperature was set at 60˚C
above the melting point of the alloy during this effort.
The mold temperature was maintained around 200˚C.
Stampo in ghisa progettato per rilevare l’insorgenza di
cricche a caldo. Sono state valutate le leghe commerciali
713 e 518; la prima è risaputa essere sensibile alla
criccatura a caldo, mentre la seconda ha una buona
resistenza verso tale fenomeno. In questa prova la colata
è stata eseguita a una temperatura di 60˚C al di sopra
del punto di fusione della lega.La temperatura dello
stampo è stata mantenuta intorno ai 200˚C.
Hot tearing susceptibility of alloys is greatly influenced
by solidification behavior of molten metal in the mushy
zone. Solidification can be divided into four stages [15]: (i)
Mass feeding where the liquid and solid are free to move; (ii)
Interdendrtic feeding when the dendrites begin to contact
each other, and a coherent solid network forms; (iii)
Interdendritic separation. With increasing fraction solid, the
liquid network becomes fragmented. If liquid feeding is not
adequate, a cavity may form. As thermal contraction occurs,
strains are developed and if the strain imposed on the network
is greater than a critical value, a hot tear will form and
propagate. Lastly, in stage (iv), Interdendritic bridging or
solid feeding occurs. Simply stated, hot tearing occurs if the
solidification shrinkage and thermal deformation of the solid
cannot be compensated by liquid flow.
Measuring the development of strains and the evolution
of hot tearing during solidification is not trivial. The
Metal Processing Institute is a member of the Light Metals
Alliance, and we have teamed up with our alliance partner
CANMET to address hot tearing in aluminum alloys. The
constrained bar mold used in this study was developed at
CANMET Materials Technology Laboratory (MTL) and designed
to measure load and temperature during solidification.
Fig. 1 shows one of the mold plates and testing setup.
The mold is made of cast iron and coated with insulating mold
wash. The test piece has two arms. One test arm (12.5mm) is
constrained at one end with heavy section (22.5mm) to keep
the bar from contraction, so the tension will be developed
and hence cracking could be induced during solidification. The
other arm is for load and temperature measurement with
one end connected to a load cell. This opened end of the
mold is closed with a graphite cylinder block which can move
freely in horizontal direction. The block is connected to the solidifying
material on inner side with a screw and on external
Alluminio e leghe
s
Fig. 2
(a) Temperature-load-time curves of alloy 713;
(b) Derivative of Load vs. time curve.
(a) Curve temperatura-carico-tempo per la lega 713;
(b) andamento della derivata del carico vs. tempo.
side with a load cell. Two K-type thermocouples are used for
the temperature measurement. One is positioned at the riser
end and the other at the end of the bar as shown in Fig. 1.
After pouring the melt into the mold, the temperature and
load were recorded with a computer data acquisition system.
Fig. 2 and 3 show the measured temperatures and load recorded
during casting as a function of time for alloy 713 and 518
respectively. The load represents the tension force developed
in the casting during solidification. The cooling curve T1 was
recorded with thermocouple tip positioned at the riser end
and T2 with thermocouple tip at the end of the bar as shown in
Fig. 1. A rapid rise in temperature (both curves) was observed
immediately after pouring and the temperature started falling
shortly. It’s noticed that negative loads (compressive forces)
were developed shortly after pouring for the tests, probably
due to the pressure head of the melt [16]. When the rod begins
to solidify but cannot contract freely, the tension force increases.
Fig. 2(b) and 3(b) are derivatives of load vs. time curve to
determine onset of hot tearing. An obvious change in the rate
suggests that cracking might occur there.
From Fig. 2b, load began developing at proximately 9 seconds
and the solidification temperature was around 617˚C (Fig.2a),
then increased rapidly. It is shown that the rate changed
abruptly to zero at 16.5 seconds, suggesting a severe tear occurred
there.
Hot tearing occurred at around 530˚C, corresponding to 94%
solid, according to Pandat Scheil solidification calculation.
