DETERMINATION OF THIN FILM'S MECHANICAL PROPERTIES

1
DOKUZ EYLUL UNIVERSITY
GRADUATE SCHOOL OF NATURAL AND APPLIED SCIENCES
DETERMINATION OF THIN FILM’S
MECHANICAL PROPERTIES
by
Osman ÇULHA
December, 2006
İZMİR

2
DETERMINATION OF THIN FILM’S
MECHANICAL PROPERTIES
A Thesis Submitted to the
Graduate School of Natural and Applied Sciences of Dokuz Eylul University
In Partial Fulfillment of the Requirements for the Degree of Master of Science
in Metallurgical and Materials Engineering, Metallurgical and Materials
Program
by
Osman ÇULHA
December, 2006
İZMİR

3
M. Sc. THESIS EXAMINATION RESULT FORM
We have read the thesis entitled “DETERMINATION OF THIN FILM’S
MECHANICAL PROPERTIES” completed by OSMAN ÇULHA under revision of
PROF. DR. TEVFİK AKSOY and we certify that in our opinion it is fully adequate, in
scope and in quality, as a thesis for the degree of Master of Science.
Prof. Dr. Tevfik AKSOY
Supervisor
(Jury Member)
(Jury Member)
Prof.Dr. Cahit HELVACI
Director
Graduate School of Natural and Applied Science
ii

4
ACKNOWLEDGEMENTS
I sincerely thank for the people who mentally support and encourage me, aid me
in my pursuing of the M. Sc. degree, and help in my academic accomplishment.
Firstly, I would like to thank Prof. Dr. Tevfik AKSOY for his supervision,
guidance, patience, and support in this work.
I also would like to thank Assoc. Prof. Dr. Mustafa TOPARLI and my all
colleagues especially Bahadır UYULGAN, Faruk EBEOGLUGİL and Erhan
ÖZKAN for their cooperation, friendship and patience.
Finally, I would like to thank my all family and my engaged Aslı ŞIK for their
support and persistence.
Osman ÇULHA
iii

5
DETERMINATION OF THIN FILM’S MECHANICAL PROPERTIES
ABSTRACT
The surface of industrial component may require treatment to enhance surface
characteristic. Surface treatments that cause microstructure changes in the bulk
material include heating and cooling/quenching through induction, flame, laser and
electron beam techniques, or mechanical treatments.
Boriding is one of the thermo chemical surface treatment processes which is
extensively used for ferrous and nonferrous materials. In this study, characterization
and mechanical properties of FeB and Fe 2 B phases, which were formed on surface of
SAE 1020 and SAE 1040 quality steel by boriding technique, were investigated.
The aim of this study was to obtain micro structural characteristics and investigate
the mechanical properties such as hardness, young’s modulus and fracture toughness
of the boride layer. The produced double layer (FeB and Fe 2 B) was extensively
analyzed with respect to X-ray diffraction (XRD), scanning electron microscopy
(SEM) including energy dispersive spectroscopy (EDS), microhardness and surface
roughness. Mechanical properties of layers was examined by Shimadzu Dynamic
Ultra-micro Hardness test machine for estimating young’s modulus due to loadunload
sensing analysis, in addition to mechanical investigation hardness-depth
curves of the layer was obtained. Fracture toughness properties of surface layer FeB
was calculated by Vickers Fracture Toughness method with measuring crack length
after loading stage is finished.
Finite element modeling was performed to determine the yield stress and stress
analysis of borided substrates.
Keywords: Boriding, micro-indentation, Fracture toughness, Finite Element Method
iv

6
İNCE FİLMLERİN MEKANİK ÖZELLİKLERİNİN BELİRLENMESİ
ÖZ
Endüstride kullanım alanına sahip birçok malzeme yüzey özelliklerinin
iyileştirilmesine ihtiyaç duymaktadır. Yüzey iyileştirme işlemleri, yüzeyin mikro
yapısal özelliklerini termokimyasal yöntemle (ısıtma ve soğutma işlemleriyle)
değiştirmeye yöneliktir. Başlıca ısıtma işlemleri alevle, lazerle ve elektron beam
tekniklerini içermekte, soğutma ise su verme, havada ve yağda soğutma işlemlerini
kapsamaktadır.
Borlama işlemi, demir ve demir dışı metallere uygulanabilen termokimyasal
yüzey işlemlerindendir. Bu çalışmada, SAE 1020 ve SAE 1040 kalite çelikler
borlama işlemine tabi tutularak oluşturulan FeB ve Fe 2 B fazlarının karakterizasyonu
ve mekanik özelliklerinin belirlenmesi hedeflenmiştir.
Bu çalışmanın amacı, borür tabakalarının mikro yapısal özelliklerinin yanı sıra
sertlik, elastisite modülü ve kırılma tokluğu gibi mekanik özelliklerinin
belirlenmesidir. Üretilen FeB ve Fe 2 B içeren çiftli tabaka, taramalı elektron
mikroskobu ve X ışınları difraksiyon cihazı, mikro sertlik ve yüzey pürüzlülüğü
cihazları ile detaylı olarak incelenmiştir. Tabakaların mekanik özelliklerinden olan
elastisite modülü ve sertlik-derinlik değişimi Shimadzu Dynamic Ultra-mikro
Hardness test cihazı kullanılarak belirlenmiştir. FeB tabakasına ait kırılma tokluğu
değeri ise Vickers kırılma tokluğu deney prosedürü uygulanarak elde edilmiştir.
Sonlu elemanlar metodu ise FeB tabakasına ait akma gerilmesi tayini ve gerilme
analizi yapmak için kullanılmıştır.
Anahtar Kelimeler: Borlama, Mikro Indentation, Kırılma Tokluğu, Sonlu Elemanlar
metodu
v

7
CONTENTS
Page
THESIS EXAMINATION RESULT FORM .............................................................. ii
ACKNOWLEDGEMENTS ........................................................................................iii
ABSTRACT................................................................................................................ iv
ÖZ ................................................................................................................................ v
CHAPTER ONE – INTRODUCTION .................................................................... 1
CHAPTER TWO – BORONIZING ……………………………………………….4
2.1 Introduction……………………………………………………………………4
2.2 Characteristic Features of Boride Layers……………………………………...4
2.3 Advantages of Boriding……………………………………………………….6
2.4 Boriding of Ferrous Materials……………………………………………….10
2.5 Characteristics of FeB and Fe 2 B Layers……………………………………..10
2.6 The Boriding Process………………………………………………………...12
2.7 Ferrous Materials for Boriding…………………………………………….....13
2.8 Influence of Alloying Elements……………………………………………...13
2.9 Heat Treatment after Boriding……………………………………………….15
vi

8
2.10 Boriding of Nonferrous Materials……………………………………….….15
2.11 Effect of Alloying Elements ……………………………………………….16
2.12 Thermo chemical Boriding Techniques ……………………………………16
2.12.1 Pack Boriding ……………………………………………………….16
2.12.1.1 Case Depth…………………………………………………18
2.12.1.2 Borudif Process………………………………………….…20
2.13 Applications of Thermo chemical Boriding…………………………….….20
CHAPTER THERE – MECHANICAL PROPERTIES OF MATERIALS.......23
3.1 Hardness Test……………………………………………………………….....23
3.1.1 Vickers Hardness Test…………………………………………………..23
3.1.2 Rockwell Hardness Test……………………………………………...…24
3.1.3 Brinell hardness and Meyer hardness………………………………...…25
3.2 Determination of Young’s modulus by load-depth sensing indentation……...26
3.3 Determination of Elastic modulus by Vickers indentation………………...….30
3.4 Correlation between yield strength and hardness………………………….….32
3.5 Vickers Indentation Cracks and Fracture Toughness……………………...….33
CHAPTER FOUR – EXPERIMENTAL STUDIES…………………………..…35
4.1 Purpose…………………………………………………………………….…35
4.2 Materials…………………………………………………………………..….36
vii

9
4.3 Characterization Studies……………………………………………….……..36
4.3.1 X-Ray Diffractions (XRD)……………………………………………..36
4.3.2 Scanning Electron Microscopy / (EDS)…………………………….….36
4.4 Surface Roughness Test………………………………………………...……37
4.5 Dynamic Ultra-micro Hardness Test DUH-W201/W201S………………..…37
4.6 Finite Element Modeling……………………………………………….…….37
CHAPTER FIVE – RESULTS AND DISCUSSION……………………………38
5.1 Characterization of the Layers………………………………………………38
5.1.1 XRD…………………………………………………………………..38
5.1.2 SEM and Thickness examination……………………………………..39
5.2 Surface Roughness Measurement Layer…………………………………….46
5.3 Ultra Micro Hardness Examination of Layer………………………………..46
5.4 Fracture Toughness Measurement…………………………………………...59
5.5 FEM Model…………………………………………………………………..72
CHAPTER SIX – CONCLUSION………………………………………………..80
6.1 General Results………………………………………………………………80
6.2 Future Plans…………………………………………………………………..82
REFERENCES…………………………………………………………….……….83
viii