The technique developed to measure hot tearing tendency is
a valuable tool to differentiate between alloys and to use it to
optimize alloys for high integrity castings.
la metallurgia italiana >> marzo 2009 39

Alluminio e leghe

Memorie >>
s
Fig. 5
Phase diagram of Aluminum-10% Silicon and
low solubility of Fe.
Diagramma di stato dell’ alluminio-10% silicio con bassa
solubilità del ferro.
the die surface and reduces the contact area and the reaction
between the two.
High temperatures and high melt velocity are process conditions
which lead to soldering.
Of the two, high temperature is the most important to avoid in
order to prevent soldering.
This can most effectively be done through careful design of
the die. By configuring the part and optimizing the design of
the die cooling system, the potential for soldering can be greatly
reduced. It is very important to consider this during the
design phase of a die because once a die is manufactured it
is very difficult to reduce any hot spots. Other potential solutions
include using additional spray in the high solder areas to
reduce temperature or the use of inserts with high conduction
coefficients.
Impingement velocity is important to control as well. The die
surface should be coated with lubricants and is likely oxidized
from prior treatment. A high impingement velocity can wash
these protective coatings off of the die surface, exposing the
die steel to the aluminum alloy and begin erosion of the die
surface. Both of these effects will promote the beginning of die
soldering.
SSM processing can help to reduce both the temperature and
velocities apparent in the casting system, and should help reduce
die soldering [12].
Die coatings can be useful as a diffusion barrier between the
steel in the die and the aluminum in the cast alloy. An effective
coating must be able to withstand the harsh conditions at
the surface of the die, however. Coatings which are sometimes
used include CrN+W, CrN, (TiAl)N and CrC [19]. Additionally,
surface treatments such as nitriding and nitro-carburizing
can help to strengthen the surface and prevent erosion, which
accelerates the soldering process by roughening the surface
and creating local temperature excursions at the peaks of the
die surface which solder very quickly.
Accurate modeling of the casting process during the design
phase is very important to an effective control against die soldering.
All of the previously mentioned controls require ad-
Alluminio e leghe
s
Fig. 6
Change in viscosity of an Al-Si alloy with the
addition of 230ppm Sr [18].
Variazione della viscosità di una lega Al-Si con l’aggiunta
di 230 ppm di stronzio. [18].
ditional cost during the design and manufacturing of the die,
and it must be understood how badly soldering will affect the
process before the costs of any of those controls can be justified.
CONCLUSIONS
Though these three alloy characteristics seem tangentially
related, they are factors that influence castability. In order to
control these castability indices, it is necessary to develop experimental
methods until robust quantitative analysis is possible.
Once quantitative data can be extracted, the improvement
in our understanding will occur. In the case of die soldering,
multiple possible avenues to reduce the problem have been
identified. Even when the initial intention was to resolve problems
occurring in sand and permanent mold castings, such as
hot tearing, the information gleaned about how stresses develop
in liquid metal has wider applicability. Though die casting
usually assures good fluidity through the use of pressure, if
fluidity (and the factors which influence its variation) are well
understood, it is possible to operate within tighter processing
windows.
REFERENCES
1] D.V. RAGONE, C.M. ADAMS, H.F. TAYLOR, AFS Trans. 64,
(1956), p.640.
2] D.V. RAGONE, C.M. ADAMS, H.F. TAYLOR, AFS Trans. 64,
(1956), p.653.
3] M.C. FLEMINGS, Brit. Foundryman 57, (1964), p.312.
4) M.C. FLEMINGS, Solidification Processing. McGraw-Hill,
New York (1974).