1
CHAPTER ONE
INTRODUCTION
Although well known for approximately 100 years, classical hardness test has
gained new popularity during the last decades.
The advantage of the indentation test, in comparison with a uniaxial tensile test, is
of course the relative simplicity of the experimental setup. On the other hand, an
obvious drawback is the very complicated mechanical problem arising owing to
inelastic and/or inhomogeneous deformation in the indented materials. Therefore,
until recently the interpretation of indentation tests have relied heavily on semiempirical
formulae. The work by Tabor is perhaps the best example of this, with no
or little theoretical foundation. With the advent of modern computers and advanced
numerical methods, however, the understanding of the mechanics involved during
ball indentation (Hill et al., 1989; Kral et al., 1993; Larsson, 1994), cone indentation
(Laursen and Simo, 1992) and Vickers Indentation (Giannakopoulos and Suresh,
1994), has increased rapidly in recent years.
These new interests in the mechanical behavior of indentation testing are to a
large extent a result of the increased use of new materials. These materials are
notoriously difficult to characterize through uniaxial or other standard tests, which in
many cases make indentation testing the only possible alternative for determining
their mechanical properties.
The qualities of small size and non-destructive test capability make the
indentation technique superior to the tension test. For very small volumes of
material, the uniaxial test is inapplicable. Furthermore, the structural materials may
not be removed to do the tension test in most cases, for instance, the materials used
in the electronic solders or engineering welds. Indentation technique can evaluate the
material properties while keeping the structural integrity.
1

2
Sharp indentation tests also serve to initiate and control fracture in brittle
materials. Much effort has been devoted to fracture toughness determination. Other
applications that refer to, or use indirectly, the Vickers hardness include abrasive
wear problem, armor plating, machining, etc. Most tests are of static type; however,
several dynamic tests are also performed.
Another main application area of indentation technique lies in the thin film –
substrate systems research. Thin film coatings have been widely used in various
industries in recent decades. The coated products have become crucial to machine
parts performance and durability, multiplying the life span of some parts by 10 to 20
times.
For instance, the mechanical properties of FeB and Fe 2 B layers are important for
industrial application. Pack boriding process was applied to SAE 1020 and SAE
1040 quality steel for achieving advantages such as:
• Hardness of the boride layer can be retained at higher temperatures than, for
example, that of nitrided cases.
• A wide variety of steels, including through-hardenable steels, are compatible
with the processes (Mater. Eng, 1970)
• Boriding, which can considerably enhance the corrosion-erosion resistance of
ferrous materials in nonoxidizing dilute acids as indicated in chapter two Fig.
2.4 and alkali media, is increasingly used to this advantage in many industrial
applications (W.J.G. Fichtl).
• Borided surfaces have moderate oxidation resistance (up to 850) and are quite
resistant to attack by molten metals.
• Borided parts have an increased fatigue life and service performance under
oxidizing and corrosive and corrosive environment.
After pack boriding process, multilayer diffusion phases (FeB and Fe 2 B)
formed on steel substrate. The structural compositions of layers consist of boron
rich phase (FeB) and iron rich phase (Fe 2 B), respectively. FeB formation begins

3
from surface, Fe 2 B phase forms in deeper region because of decreasing boron
concentration form surface to inside of substrate. Upper part of multilayer
consists of FeB phase which is more brittle than lower layer (Fe 2 B).
Mechanical properties (Young’s modulus, Hardness, Yield strength) of FeB
layer are difficult to characterize through uniaxial or other standard tests. So,
micro indentation method became useful for determination of mechanical
properties of layers.

4
CHAPTER TWO
BORONIZING
2.1 Introduction
Boriding, or boronizing, is a thermo chemical surface hardening process that can
be applied to a wide variety of ferrous, nonferrous, and cermet materials. The process
involves heating well-cleaned material in the range of 700 to 1000 o C, preferably for
1 to 12 h, in contact with boronaceous solid powder (boronizing compound), paste,
liquid or gaseous medium. Other developments in the thermo chemical boriding
include gas boriding techniques such as plasma boriding and fluidized bed boriding.
There is a current trend toward the use of multicomponent boriding (Asm Handbook,
vol.4, 2000).
This article presents mainly various media used for thermo chemical boriding,
their advantages, limitations and applications. Physical and chemical vapor
depositions, plasma spraying and ion implantation are alternative nonthermochemical
surface coating process for the deposition of boron or codepozition of boron and
metallic elements onto suitable metallic or nonmetallic substrate material (Asm
Handbook, vol.4, 2000).
2.2 Characteristic Features of Boride Layers
During boriding, the diffusion and subsequent absorption of boron atoms into
metallic lattice of component surface form interstitial boron compounds (A.G. von
Matuschka, 1980). The resulting layer may consist of either a single phase boride or
polyphase boride layer. The morphology, growth and phase composition of boride
layer can be influenced by alloying element in the base material as represented in Fig
2.1 (R. Chatterjee and Fischer, 1981). The micro-hardness of the boride layer also
depends strongly on the composition and structure of the boride layer and
composition of the based material as shown in Table 2.1 (Galibois et al, 1980).
4

5
Figure 2.1 Effect of steel composition on the morphology and thickness of the boride layer
(Asm Handbook, vol.4, 2000).
Table 2.1 Melting points and microhardness of different boride phases formed
during boriding of different substrate materials (Asm Handbook, vol.4, 2000)

6
2.3 Advantages of Boriding
Boride layer posses a number of characteristic features with special advantages
over conventional case-hardened layers. One basic advantage is that boride layers
have extremely high hardness values (between 1450 and 5000 HV) with high melting
points of constituent phases as demonstrated in Table 2.1. The typical surface
hardness values of borided steels compared with other treatments and other hard
materials are listed in Table 2.2. This clearly illustrates that the hardness of boride
layers produced on carbon steels is much greater than that produced by any other
conventional surface treatments; it exceeds that of the hardened tool steels, hard
chrome electroplate and is equivalent to WC (Asm Handbook, vol.4, 2000).
Table 2.2 Typical surface hardness values of borided steels compared
with other treatments and hard materials (Asm handbook, vol. 4, 2000)
The combination of hard surface hardness and a low surface coefficient of friction
of the boride layer also make significant contribution in combating the main wear
mechanism: adhesion, tribooxidation abrazyon and surface fatigue (W.J.G. Fichtl;
1983, K.H. Habing and R.Chatterjee-Fische; 1981). This fact has enabled the mold

7
makers to substitute easier-to-machine steels for the base metal and to steel obtain
wear resistance and ant galling properties superior to those of the original material
(D.J. Bak, 1981). Figure 2.2 shows the effect of boriding on abrasive wear resistance
of borided C45 steel, titanium and tantalum as a function of number of revolutions
(or stressing period) based on wear test. Figure 2.3 shows the influence of steel
composition on abrasive wear resistance.
Figure 2.2 Effect of boriding on the wear resistance. a) 0.45 % C (C45) steel borided 900 o C.
b)Titanium borided at 1000 o C for 24h c) Tantalum borided at 1000 o C for 8h (Asm Handbook
vol. 4, 2000)
Figure 2.3 Effect of steel composition on wear resistance under
abrasive wear. Test conditions: DP-U grinding tester, SiC paper
220, testing time 6 min. (Asm handbook, vol.4, 2000).

8
Other advantages of boriding include:
• Hardness of the boride layer can be retained at higher temperatures than, for
example, that of nitrided cases.
• A wide variety of steels, including through-hardenable steels, are compatible
with the processes (Mater. Eng, 1970)
• Boriding, which can considerably enhance the corrosion-erosion resistance of
ferrous materials in nonoxidizing dilute acids as indicated in Fig. 2.4 and
alkali media, is increasingly used to this advantage in many industrial
applications (W.J.G. Fichtl).
• Borided surfaces have moderate oxidation resistance (up to 850) and are quite
resistant to attack by molten metals.
• Borided parts have an increased fatigue life and service performance under
oxidizing and corrosive and corrosive environment.
Figure 2.4 Corroding effects of mineral acids on boronized and nonboronized steels
(Asm handbook, vol. 4, 2000).

9
Disadvantages of boronizing treatment are:
• The techniques are inflexible and rather labor intensive, making the process
less cost effective than other thermo-chemical surface hardening treatments
such as gas carburizing and plasma nitriding. Both gas carburizing and
plasma nitriding have the advantage over boronozing because those two
processes are flexible systems, offer reduced operating and maintenance cost,
required shorter processing time, and are relatively easy to operate. It is
therefore, suited to engineering components that need high hardness and
outstanding wear and corrosion resistance of the boride layers, and/or where
cheaper labor is available (P. Dearnly, T. Bell, 1985).
• The layers thickness magnitude depends on the base material composition but
remains consistent for a given combination of material and treatment cycle.
However, it can be predicted for given part geometry and boronozing
treatment. For treatment of precision parts, where little stock removal is
permitted, an allowance of 20 to 25 % dimensional increase of final boride
layer thickness must be provided.
• Partial removal of the boride layer for closer tolerance requirements is made
possible only by a subsequent diamond lapping because conventional
grinding causes fracture of the layer. Thus, precise boronozing is mostly
practiced for components with large cross-sectional area (P. Dearnly, T. Bell,
1985).
• Boriding of most steels provides a marginal increase, if any, in the bending
fatigue endurance limit, although some improvement in the corrosion fatigue
strength has been noticed.
• In general, the rolling contact fatigue properties of borided alloy steel parts
are very poor compared to carburized and nitrided steel at high contact loads.
This is why boronozing treatment of gears is limited to those screw designs
where transverse loading of gear teeth is minimized (P. Dearnly, T. Bell,
1985).
• There is frequently a need to harden and temper the tool after boriding (H.C.
Child, 1981), which requires a vacuum or inert atmosphere to preserve the
integrity of the borided layer.

10
2.4 Boriding of Ferrous Materials
Unlike carburizing treatment on ferrous materials, where there is gradual
decrease in composition from the carbon-rich surface to the substrate, the boriding of
ferrous materials results in the formation of either a single-phase or double-phase
layer of borides with definite compositions. The single phase boride layers consist of
Fe 2 B, while the double-phase layer consist of an outer dark etching phase FeB and
inner bright etching phase Fe 2 B. The formation of either a single or double phase
depends on the availability of boron (R. Chatterjee-Fischer, 1977).
2.5 Characteristics of FeB and Fe 2 B Layers
The formation of a single Fe 2 B phase (with saw tooth morphology due to
preferred diffusion direction) is more desirable than a double-phase layer with FeB.
The boron rich FeB phase is considered undesirable, in part because FeB is more
brittle than the iron sub boride Fe 2 B layer. Also, because FeB and Fe 2 B formed under
tensile and compressive residual stress, respectively, crack formation is often
observed at or in the neighborhood of the FeB / Fe 2 B interface of double-phase layer.
These cracks may lead to flaking and spalling when a mechanical strain is applied
(Galibois et al, 1980) or even separation, as shown in Fig.2.5, when a component is
undergoing a thermal and/or mechanical shock. Therefore, the boron-rich FeB phase
should be avoided or minimized in the boride layer (Galibois et al, 1980).
Figure 2.5 Separation of two-phase boride layer
on low carbon (St 37) steel (borided at 900 o C, 4 h
and 200x), (Asm Handbook, vol.4, 2000).

11
It has also been reported that the tribological properties depend on the
microstructure of the boride layer. The dual-phase FeB-Fe 2 B layers are not inferior to
those of monophase Fe 2 B layers, provided that the porous surface zone directly
beneath the surface is removed (W. Liluental et al, 1983). Alternatively, a thinner
layer is favored because of less development of brittle and porous surface zone
formation and flaking.
Typical properties of the FeB phase are:
• Microhardness of about 19 to 21 GPa.
• Modulus of elasticity of 590 GPa.
• Density of 6.75 g/cm 3 .
• Thermal expansion coefficient of 23 ppm/ o C
• Composition with 16.23 wt % boron.
• Orthorhombic crystal structure with 4 iron and 4 boron atoms per unit cell
• Lattice parameters: a= 4.053 A o , b= 5.495 A o and c= 2.946 A o
The formation of single phase Fe 2 B layers with a saw tooth morphology is
desirable in the boriding of ferrous materials (D.N. Tsipas et al, 1988). A single Fe 2 B
phase can be obtained from a double FeB-Fe 2 B phase by a subsequent vacuum or salt
bath treatment for several hours above 800 o C, which may be followed by oil
quenching to increase substrate properties (P.A. Dearnly et al, 1986).
Typical properties of Fe 2 B are:
• Microhardness of about 18 to 20 GPa
• Modulus of elasticity of 285 to 295 GPa
• Thermal expansion coefficient of 7.65 ppm/ o C
• Density 7.43 g/cm 3
• Composition with 8.83 wt% boron
• Body-centered tetragonal structure with 12 atoms per unit cell
• Lattice parameters: a=5.078 A o , c= 4.249 A o
The solubility of boron in ferrite and austenite is very small (< 0.008 % at 900
o C) according to the Fe-B phase diagram (T.B. Massalski, 1986). According to the
Brown and Nicholson (M.E. Nicholson, 1954), the phase diagram exhibits an / γ /

12
Fe 2 B peritectoid reaction at about 912 o C, based on their findings of higher solubility
of boron in ferrite than in austenite at the reaction temperature. However, later work
revealed a higher solubility of boron in austenite compared to that in ferrite, thereby
suggesting the reaction is eutectoid in nature (T.b. Cameron and J.EMorral, 1986).
2.6 The Boriding Process
The boriding process consists of two types of reaction (R.Chatterjee-Fischer,
1989). The first reaction takes place between the boron-yielding substance and the
component surface. The nucleation rate of the particles at the surface is a function of
the boriding time and temperature. This produces a thin, compact boride layer.
The subsequent second reaction is diffusion controlled, and the total thickness of
the boride layer growth at a particular temperature can be calculated by the simple
Formula:
d = k t
Where d is the boride layer thickness in centimeters; k is a constant, depending on
the temperature, and t is the time in seconds at given temperature. The diffusivity of
boron at 950 o C is 1.82x10 –8 cm 2 /s for the boride layer and 1.53x10 –7 cm 2 /s for the
diffusion zone. As a result, the boron-containing diffusion zone extends more than 7
times the depth of boride layer thickness into the substrate (H. Kunst et al, 1967). It
has been proposed that a concentration gradient provides the driving force for
diffusion-controlled boride layer growth (M.J Lu, 1983).
Diffusion case thicknesses range to approximately 0.13 mm for ferrous alloys,
depending on alloy compositions and configurations. A lower case depth is required
for the high-carbon and/or high-alloy tool steels, whereas higher case depths may be
needed for the low-or medium-carbon steels. When case depth is about 320 to 350
µm, subsequent heat treatment is not performed.

13
2.7 Ferrous Materials for Boriding
With the exception of aluminum-and silicon-bearing steels, industrial boriding
can be carried out on most ferrous materials such as structural steels; case hardened,
tempered, tool, and stainless steels; cast steels; Armco (commercially pure) iron;
gray and ductile cast irons; and sintered iron and steel (W. Fitchl, 1972). Because
boriding is conducted in the austenitic range, air-hardening steels can be
simultaneously hardened and borided. Water-hardening grades are not borided
because of the susceptibility of the boride layer to thermal shock. Similarly,
resulfirized and leaded steels should not be used because they have tendency toward
case spalling and case cracking.
2.8 Influence of Alloying Elements
The mechanical properties of borided alloys depend strongly on the composition
and structure of the boride layers. The characteristic saw tooth configuration of the
boride layer is dominant with pure iron, unalloyed low-carbon steels, and low alloy
steels. As the alloying element and/or carbon content of the substrate steel is
increased, the development of a jagged boride/substrate interface is suppressed, and
for high alloy steels a smooth interface is formed (A.J. Ninham and I.M. Hutchings,
1989 and Fig.2.1). Alloying element mainly retard the boride layer thickness (or
growth) caused by restricted diffusion of the formation of diffusion barrier.
Carbon does not dissolve significantly in the boride layer and does not diffuse
through the boride layer. During boriding, carbon is driven from the boride layer to
the matrix and forms, together with boron, boroncementite Fe 3 (B, C) as a separate
layer between Fe 2 B and the matrix (J.J. Smit, 1984C.M. Brakman, 1988).
Like carbon, silicon and aluminum are not soluble in the boride layer, and these
elements are pushed from the surface by boron and are displaced ahead of the boride
layer into the substrate, forming ironsilicoborides (FeSi 0.4 B 0.6 ) and Fe 3 SiB 2
underneath the Fe 2 B layer. Steels containing high contents of these ferrite forming

14
elements should not be used for boriding because they reduce the wear resistance of
normal boride layer (A.G. von Matuschka, 1980); they produce a substantially softer
ferrite zone beneath the boride layer than that of the core (H.C. Fielder and W.J.
Hayes, 1970). At higher surface pressure, this type of layer buildup results in the socalled
egg shell effect, that is at greater thickness extremely hard and brittle boride
layer penetrates into the softer intermediate layer and is consequently destroyed
(W.J.G. Fichtl, 1989).
A reduction in both the degree of interlocking tooth structure and of boride depth
can occur with high nickel containing steels. Nickel has been found to concentrate
below the boride layer; it enters the Fe 2 B layer and in some instances promotes the
precipitation of Ni 3 B from FeB layer (K.H. Habig, 1981). It also segregates strongly
to the surface from the underlying zone corresponding to the Fe 2 B layer. This is quite
pronounced in both Fe-14Ni and austenitic stainless steels. Consequently, gas
boronozing of austenitic stainless steel appears to be more appropriate treatment for
producing a low porosity, homogeneous, single-phase Fe 2 B layer because of the
lower boron activity of the gaseous mixture (P. Goeurits, 1982).
Chromium considerably modifies the structure and properties of iron borides. As
the chromium content in the base material increases, progressive improvements in
the following effects are observed:
• Formation of boron rich reaction products
• Decreased in boride depth
• Flattening or smoothening of the coating/substrate interface (M. Garbuccchio,
1985).
A reduction of boride thickness has also been noticed in ternary Fe-12Cr-C steels
with increasing carbon content (M. Garbuccchio, 1985). Manganese, tungsten,
molybdenum and vanadium also reduce the boride layer thickness and flatten out the
tooth-shaped morphology in carbon steels. The distribution of titanium, cobalt, sulfur
and phosphorus in the boride layer has not been well established.