5] M.C. FLEMINGS, E. NIYAMA, H.F. TAYLOR, AFS Trans. 69,
(1961), p.625.
la metallurgia italiana >> marzo 2009 41

Alluminio e leghe

Memorie >>
Magnesio e leghe
A PROBABILISTIC APPROACH FOR
MODELLING OF FRACTURE IN
MAGNESIUM DIE-CASTINGS
C. Dørum, D. Dispinar, O.S. Hopperstad, T. Berstad
Quasi-static tensile tests with specimens cut from a generic cast component are performed to characterise the
mechanical behaviour of the high-pressure die cast magnesium alloy AM60. The experimental data is used
to establish a probabilistic methodology for finite element modelling of thin-walled die castings subjected to
quasi-static loading. The cast magnesium alloy AM60 is described using an elastic-plastic constitutive model
consisting of a high-exponent, isotropic yield criterion, the associated flow law and an isotropic hardening rule.
A novel probabilistic approach for modelling of fracture in thin-walled magnesium die-castings using finite
element analysis is developed. The Cockcroft-Latham criterion for ductile fracture is adopted with the fracture
parameter assumed to follow a modified weakest link Weibull distribution. Comparison between the experimental
and predicted behaviour of the cast magnesium tensile specimens gives very promising results.
KEYWORDS: magnesium, die-castings, ductile fracture, weibull distribution, finite element analysis
INTRODUCTION
With increased focus on environmental issues, structural designers
in the transport industry are forced to search for light-weight
solutions. New materials are considered for vehicle design if
they provide benefits at an affordable cost. The cold-chamber
high pressure die casting method is an important production
method for aluminium and magnesium castings; particular suitable
for fully automatic, high productivity, high volume production
of complex near net shape parts.A major challenge with this
production method is to optimise the process parameters with
respect to the part design and the solidification characteristics
of the alloy in order to obtain a sound casting without casting
defects. Unbalanced filling and lack of thermal control can cause
bifilms, porosity and surface defects due to turbulence and solidification
shrinkage. Consequently, the fracture behaviour of
cast components can be of stochastic character.
To be efficient in the development of new products it is necessary
to use finite element (FE) analysis to ensure a structural design
that exploits the material. In order to be able to obtain a reli-
C. Dørum
SINTEF Materials and Chemistry, No-0314 Oslo, Norway
Structural Impact Laboratory (SIMLab), Centre for Research-based
Innovation, No-7491 Trondheim, Norway
D. Dispinar
SINTEF Materials and Chemistry, No-7465 Trondheim, Norway
O.S. Hopperstad
Structural Impact Laboratory (SIMLab), Centre for Research-based
Innovation, No-7491 Trondheim, Norway
T. Berstad
SINTEF Materials and Chemistry, No-7465 Trondheim, Norway
Structural Impact Laboratory (SIMLab), Centre for Research-based
Innovation, No-7491 Trondheim, Norway
able prediction of the structural behaviour using such analyses,
an accurate description of the material behaviour is essential.
Hence, a reliable failure criterion is also required, that enables
the designer to exploit the potential of the cast material. This
work presents a new probabilistic approach for finite-element
modelling of the structural behaviour of thin-walled cast magnesium
components.
Fig. 1 shows the geometry of the generic AM60 component in-
s
Fig. 1
Illustration of generic cast component: Length
= 400 mm, thickness = 2.5 mm, width = 80 mm,
height = 40 mm.
Illustrazione di un generico componente pressocolato:
Lunghezza = 400 mm, spessore = 2.5 mm, larghezza =
80 mm, altezza = 40 mm.
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s
Fig. 4
SEM images from the fracture surface of
samples showing the crumpled oxides.
Immagini SEM della superficie di frattura dei campioni
che mostrano gli ossidi.
criterion is reached. The model has been implemented in explicit
finite element code LS-DYNA [4].
The high-exponent isotropic yield criterion [5, 6] is written in
the form
(1)
where σ 1 and σ 2 are principal stresses in plane stress and k is
a material parameter. The flow stress σ y is defined by the isotropic
hardening rule
(2)
where ε e is the effective plastic strain, σ 0 is the proportionality
limit, and Q i and C i are hardening parameters. Using a least
squares method, the hardening parameters were determined
from the Cauchy stress versus logarithmic plastic strain curves
in Fig. . Any variation in flow stress with position in the casting
was not accounted for in the FE simulations, and thus a mean
hardening curve was applied.
In the present model, a criterion of ductile fracture proposed
by Cockcroft and Latham [7] is added. The fracture criterion is
coupled with the element-erosion algorithm available in LS-
DYNA [4]. As the fracture criterion is reached in an element,
this element is removed (eroded) from the finite element model.
The fracture criterion can be expressed as
(3)
where σ 1 is the maximum principal stress and W c is the critical
value of the integral W. Hence, fracture occurs when W = W c .
Henceforth, W c will be referred to as the fracture parameter,
while W will be denoted the Cockcroft-Latham integral. It is
seen that fracture cannot occur when the maximum principal
stress is compressive and that neither stresses nor strains alone
are sufficient to cause fracture. Furthermore, the fracture strain
increases with decreasing stress triaxiality (in the shear tests,
the stress triaxiality is significantly reduced compared to the
uniaxial tension test).
The uniaxial tensile test specimens failed before the point of diffuse
necking for the AM60 alloy, and, accordingly, the stress and
strain field are uniform up to fracture. Hence, the fracture parameter
is obtained as the area under the work-hardening curve.
Zhou and Molinari [8, 9] propose a micro-cracking model for
brittle materials (ceramics) considering the stochastic distribution
of internal defects. The model introduces a Weibull distri-
Magnesio e leghe
s
Fig. 5
Picture of test specimens with flow lines on the
surface.
Immagine dei campioni di prova con linee di scorrimento
sulla superficie.
bution [10] of the local strength of cohesive elements. Thus, the
probability of introducing a weak cohesive element increases
with the cohesive element size. Inspired by this idea, the fracture
parameter of a finite element is assumed to follow a modified
weakest-link Weibull distribution in the current study.
The Weibull distribution gives the fracture probability P (σ) of
a material volume to under effective tensile loading, i.e.
(4)
where V is the volume, V 0 is the scaling volume, σ 0 is the scaling
stress, and m is the Weibull modulus. Since cast magnesium
is not a brittle material, the use of a critical fracture stress
is not justified. Instead, the Cockcroft-Latham ductile fracture
criterion is adopted, and the fracture probability of a material
volume is recast as
(5)
where W c0 is the scaling value of the fracture parameter. By
using a random number generator and inverse sampling, this
Weibull distribution of fracture parameters can then be assigned
to the integration points in the FE mesh. With this approach,
a small element in the FE model will most probably be
given more ductile material properties than a larger element.
Fig 6. compares the numerical predictions with the experimental
results from the tensile tests on cast magnesium AM60.
Here, the uniaxial tension test specimens were modelled by
720 shell elements (i.e., a characteristic element size equal to
1.0 mm) It is seen that the observed experimental scatter is well
reproduced numerically.
CONCLUDING REMARKS
The quasi-static behaviour of high-pressure die cast magnesium
alloy AM60 has been studied through tensile tests. The specimens
were taken from various positions in the cast profile. The
experimental data were used to develop a probabilistic method
for finite element modelling of thin-walled die castings subjected
to quasi-static loading. The ductility of the specimens cut from
the castings depends on the position in the casting. There are also
significant variations in ductility when comparing the measured
characteristics of specimens cut from different castings that were
cast under equal casting conditions. Thus, as a result of unstable
flow of the liquid magnesium in the mould cavity, the mechanical
properties of the casting are of stochastic nature. By combin-
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Acciaio inossidabile
THE EFFECT OF AUSTENITE VOLUME
FRACTION ON THE DEFORMATION
RESISTANCE OF 409 STAINLESS
STEELS DURING HOT-STRIP ROLLING
D. Chae, S. Lee, S. Son
The mill log data obtained from the hot-strip rolling of 409 stainless steels were analyzed in order to investigate
the effect of the chemical composition on the deformation resistance. The results showed that the deformation
resistance depended sensitively on the austenite stabilizing capability of the material chemistry, suggesting
the austenite volume fraction as a dominant factor in controlling the deformation resistance. Deformation
resistance ratio (DRR) was defined as a ratio of the deformation resistance of a two-phase (ferritic+austenitic)
microstructure to that of a fully ferritic microstructure. The dependence of DRR on the austenite volume
fraction appeared to be linear, which was also observed by the plane strain compression tests performed on the
laboratory specimens with various austenite volume fractions. The implication of this result is that during the
hot-strip rolling of 409 stainless steels with a two-phase microstructure, these steels are likely to deform in an
equal-strain manner.