15
2.9 Heat Treatment after Boriding
Borided parts can be quench hardened in air, oil, and salt bath or polymer
quenchant and subsequently tempered. Heating to the hardening temperature should
be carried out in an oxygen-free protective atmosphere or in a neutral salt bath
(R.Chatterjee-Fischer, 1989)
.
2.10 Boriding of Nonferrous Materials
Nonferrous materials such as nickel, cobalt and molybdenum based alloys as well
as refractory metals and their alloys and cemented carbides can be borided. Copper
cannot: therefore, it provides a good stopping-off material. Of special interest is the
boriding of nickel alloys and titanium and its alloys. Usually, boriding of nickel plate
is done in gaseous BCl 3 -H 2 -Ar mixture in the temperature range of 500 to 1000 o C
(S. Motojima et al, 1981), whereas Permalloy is pack borided with 85% B 4 C and
15% Na 2 CO 3 , or 95 % B 4 C and 5 % Na 2 B 4 O 7 powder mixture at 1000 o C for 6 h in
H 2 atmosphere.
Boriding of titanium and its alloys is carried out preferably between 1000 and
1200 o C. Here pack boriding in oxygen-free amorphous boron in combination with
high vacuum (0.0013 Pa, or 10 –5 torr) and high purity argon atmosphere or gas
boriding with H 2 -BCl 3 -Ar gas mixture is preferred. The microhardness readings of
boride layers formed on titanium and refractory metals are very high compared to
those formed on cobalt and nickel (see Table 2.1). The wear properties of sintered
carbides can be increased by boriding because of the acceptance of boron by soft
cobalt and nickel binders (Lindberg heating treating company, “Boroalloy Process”).
Boride layers formed on tantalum, niobium, tungsten, molybdenum and nickel do
not exhibit tooth-shaped morphology as with titanium. When cemented carbides such
as WC-Co are commercially pack borided with a powder mixture containing 40%
B 4 C, 45% SiC and 5% KBF 4 , three distinct zone are found to be formed in the boride

16
layer, comprising the exterior (zone I), intermediate (zone II), and interior (zone III)
regions as represented in Table 2.3.
Table 2.3 Three distinct zones formed in borided cemented carbide
materials (Asm Handbook, vol.4, 2000)
2.11 Effect of Alloying Elements
As in iron and steel, suitable alloying additions raise the hardness of boride layer
formed on these metals: this is presumably caused by the formation of solid-solution
borides.
Additions of alloying elements in nickel, cobalt and titanium retard the rate of
boride layer growth and in the case of multiphase layers, proportion of boride layer
with high boron content (for example TiB 2 in titanium) increases. The tooth-shape
morphologies in the cases of cobalt and titanium are also retarded with alloying
addition, and the layers appear more uniform ( R. Chatterjee et al, 1976).
2.12 Thermo chemical Boriding Techniques
Because extensive investigations have been carried out on boriding of ferrous
materials, the techniques described below focus on mainly on the boriding of ferrous
materials.
2.12.1 Pack Boriding
Pack boriding (V.A. Volkov et al, 1975 and N. Komutsu et al, 1974) is the most
widely used boriding process because of its relative ease of handling, safety and the
possibility of changing the composition of the powder mix, the need for limited

17
equipment, and the resultant economic savings (A.G. von Matuschka, 1980). The
process involves packing the annealed, cleaned, smooth parts in a boriding powder
mixture contained in a 3 to 5 mm thick, heat resistance steel box so that surfaces to
be borided are covered with an approximately 10 to 20 mm thick layer. Many
different boriding compounds have been used for packing boriding. They include
solid boron-yielding substances, diluents and activators.
The common boron-yielding substances are boron carbide (B 4 C), Ferro boron and
amorphous boron: the last two have greater boron potential, provide a thicker layer,
and are more expensive than B 4 C (N. Komutsu et al, 1974). Silicon carbide (SiC) and
alumina (Al 2 O 3 ) serve as diluents, and they do not take part in the reaction. However,
SiC controls the amount of boron and prevents the caking of the boronozing agent.
NaBF 4 , KBF 4 , (NH4) 3 BF 4 , NH 4 Cl, Na 2 CO 3 , BaF 2 and Na 2 B 4 O 7 are the boriding
activators. There are special proprietary brands of boriding of boriding compounds,
such as different grades of Ekabor, available on the market that can be used with
confidence.
Typical compositions of commercial solid boriding powder mixtures are:
• 5% B 4 C, 90% SiC, 5% KBF 4
• 50 % B 4 C, 45 % SiC, 5% KBF 4
• 85 % B 4 C, 15 % Na 2 CO 3
• 95 % B 4 C, 5 % Na 2 B 4 O 7
• 84 % B 4 C, 16 % Na 2 B 4 O 7
• Amorphous boron (containing 95 to 97 % B)
• 95 % Amorphous boron, 5 % KBF 4
The parts conforming to the shape of the container are packed, covered with a lid,
which rests inside the container and is weighted with an iron slug or stone to ensure
an even trickling of the boriding agent during the boriding treatment, as shown in
Fig. 2.7. It is than heated to the boriding temperature in an electrically heated box or
pit furnace with open or covered heating coils or a muffle furnace for a specified

18
time. The container should not exceed 60 % of the furnace chamber volume. In
principle, boriding should be accomplished in such a way that high internal stresses
are relieved, which in turn, eliminates cracks or spalling. With the packing process,
the powder may be reused by blending with 20 to 50 wt% of fresh powder mixture.
In this case, the powder should be discarded after 5 or 6 cycles (Asm Handbook,
vol.4, 2000).
Figure 2.7 Diagram of packing of a single
geometrical part in pack boriding box (A.G.
von Matuschka, 1980).
2.12.1.1 Case Depth
Figure 2.8 and Figure 2.9 show that the thickness of the boride layer depends on
the substrate material being processed, boron potential of the boronozing compound,
boronizing temperature, and time. In ferrous materials, the heating rate especially

19
between 700 o C and the boriding temperature (800 to 1000 o C) should be high in
order to minimize the formation of FeB (R. Chatterjee, 1989).
Figure 2.8 Diagram showing the influence of the B 4 C
content of the boriding powder on the Proportion of
FeB phase in the boride layer of various steels borided
with pack powder at 900 o C (R. Chatterjee, 1976).
Figure 2.9 Effect of pack boriding temperature and time
on the boride layer thickness in C45 (Asm Handbook,
vol.4, 2000)
It is usual practice to match the case depth with the intended application and base
material. As a rule, thin layers (15 to 20 µm) are used for protection against adhesive

20
wear, whereas thick layers are recommended to combat erosive wear. The commonly
produced case depths are 0.05 to 0.25 mm for low alloy and low carbon steels and
0.025 to 0.076 mm for high alloy steels. However, case depths>0.089 mm are
uneconomical for highly alloyed materials such as stainless steels and some tool
steels (Lindberg heating treating company, “Boroalloy Process”).
2.12.1.2 Borudif Process
In another modified pack boriding treatment, called the Borudif Process, steel
parts are packed in a 1:4 mixture of B 4 C-SiC, and the moderate activator, BF 3 plus
(BOF) 3 gas is passed through the pack at 850 to 1000 o C for 4 h (ref 43). The process
offers a wide range of boriding potential because of the easy control of (BOF) 3 gas
concentration that facilitates the treatment of a wide variety of substrate materials (P.
Goeuriot et al, 1981).
2.13 Applications of Thermo chemical Boriding
Table 2.5 shows that presently borided parts have been used in a wide variety of
industrial applications, because of the numerous advantages properties of boride
layers. In sliding and adhesive wear situation, boriding is applied to:
• Spinning steel rings, steel rope and steel thread guide bushings
• Grooved gray cast iron drums for textile machinery.
• Four-holed feed water regulating valves
• Burner nozzles swirl elements, and injector tops for steel oil burners in the
chemical industry.
• Drive, worm, and helically toothed steel gears in various high-performance
vehicle and stationary engines.