KEYWORDS: 409 stainless steel, mean flow stress, deformation resistance, two-phase material, austenite potential
and mill log analysis
INTRODUCTION
In hot-strip rolling, precise thickness control requires an accurate
prediction of the roll force. The accurate prediction of
the roll force, in turn, depends on the accurate calculation of
the hot flow strength of the material because it significantly
affects the pressure at the interface between the work roll and
the rolled material. In order to calculate the hot flow strength
as a function of rolling parameters which are representative
of each rolling pass, an average flow strength, hereafter called
UNS
S40910
S40920
S40930
S40945
S40975
C
0.030
N
0.030
Dongchul Chae, Soochan Lee,
Seunglak Son
Posco, Pohang, Korea
Paper presented at the 3rd International Conference Thermomechanical
Processing of Steels, organised by AIM, Padova, 10-12 September 2008
‘deformation resistance’ is defined over the total applied strain
during a rolling pass[1, 2]. Industrially, deformation resistance
is analyzed using mill log data to develop and refine rolling
mill models.
The 409 stainless steels are characterized by relatively low
carbon and nitrogen contents with approximately 11% chromium
in their chemical compositions (Tab. 1). Titanium (and/
or niobium) is usually added enough to tie up carbon and nitrogen
atoms. Due to the fact that titanium is a strong ferrite
former, these steels are fully ferritic in an annealed condition.
Composition percentage, max or range
Cr Ni
Other elements
10.5- 0.5 Ti 6(C+N) min, 0.5 max; Nb 0.17 max
11.7 0.5 Ti 8(C+N) min, Ti 0.15-0.5; Nb 0.10 max
0.5 Ti+Nb 0.08+8(C+N) min, 0.75max; Ti 0.05 min
0.5 Nb 0.18-0.40, Ti 0.05-0.20
0.5-1.0 Ti 6(C+N) min, 0.75 max
s
Tab. 1
Chemical compositions of ferritic
stainless steel grades containing 11% chromium
in ASTM A 240/A 240M-00.
Composizione chimica dei gradi di acciaio inossidabile ferritico
contenenti 11% cromio negli ASTM A 240/A 240M-00.
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s
Fig. 2
The effect of chemical composition on the
austenite volume fraction. The two ingots were quenched
after the heat treatment at 950~1250°C for 1 hour.
Effetto della composizione chimica sulla frazione in volume
del austenite. I due lingotti sono stati temprati dopo
trattamento termico a 950~1250°C per 1 ora.
Acciaio inossidabile
DISCUSSION
The Shaffler diagram represents the as-solidified microstructure
after a rapid cooling as a function of Cr-equivalent
and Ni-equivalent [10]. It is noteworthy that the
compositional range of 409 stainless steels is near the twophase
(martensitic+ferritic) boundary in the Shaffler diagram.
Because the presence of martensite implies the thermal
transformation from austenite, it is likely that the material
may have an austenite phase at high temperatures. This
seems to be closely related to one of the significant features
of the Fe-Cr equilibrium diagram, that is, the phase
boundary between austenite and ferrite fields, known as
the gamma loop. Based on the Fe-Cr equilibrium diagram,
approximately 11% Cr is near the nose of the gamma loop.
Therefore, austenite is likely to form at high temperatures.
In addition, it is also well known that the role of austenite
stabilizing elements such as carbon and nitrogen is to shift
the gamma loop to higher chromium content. Therefore,
the small fluctuation in the austenite stabilizing capability
of the material chemistry can possibly affect the phase balance
between ferrite and austenite significantly.
It can be seen in Fig. 2 that, for two 409 ingots produced
in a laboratory, the high temperature microstructure consists
of two phases, austenite+ferrite, and the maximum
amount of austenite occurs at about 1050 °C.
The microstructures of the 12mm thick hot rolled plates
exposed at 1050°C for 5 minutes
are shown in Fig. 3. The microstructure
of the material A (in Tab.
2) is observed to be fully ferritic
as shown in Fig. 3(a). The region
of light contrast in Fig. 3(b) corresponds
to the elongated martensite
transformed from austenite.
As the Ni-equivalent of the materials
increases, the banded array
of alternating layers of ferrite and
martensite (i.e., austenite) becomes
distinctive. The consequence of the
highest Ni-equivalent is the formation
of a 100% martensitic microstructure
(Fig. 3(e)).
Fig. 3
Light micrographs
quenched after the heat
treatment at 1050°C for 5
minutes in the case of (a) the
material A, (b) the material
B, (c) the material C, (d) the
material D and (e) the material
F in Table 2. Murakami etchant
was used for etching.
Micrografie di pezzi temprati
dopo trattamento termico a
1050°C per 5 minuti nel caso del
(a) materiale A, (b) materiale B,
(c) materiale C, (d)
materiale D e (e) materiale F
di Tebella 2. Per l’attacco chimico
è stato utilizzato il reagente di
Murakami.
la metallurgia italiana >> febbraio 2009 57
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Acciaio inossidabile

Material
M1
M2
M3
M4
lowing equation:
Memorie >>
C
0.005
~
0.015
[4]
where, P, B, Ld and Qp are roll
force, strip width, the projected
length of the arc of contact between
the roll and the rolled material
and Sims geometrical factor, respectively.
It should be reminded
that the MFS obtained from the
stress-strain curve can be compared
to the deformation resistance
obtained from the rolling data. Thus, MFS in Equ. 2 were
replaced with Km. Then, the DRR is also defined from the
deformation resistance calculated from the mill log data.
where, Km,α and Km,γ are the deformation resistances of a
single ferritic microstructure and a single austenitic microstructure,
respectively.
The dependence of deformation resistance on the entry
temperature was analyzed from the first rolling pass, F1,
to the last rolling pass, F7. Four materials were examined
in details. The differences in the composition and the microstructure
are summarized in Tab. 3. Among the four
materials, it was observed that M2 and M4 had two phase
(ferrite+austenite) microstructures right before the entry to
F1. The austenite volume fractions of M2 and M3 materials
right before the entry to F1 were measured to be 7% and
33% as shown in Fig. 6, respectively while austenite formation
was not found from the M1 and M4 materials.
The rolling conditions for the materials, M2, M3 and M4,
were also much similar to those of M1 except for the entry
temperatures. It is a usual occurrence that if the rolling
temperature increases, then the deformation resistance decreases.
This usual expectation is confirmed from the observation
that the deformation resistance of the fully ferritic
M4 is lower at the same pass than that of the fully ferritic
M1 because the rolling temperature of M4 was higher than
that of M1. However, comparing the responses of M1 with
those of M2 and M3, a significant trend is observed (Fig.
[5]
Acciaio inossidabile
Chemical composition (wt%)
N C+N Cr-equivalent Ni-equivalent
0.005 0.010 12.95 0.68
~ ~ 12.48 0.75
0.010 0.025 12.20 0.93
13.03 0.48
Austenite volume
fraction before F1 (%)
~0
7
33
~0
s
Tab. 3
Chemical compositions of the four materials(wt%) and their austenite volume fraction observed before the entry to the
first pass, F1, of the hot–strip finishing mill. Ni-equivalent and Cr-equivalent are defined as Ni+0.5Mn+30C+0.3Cu+25N and
Cr+2.0Si+1.5Mo+5.5Al+1.5Ti, respectively.