21
Table 2.5 Proven applications for borided ferrous materials (Asm Handbook Vol. 4, 2000)
As abrasive wear resistant materials, borided stainless steel are used for parts
such as screw cases and bushings, perforated and slotted hole screens, rollers, valve
components, fittings, guides, shafts, and spindles, and borided Ti-6Al-4V for parts
such as leading edge rotor blade cladding for helicopter applications in this category
include:
• Nozzles of bag filling equipment.
• Extrusion screws, cylinders, nozzles, and reverse-current blocks in plastic
production machinery.

22
• Bends and baffle plates for conveying equipment for mineral-filled plastic
granules in the plastics industry.
• Punching dies , pres and drawing matrices, and necking rings
• Pres dies, cutting templates, punched plate screens
• Screw and Wheel gears, bevel gears
• Steel molds, extruder barrels, plungers and rings.
• Extruder tips, nonreturn valves and cylinders.
• Casting fillers for processing nonferrous metals

23
23
CHAPTER THREE
MECHANICAL PROPERTIES OF MATERIALS
3.1 Hardness Test
Indentation tests, in many cases referred to as hardness tests, have for a long time
been a standard method for material characterization. The hardness test consists of
loading an indenter made of diamond or any other hard material (e.g., Tungsten
Carbide) and pressing it into the surface of a softer material to be examined. The
further into the material the indenter sinks (for a given load), the softer the material is
and the lower its yield strength. Hardness is not an intrinsic material property
dictated by precise definitions in terms of fundamental units of mass, length and
time. A hardness property value is the result of a defined measurement procedure.
Hardness tests are the most commonly used non-destructive testing procedures in the
metal industry and in research because they provide an easy, inexpensive and reliable
method of evaluating basic properties of developed or new materials. The hardness
test indenter is so small that it scarcely damages the bulk material; therefore, it can
be used for routine batch tests on small samples of materials to ascertain that they are
up to specifications on yield without damaging them. The usual method to achieve a
hardness value is to measure the depth or area of an indentation left by an indenter of
a specific shape, with a specific force applied for a specific time. There are three
principal standard test methods for expressing the relationship between hardness and
the size of the impression, these being Vickers, Rockwell and Brinell (Ziheng Yao,
2005).
3.1.1 Vickers Hardness Test
The Vickers indenter is a square based pyramid with an angle of 136 degrees
between the faces and a ratio of diagonals of 1:1 (as shown in Figure 3.1). The
Vickers hardness number is one of the most widely used measures of hardness in
engineering and science. The Vickers diamond hardness, VDH, is calculated using

24
the indenter load and the actual surface area of the impression. The resulting quantity
is usually expressed in kgf/mm 2 (Ziheng Yao, 2005).
F= Load in kgf
d = Arithmetic mean of the two diagonals, d 1 and d 2 in mm (Figure 3.1)
HV = Vickers hardness
Figure 3.1 Scheme of Vickers hardness test and Vickers impression
3.1.2 Rockwell hardness test
The Rockwell hardness test method consists of indenting the test material with a
diamond cone or hardened steel ball indenter. The indenter is forced into the test
material under a preliminary minor load F 0 usually 10 kgf. When equilibrium has
been reached, an indicating device, which follows the movements of the indenter and

25
so responds to changes in depth of penetration of the indenter, is set to a datum
position. While the preliminary minor load is still applied an additional major load is
applied with resulting increase in penetration. When equilibrium has again been
reach, the additional major load is removed but the preliminary minor load is still
maintained. Removal of the additional major load allows a partial recovery, so
reducing the depth of penetration. The permanent increase in depth of penetration,
resulting from the application and removal of the additional major load is used to
calculate the Rockwell hardness number as shown in Figure 3.2 (Ziheng Yao, 2005) .
Figure 3.2 Schematic diagram of Rockwell hardness test
3.1.3 Brinell hardness and Meyer hardness
The Brinell hardness test method consists of indenting the test material with a 10
mm diameter hardened steel or carbide ball subjected to a load of 3000 kg. For softer
materials the load can be reduced to 1500 kg or 500 kg to avoid excessive
indentation. The full load is normally applied for 10 to 15 seconds in the case of iron
and steel and for at least 30 seconds in the case of other metals. The diameter of the
indentation left in the test material is measured with a low powered microscope. The
Brinell harness number is calculated by dividing the load applied by the surface area
of the indentation as seen in Figure 3.3.

26
Figure 3.3 Schematic diagrams Brinell hardness test
The Meyer hardness is similar to the Brinell hardness except that the projected
area of contact rather than the actual curved surface area is used to determine the
hardness. In this case, the hardness number is equivalent to the mean contact pressure
between the indenter and the surface of the specimen. The mean contact pressure is a
quantity of considerable physical significance. The Meyer hardness is given by:
P P
H = Pm = =
(3.1)
2
A πa
A is the projected contact area and a is the real contact radius during indentation
(Ziheng Yao, 2005).
3.2 Determination of Young’s modulus by load-depth sensing indentation
For the past two decades, the advent of nano- and micro- scale science,
engineering and technology coupled with substantial progress in instrumentation has
resulted in ‘instrumented’ indentation or ‘load-depth sensing’ indentation. Figure 3.4
shows that typical load-depth sensing indentation test graph. It primarily consists of a
controlled load (P) applied through a diamond tip that is in contact with a specimen.
The penetration depth (h s ) of the tip into the material is recorded as a function of the
applied load. There is no question that the loading part is elastic-plastic response.
The unloading part is usually considered pure elastic rebound of the material. It is
only related to the elastic property of the material. If the area in contact is assumed to
remain constant during initial unloading, the elastic behavior may be modeled as that

27
of a blunt punch indenting an elastic solid (Loubet et al, 1984, Doerner and Nix,
1986 ). They obtained:
Figure 3.4 Characteristic region of load-depth diagram
S
dP A
= = 2Er
(3.2) and where
dh π
s
2
1 1−ν
1−ν
0
= + (3.3)
E E
r
E o
2
Er is called reduced modulus or combined modulus, S=dP/dhs is the
experimentally measured stiffness of the upper portion of the unloading data, which
is the slope of the curve fitted straight line of the initial part of unloading, A is the
projected contact area of the indenter at maximum loading condition, E and ν are
Young’s modulus and Poisson’s ratio for the specimen, and E o and νo are the same
parameter for the indenter.
The substitution of reduced modulus in Equation (3.2) for indentation test data is
valid (A.C. Fischer-Cripps, 2001). Because of the utilization of the slope or
unloading stiffness, it makes no difference whether or not the deflection of the
indenter is accommodated explicitly or transferred to that occurring within the

28
specimen by artificially reducing the specimen modulus from its true value to lower
value, the reduced modulus. Usually the indenter is assumed to be perfectly rigid and
E 0 =∞. So, the Equation (3.3) becomes:
E
E r
=
(3.4)
2
1−ν
and plug into the equation (3.2) we can get:
2
1−ν
π dP
E = (3.5)
2 A dh s
The equation (3.5) is applicable to any indenter that can be described as a body of
revolution of a smooth function (Pharr, Oliver and Brotzen, 1992). To evaluate
independently the projected contact area, it is proposed a simple empirical method
based on extrapolating the initial linear portion of the unloading curve to zero load
and using the extrapolated depth with the indenter shape function to determine the
contact area (A, A.K. Bhattacharya and W.D. Nix, 1988). One of the more
commonly used methods to get contact area by analyzing micro indentation loaddisplacement
data is that of Oliver and Pharr method, which expands on ideas
developed by Loubet et al. and A. K. Bhattacharya and Nix. They found the loaddisplacement
curves during unloading are not linear for most materials, even in the
initial stages. The analysis begins by fitting the unloading curve to the power-law
relation:
P h s
h f
)
m
= B( −
(3.6)
where P is the indentation load, hs is the displacement, B and m are empirically
determined fitting parameters, and hf is the final displacement after complete
unloading. By differentiating above equation at the maximum depth of penetration,
hs=hm, giving stiffness S:
dp
dh
m−1
S = ( hs
= hm
) = mB(
hm
− h
f
)
(3.7)
s
The depth along which contact is made between the indenter and the specimen, hc,
can also be estimated from the load-displacement data using:

29
h
c
Pm
= hm
− ε
(3.8)
S
where Pm is the peak indentation load and ε is a constant which depends on the
geometry of the indenter. With these basic measurements, the projected contact area,
A, is derived by evaluating an indenter shape function at the contact depth, hc, that is
A=f(hc). Finally, substitute S in Equation (3.7) and the projected area A into
Equation (3.5) to obtain the Young’s modulus of the specimen. It is important to note
these equations were derived from pure elastic contact solution derived by Sneddon,
and how well they work for elastic/plastic indentation is not entirely clear. One
important way in which the elastic solution fails to properly describe elastic/plastic
behavior concerns pileup and sink-in of material around the indenter. In the pure
elastic contact solution, material always sinks in, while for elastic/plastic contact,
material may either sink in or pile up. Since this has important effects on the
indentation contact data, it is not surprising that the Oliver-Pharr method has been
found to work well for hard ceramics, in which sink-in predominates, but significant
errors can be encountered when the method is applied to soft metals that exhibit
extensive pileup. It is discussed the influences of pileup on the measurement of
Young’s modulus and pointed out that when pileup is large, the areas deduced from
analysis of the load displacement curves underestimates the true contact areas by as
much as 60% (A.Bolshakov and G.M. Pharr, 1998). This, in turn, leads to
overestimate the hardness and elastic modulus. The parameter,
h f
h max
which can be
measured experimentally and correlated with the material parameters E, ν, σ y and n
dσ
( ) which control indentation deformation, can be used as an indication of
dε
whether or not pileup is an important factor. Pileup is significant only when
h f
h max
>0.7 and the material does not appreciably work harden. When,
h f
h max

30
Recently, it is shown that since the boundary conditions used to derive elastic
contact models employed in indentation allow for inward displacement of the
surface, a correction factor, β, needs to be added to Equation (3.2) which becomes:
dP A
S = = 2βEr
(3.9)
dh π
s
So the formula to find Young’s modulus (3.5) becomes (Hay et al, 1998):
2
(1 − 2ν
) π dP
E = (3.10)
2β
A dh s
3.3 Determination of Elastic modulus by Vickers indentation
Regarding mechanical properties, hardness testing provides useful information on
the strength and deformative characteristics of the materials (elastic modulus, elastic
recovery, hardness, etc.). Hardness is a mechanical parameter which is strongly
related to the structure and composition of solids. Hence, microhardness is not only a
mechanical characteristic routinely measured but it has also been developed as an
investigation method of structural parameters in recent years. Therefore, hardness
experiments have become more and more important to characterize a material
(Giannakopoulos, A.E. and Suresh, 1999 and Çimenoglu et al, 2003).
The characteristic ability of a material to resist penetration of an indenter
allows evaluation of a parameter that we know hardness. The indentation hardness of
materials is measured in several ways by forcing an indenter having specific
geometry (ball, cone, and pyramid) into the specimens’ surface.
The conventional microhardness value can be determined from the optical
measurement of the residual impression left behind upon load release. In recent
decades, the development of depth-sensing indentation equipment has allowed the
easy and reliable determination of two of the most commonly measured mechanical
properties of materials, the hardness and Young’s modulus. The depth-sensing (or

31
dynamic) micro indentation method offers great advantages over conventional
Vickers microhardness testing in two aspects. Firstly, apart from microhardness (or
micro strength), the method can also provide well-defined mechanical parameters
such as elastic modulus of the interfacial zone. Secondly, as load and depth of an
indentation are continuously monitored, optical observation and measurement of
diagonal length of the indent/impression, which can be difficult and subjected to
inaccuracy, is no longer required (Uzun et al, 2005).
Figure 3.5 Schematic plots of (a) cross-section of an indentation and (b) a typical
load–displacement curve showing the values used in the Oliver and Pharr method
(Uzun et al, 2005).
Two mechanical properties, namely, elastic modulus E and microhardness H can
be obtained with the load and penetration depth data. A typical load–penetration
depth curve is shown in Fig. 3.5. During indenter loading, test material is subjected
to both elastic and plastic deformation. The three key parameters needed to

32
determine the hardness and modulus are the peak load (P max ), the contact area (A c )
and the initial unloading contact stiffness (S). Similar to the conventional
microhardness testing, the micro indentation hardness is usually defined as the ratio
of the peak indentation load, Pmax, to the projected area of the hardness impression,
Ac, i.e (Uzun et al, 2005).
P P
H =
26
c
max max
= (
2
Ac
.43h
2
A
c
= 26.43hc
) (3.11)
Different approaches for deducing the contact depth, hc, from the resultant load
displacement curve have been purposed and perhaps the most widely used one is that
of Oliver and Pharr. The Oliver and Pharr data analysis procedure begins by fitting
unloading curve to an empirical power-law relation.
P = α(h − h f ) m (3.12)
Where P is the indentation load, h is the penetration depth, h f is the final unloading
depth and α and m are empirically determined fitting parameters. Using the initial
part of the unloading curve, both stiffness and contact depth are determined by
differentiating Eq. (3.12) at the maximum depth of penetration, h = h max . Then, the
stiffness of the contact is given by
S
dP 2
= = E r
A c
(3.13)
dh π
Where E r is the reduced elastic modulus.
3.4 Correlation between yield strength and hardness
The development of indentation methodologies for the micro mechanical
characterization of materials requires a precise understanding of the correlation
between uniaxial mechanical properties and hardness. One of such fundamental
correlations was found by Tabor for pyramidal (Vickers) indenters. Considering

33
indentation experiments conducted in specimens of pure copper and mild steel which
were previously subjected to different amounts of strain hardening, Tabor proposed
that hardness is, to a great extent, proportional to the uniaxial stress at a plastic strain
of 0.08. Namely,
H = Cσ
(3.14)
r
Where H is the Vickers hardness of the material, C=3.3, and σ
r
is the uniaxial stress
corresponding to a characteristic uniaxial strain (ε r ) of 0.08 (Ziheng Yao, 2005).
3.5 Vickers Indentation Cracks and Fracture Toughness
Fracture toughness determination with Vickers Hardness indentation was
proposed by Evan and Charles and later extended and modified by Niihara et al,
Antis et al and Lawn et al. With this method the fracture toughness is calculated from
the length of cracks which develop during a Vickers indentation test and can be
measured optically at the specimen surface. In a modified procedure a strength test is
performed with the damaged specimen and K IC is determined from the indentation
load and bending strength.
In Figure 3.6 the development of the Vickers indentation cracks is illustrated
according to Binner and Stevens:
Below the Vickers pyramid a deformation zone develops (a). During loading and
unloading two perpendicular cracks are initiated starting at the deepest location on
deformation zone (b) and propagate to the specimen surface (c). The final crack is
nearly semicircular (d). The crack length at the surface (e) 2c, the length of the
indentation diagonal 2a.
1 / 2 − 3 / 2
⎡ E ⎤ ⎡ c ⎤
K IC
∞ H a
⎢
⎣ H ⎥
⎦
⎢
⎣ a ⎥
(3.15)
⎦

34
From theoretical consideration it follows that for the fracture toughness with
the hardness H calculated from the indentation load F and the diagonal a of the
impression.
F
H = (3.16)
2
2a
The exponent for E/H was given in earlier investigation as 0.4. Different
values are proposed in the literature for the prefactor in (2.18). The best agreement
with experimental data was found by Anstis et al. to be:
1/ 2
−3 / 2
⎛ E ⎞ ⎛ c ⎞
K IC
= 0.032H
a⎜
⎟ ⎜ ⎟
(3.17)
⎝ H ⎠ ⎝ a ⎠
Figure 3.6 Development of a Vickers indentation crack under an increasing load P and after load
removal, according to D. Munz and T. Fett

35
CHAPTER FOUR
EXPERIMENTAL STUDIES
4.1 Purpose
Serviceable engineering components not only rely on their bulk material
properties but also on the design and characteristics of their surface. This is
especially true in high hardness components, as their surfaces must perform many
industrial functions in a variety of complex environments. The surface of industrial
component may require treatment to enhance the surface characteristic. Surface
treatments that cause microstructure changes in the bulk material include heating and
cooling/quenching through induction, flame, laser and electron beam techniques, or
mechanical treatments.
One of the thermo chemical surface treatments of steel based material is boriding
process. In our study, our samples were borided by pack boriding process, as
mentioned Chapter Two, for 2, 4 and 6 h at 900 o C.
The objective of this study was to obtain micro structural characteristics and
investigate the mechanical properties such as hardness, young’s modulus and fracture
toughness of the layer. The produced double layer (FeB and Fe 2 B) new surface was
extensively analyzed with respect to X-ray diffraction (XRD), scanning electron
microscopy (SEM) including energy dispersive spectroscopy (EDS), microhardness
and surface roughness. Mechanical properties of layers was examined by Shimadzu
Dynamic Ultra-micro Hardness test machine for estimating young’s modulus due to
load-unload sensing analysis, in addition to mechanical investigation hardness-depth
and hardness-force curves of the layer was obtained. Fracture toughness properties of
surface layer FeB was calculated by Vickers Fracture Toughness method with
measuring crack length after loading stage is finished.
35

36
4.2 Materials
SAE 1020 and SAE 1040 quality steel was used as a substrate material for
boriding process for 2, 4 and 6 h at 900 o C. According to the spectrometer results,
chemical compositions of substrate materials are listed Table 3.1.
Table 3.1 Chemical composition of substrate materials
Steel Quality % C % Si % Mn % Cu % P % S
SAE 1020, DIN 17100, EN 10056 0.20 0.20 0.85 0.15 0.04 0.035
SAE 1040, DIN 17100, EN 10056 0.40 0.20 1.10 0.20 0.04 0.035
4.3 Characterization Studies
4.3.1 X-Ray Diffractions (XRD)
X-ray diffraction (XRD) patterns of thin films were determined by means of multi
purpose Rigaku D/Max-2200/PC Model diffractometer with a Cu K α radiation by
using multi purpose thin film attachment. Measurements were performed by applying
40 kV voltages and 36 mA current.
4.3.2 Scanning Electron Microscopy / Energy Dispersive Spectroscopy (EDS)
The surface morphologies of FeB and Fe 2 B layers were examined by a Scanning
Electron Microscope (JEOL-JSM 6060 SEM) with an Energy Dispersive X-ray
spectroscopy (IXRF System EDS) system attachment. Accelerating voltage of 20 kV
was used for the SEM imaging and SEM/EDS analyses.