Composizione chimica dei quattro materiali (% in peso) e loro frazione di austenite (in volume) rilevata dopo ingresso alla prima
passata, F1, nel laminatoio di finitura. Ni e Cr equivalenti definiti rispettivamente come Ni+0.5Mn+30C+0.3Cu+25N e
Cr+2.0Si+1.5Mo+5.5Al+1.5Ti.
s
Fig. 6
Microstructures of (a) the material M2 and
(b) M3. Martensitic phases (i.e., high temperature
austenitic phases) with dark contrast are elongated
along the rolling direction.
Microstrutture di (a) il materiale M2 e (b) materiale
M3. Con contrasto scuro le fasi martensitiche (ex fasi
austenitiche alle alte temperature) con contrasto scuro
sono allungate lungo la direzione di laminazione.
7(a)). The deformation resistances of the two-phase materials,
M2 and M3 are higher than that of the fully ferritic
material, M1, even though the rolling temperatures were
higher in the case of M2 and M3.
This unexpected behaviour can be understood by the presence
of the higher volume fraction of hard austenite in the
microstructure of the materials, M2 and M3.
An attempt was made to calculate the DRR values during
the first pass, F1, for the materials shown in Tab. 3. In order
to calculate the DRR, the dependence of deformation
resistance on the entry temperature must be known for the
fully ferritic material. Therefore, additional mill logs were
analysed. Twenty seven materials with six different chemical
compositions were selected for the investigation. The
applied strains were in the range of 0.45~0.50 and the strain
rates were between 7.5 to 9.5/sec. Thus, the rolling conditions
except for the temperature were almost constant. The
samples for the microstructural examination were taken
right before the entry to the first pass, F1, and all the microstructures
investigated using an optical microscope were
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Acciaio inossidabile
s
Fig. 8
The dependence of DRR as a function of (a) the measured austenite volume fraction and (b) the calculated
austenite potential. The line fit in Fig. 5(b) was superimposed to compare the DRRs from the experiments with those
from the mill log analyses.
Variazione del DRR in funzione di (a) frazione in volume dell’austenite misurata e (b) potenziale di austenite calcolato. La
linea inserita in Fig. 5(b) è stata aggiunta per confrontare i DRR delle sperimentali con quelli dedotti dall’ analisi dei documenti
di produzione.
resistance ratio (DRR) was defined as a ratio of the deformation
resistance of a two-phase microstructure to that of a
fully ferritic microstructure. The DRR appears to vary linearly
with the austenite volume fraction, thus implying that
the material is likely to deform in an equal-strain manner
along the hot rolling direction.
REFERENCES
1) William L. Roberts, Hot Rolling of Steels, Marcel Dekker
Inc., 1983, p.649
2) Vladimir B. Ginzburg and Robert Ballas, Flat Rolling
Fundamentals, Marcel Dekker Inc., 2000, p.199
3) Robert G. Nooning, Jr., Master Thesis, University of
Pittsburgh, 2002, p76.
4) T. M. Maccagno, J. J. Jonas, S. Yue, B. J. McCrady, R. Slobodian
and D. Deeks, ISIJ International, Vol. 34, 1994, No.
11, p.917.
5) T. M. Maccagno, J. J. Jonas and P. D. Hodgson, ISIJ International,
Vol. 36, 1996, No. 6, p.720.
6) F. Siciliano Jr., K. Minami, T. M. Maccagno and J. J. Jonas,
ISIJ International, Vol. 36, 1996, No. 12, p.1500.
s
Fig. 9
The dependence of DRR during a first rolling pass, F1, as a function of (a) the calculated austenite potential
and (b) the Cr-equivalent & Ni-equivalent. Three dimensional data have been projected on the two dimensional planes
in Fig. 8(b).
La dipendenza del DRR durante il primo passaggio di laminazione, F1, in funzione di (a) potenziale di austenite calcolato e (b)
Cr e Ni equivalenti. I dati tridimensionali sono stati proiettati sui piani bidimensionali di Fig. 8(b).
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Acciaio inossidabile