37
4.4 Surface Roughness Tester
SJ-301 Mitutoyo surface roughness tester was used for determination of layers
roughness. Test measurement results give; Ra values of surface which means the
differences between maximum and minimum asperities height versus measurement
distance.
4.5 Dynamic Ultra-micro Hardness Tester DUH-W201/W201S
Mechanical and elastic properties (hardness, young modulus, yield stress, etc.) of
FeB layers were obtained from load-indentation curves by using ultra-micro hardness
tester.
4.6 Finite Element Modeling
Finite element simulations are performed for the problem of Vickers indenter
brought into contact against an infinite half space. Both the indenter and half space
are modeled as axial symmetry geometry. While the isotropic hardening has been
regarded as an accurate model in finite element modeling of indentation, this
assumption is merely a simplification of the true mechanical response of the material
as the shape of the yield locus in most metals varies during plastic flow. M. Mata et
al. [20] stated that ‘the hypothesis of isotropic hardening is considered to provide
accurate results in the analysis of loading histories where no reversed plastic
deformation occurs well beyond the yield surface. The response is taken to be rate
independent, so in the quasi-static calculations presented here no rate effects are
represented. The influence of large deformation is included in the analysis by using
geometry nonlinearity activated in ABAQUS. The Von Misses yield criterion is
taken for the computation.

38
CHAPTER FIVE
RESULTS AND DISCUSSION
5.1 Characterization of the Layers
5.1.1 XRD
The XRD pattern of the borided layers, which was formed on SAE 1020 and 1040
quality steel at 900 o C for 2, 4 and 6h, are given in figure 5.1 and figure 5.2,
respectively. According to this pattern, layer includes only FeB phase with (111),
(210), (101) and (111).
140
120
100
SAE 1020 Quality Steel Borided at 900 C
2h borided
4h borided
6h borided
Intensity (%)
80
60
40
20
0
39 44 49 54 59 64 69
2 Theta
Figure 5.1 Phase analysis of 1020 steel borided at 900 o C for 2h, 4h, and 6h
38

39
140
120
SAE 1040 Quality Steel Borided at 900 C
2h borided
4h borided
6h borided
100
Intensity (%)
80
60
40
20
0
39 44 49 54 59 64 69
2 Theta
Figure 5.2 Phase analysis of 1040 steel borided at 900 o C for 2h, 4h and 6h.
Both XRD examinations of borided layers show that only FeB phase exists on the
surface. But if used SEM, cross-sectional investigation of layers demonstrate that
two diffusion layers (FeB and Fe 2 B) in saw-tooth shapes.
5.1.2 SEM and Thickness examination
SEM cross-sectional investigations show that double phase layer from surface to
inside of substrate. SEM cross-sectional photograph of SAE 1020 and SAE 1040
quality steels, which was borided at 900 o C for 2, 4 and 6h, are shown in Figure 5.3.
and Figure 5.4., respectively. The structural compositions of layers consist of boron
rich phase (FeB) and iron rich phase (Fe 2 B), respectively. FeB formation begins from
surface, Fe 2 B phase forms in deeper region because of decreasing boron
concentration form surface to inside of substrate. According to the SEM photograph,
saw tooth microstructural images show both FeB and Fe 2 B phase.

40
a)
b)

41
c)
Figure 5.3 Borided layer thickness of 1020 steel for different process time SAE 1020 steel
a) 2 h borided, b) 4 h borided, c) 6 h borided
During boriding, the diffusion and subsequent absorption of boron atoms into
metallic lattice of component surface form interstitial boron compounds. The
resulting layer may consist of either a single phase boride or polyphase boride layer.
Saw-tooth shape diffusion microstructures of borided SAE 1020 and SAE 1040 are
illustrated in Figure 5.3. and Figure 5.4., respectively. From SEM-BEC photographs
can be seen that films consist of two different diffusion layers. According to the
micro structure graphs; dark region symbolizes FeB diffusion layer while bright
areas indicate Fe 2 B diffusion layer. Increase in time from 2h to 6 h, FeB and Fe 2 B
layers thicknesses increase; as expected.

42
a)
b)

43
Figure 5.4 Borided layer thickness of 1040 steel for different process time SAE 1020 steel
a) 2 h borided, b) 4 h borided, c) 6 h borided
c)
The formation of a single Fe 2 B phase (with saw tooth morphology due to
preferred diffusion direction) is more desirable than a double-phase layer (FeB and
Fe 2 B). Because of FeB is more brittle than the iron sub boride Fe 2 B layer, the boron
rich FeB phase is considered undesirable. Also, because FeB and Fe 2 B formed under
tensile and compressive residual stress, respectively, crack formation is often
observed at or in the neighborhood of the FeB / Fe 2 B interface of double-phase layer
as seen in Figure 5.3 and Figure 5.4. for borided SAE 1020 and SAE 1040 quality
steels, respectively. These cracks may lead to flaking and spalling when a
mechanical strain is applied or even separation when a component is undergoing a
thermal and/or mechanical shock.

44
Average thickness measurement of FeB and Fe 2 B layer on SAE 1020 and SAE
1040 listed in Table 5.1. Because boriding process is controlled by diffusion
parameter such as time and temperature, formed layer thicknesses change with these
parameters.
Table 5.1 Average thickness measurement of different hour borided steel
Material
Type
SAE 1020
SAE 1040
Boriding
Time
(hour)
FeB Layer
Thickness (µm)
Fe 2 B Layer
Thickness (µm)
Total Thickness
(µm)
2 23.6 27.3 50.9
4 27.1 70.1 97.2
6 42.5 90.8 133.3
2 29.8 56.8 86.6
4 37.4 63.6 101
6 49.8 69.7 119.5
SAE 1020 QUALTIY STEEL
Layer Thickness (micrometer)
100
90
80
70
60
50
40
30
20
FeB Layer Thickness
Fe2B Layer Thickness
27,3
23,6
70,1
27,1
90,8
42,5
10
0
0 1 2 3 4 5 6 7
Boriding Time (h)
a)

45
SAE 1040 QUALITY STEEL
80
70
FeB Layer Thickness
Fe2B Layer Thickness
69,7
Layer Thickness (micrometer)
60
50
40
30
20
56,8
29,8
63,6
37,4
49,8
10
0
0 1 2 3 4 5 6 7
Boriding Time (h)
b)
Figure 5.5 Thickness measurement graphs of FeB and Fe 2 B Layer on; a) SAE 1020 quality
steel and b) SAE 1040 quality steel for 2h, 4h and 6h process time.
Figures 5.5 a) and 5.5 b) show the thickness value of FeB and Fe 2 B layers on SAE
1020 and 1040 quality steel with different process time such as 2h, 4h and 6h.
According to the SEM-BEC mode cross-sectional investigation in Figure 5.3 and
figure 5.4, FeB and Fe 2 B layers occurs, respectively, towards into the substrates.
Together chemical composition of substrates with diffusion parameters influence the
layer thicknesses and characteristics (FeB and Fe 2 B formation).

46
5.2 Surface Roughness Measurement Samples
Surface roughness value of samples is important parameter for nano-indentation
test with Dynamic Ultra Micro Hardness Tester. After diffusion controlled boriding
process, material surface roughness can be high for micro-indentation test. So,
surface polishing process is applied for decreasing roughness of FeB layers. Because
of micro indentation test result’s sensitivity and homogeneity, polishing process must
be applied whole surface to minimize Ra value of samples.
Table 5.2 Surface roughness measurements of samples
Samples SAE 1020/Fe 2 B/FeB SAE 1040/Fe 2 B/FeB
Process Time (hour) 2 4 6 2 4 6
Before Polishing Ra
(micrometer)
0.90 0.95 1 0.89 0.95 1.2
After Polishing Ra
(micrometer)
0.09 0.08 0.010 0.05 0.07 0.09
5.3 Ultra Micro Hardness Examination of Samples
The mechanical properties such as hardness, Young’s modulus, fracture
toughness, ductility, etc are important parameters for industrial application.
Shimadzu Dynamic Ultra Micro Hardness Testing machine is used for determination
hardness variation and young modulus of FeB layers. The three key parameters such
as peak load (P max ), the contact area (A c ) and the initial unloading contact stiffness
(S) can be determined by load unload test mode for different loading value as
explained in chapter 3. Different load range such as 640 mN, 320 mN, 160 mN and
80 mN is applied for determination of hardness and Young’s modulus.

47
Hardness results of 2h, 4h, and 6h borided SAE 1020 quality steel at 640 mN, 320
mN, 160 mN and 80 mN peak loads, are shown in Figure 5.6.
45000
40000
35000
1020-2-640
1020-4-640
1020-6-640
Hardness (HDV)
30000
25000
20000
15000
10000
5000
0
0 0,2 0,4 0,6 0,8 1 1,2 1,4 1,6 1,8
Depth "h" (micrometer)
a)
20000
18000
16000
1020-2-320
1020-4-320
1020-6-320
14000
Hardness (HDV)
12000
10000
8000
6000
4000
2000
0
0 0,2 0,4 0,6 0,8 1 1,2
Depth "h" (micrometer)
b)

48
30000
25000
1020-2-160
1020-4-160
1020-6-160
Hardness (HDV)
20000
15000
10000
5000
0
0 0,1 0,2 0,3 0,4 0,5 0,6 0,7 0,8 0,9
Depth "h" (micrometer)
c)
35000
30000
1020-2-80
1020-4-80
1020-6-80
25000
Hardness (HDV)
20000
15000
10000
5000
0
0 0,1 0,2 0,3 0,4 0,5
Depth "h" (micrometer)
d)
Figure 5.6 Hardness - penetration depth curve of 2h, 4h and 6h borided SAE 1020 quality
steel under a) 640 mN, b) 320 mN, c) 160 mN and d) 80 mN, peak loads.

49
The figures were constructed experimentally using the data taken from the loading
part of depth setting DHV measurements at 640 mN, 320 mN, 160 mN and 80 mN
applied peak load for each sample in Figure 5.6 and Figure 5.7 for borided SAE 1020
and SAE 1040 quality steel, respectively. Because diffusion controlled process used
to form FeB layer, structure and composition can be changed from surface into
samples. As follows, surface of samples consist of too much boron concentration in
respect to inside of samples. So, composition and structure change depending on
diffusion distance from surface to inside of specimen as explained chapter 3. In
addition to hardness behaviour of FeB layer as shown figure 5.6 and figure 5.7, it is
seen that HV numbers decrease with increasing applied peak load from surface to
inside of layer. Since hardness is accepted as an inherent material property, it should
not vary with indentation load and size. However, investigations have confirmed that
HV number of materials were indentation size dependent especially at lower peak
loads. Increase in hardness with decreasing applied peak load cause from differences
in indentation depth, therefore this effect is called indentation size effect. Figure 5.6
and figure 5.7 exhibit this kind of behaviour.
18000
16000
14000
1040-2-640
1040-4-640
1040-6-640
Hardness (HDV)
12000
10000
8000
6000
4000
2000
0
0 0,2 0,4 0,6 0,8 1 1,2 1,4 1,6 1,8 2
Depth "h" (micrometer)
a)

50
45000
40000
35000
1040-2-320
1040-4-320
1040-6-320
Hardness (HDV)
30000
25000
20000
15000
10000
5000
0
0 0,2 0,4 0,6 0,8 1 1,2
Depth "h" (micrometer)
b)
50000
45000
40000
1040-2-160
1040-4-160
1040-6-160
35000
Hardness (HDV)
30000
25000
20000
15000
10000
5000
0
0 0,1 0,2 0,3 0,4 0,5 0,6 0,7 0,8
Depth "h" (micrometer)
c)

51
250000
225000
200000
1040-2-80
1040-4-80
1040-6-80
Hardness (HDV)
175000
150000
125000
100000
75000
50000
25000
0
0 0,1 0,2 0,3 0,4 0,5
Depth "h" (micrometer)
d)
Figure 5.7 Hardness - penetration depth curves of 2h, 4h and 6h borided SAE 1040 quality
steel with a) 640 mN, b) 320 mN, c) 160 mN and d) 80 mN peak loads.
With the ever increasing research and application interests in surface coatings,
micro or nano indentation experiments have been extensively used to determine the
mechanical properties of materials. Hardness and Young’s modulus are the two
frequently studied ones. FeB layers hardness examinations were made by different
applied peak loads.
The load-unload mode (load–displacement) curves shown in Figs. 5.8–5.9 (SAE
1020 and 1040 quality steel with 2h, 4h and 6h boriding process time, respectively)
represent the 640mN, 320 mN, 160mN and 80mN applied load as a function of the
displacement (elastic and plastic) of the indenter with respect to the initial position of
the surface and Table 5.3 shows the hardness, Young’s modulus (according to
equation 3.3 and 3.13), residual depth and maximum depth value of samples under
640mN, 320 mN, 160mN and 80mN applied peak loads.

52
Force (mN)
700
650
600
550
500
450
400
350
300
250
200
150
1020-2-640
100
1020-4-640
50
1020-6-640
0
0 0,2 0,4 0,6 0,8 1 1,2 1,4 1,6 1,8 2 2,2 2,4
Depth "h" (micrometer)
a)
350
300
250
Force (mN)
200
150
100
1020-2-320
50
1020-4-320
1020-6-320
0
0 0,1 0,2 0,3 0,4 0,5 0,6 0,7 0,8 0,9 1 1,1 1,2
Depth "h" (micrometer)
b)

53
180
160
140
120
Force (mN)
100
80
60
40
1020-2 160
20
1020-4 160
1020-6 160
0
0 0,1 0,2 0,3 0,4 0,5 0,6 0,7 0,8 0,9
Depth "h" (micrometer)
c)
90
80
70
Force (mN)
60
50
40
30
20
10
0
1020-2-80
1020-4-80
1020-6-80
0 0,05 0,1 0,15 0,2 0,25 0,3 0,35 0,4 0,45 0,5 0,55
Depth "h" (micrometer)
A schematic representation of indentation load (P) versus displacement (h) data
obtained during one full cycle of loading and unloading is presented in Figure 5.8-
5.9. The important quantities are the peak load (P max ), the maximum depth (h max ), the
d)
Figure 5.8 A series of load-displacement plots at different peak loads of SAE 1020/FeB
layers with 2h, 4h and 6h process time, a) 640mN, b) 320 mN, c) 160 mN, d) 80 mN

54
final or residual depth after unloading (hr), contact depth (h c ) and the slope of the
upper portion of the unloading curve (S=dP/dh). The parameter S has the dimensions
of force per unit distance and is known as the elastic contact stiffness, or more
simply, the contact stiffness.
In order to determine the elastic modulus, maximum applied load should be
sufficient to produce a permanent deformation on the coating surface. So that the
values of elastic modulus, determined from indentations, do not depend on the value
of h and, therefore, on the value of the maximum load, the indentation depth should
not exceed 10-25 % of the coating thickness, otherwise the results will be affected by
the properties of the substrate. In this study, diffusion controlled FeB layers thickness
measurement results show that maximum thickness is 42.5 µm and 49.8 µm for SAE
1020 and SAE 1040 quality steel, respectively and maximum indentation depth are
2.126 and 1.646 for SAE 1020 and SAE 1040 quality steel, respectively as shown in
Table 5.3.
Force (mN)
700
650
600
550
500
450
400
350
300
250
200
150
1040-2-640
100
1040-4-640
50
1040-6-640
0
0 0,2 0,4 0,6 0,8 1 1,2 1,4 1,6 1,8
h (micrometer)
a)

55
350
300
250
Force (mN)
200
150
100
1040-2-320
50
1040-4-320
1040-6-320
0
0 0,1 0,2 0,3 0,4 0,5 0,6 0,7 0,8 0,9 1 1,1 1,2
Depth "h" (micrometer)
b)
180
160
140
120
Force (mN)
100
80
60
40
1040-2 160
20
1040-4 160
1040-6 160
0
0 0,1 0,2 0,3 0,4 0,5 0,6 0,7 0,8 0,9
Depth "h" (micrometer)
c)

56
90
80
70
60
Force (mN)
50
40
30
20
10
1040-2-80
1040-4-80
1040-6-80
0
0 0,05 0,1 0,15 0,2 0,25 0,3 0,35 0,4
Depth (h) micrometer
Figure 5.9 A series of load-displacement plots at different peak loads of SAE 1040/FeB
layers with 2h, 4h and 6h process time, a) 640mN, b) 320 mN, c) 160 mN, d) 80 mN
d)
Eq. (3.13) as mentioned in chapter 3 is the basic equation for determination of
reduced elastic modulus by micro indentation. The key quantities are the initial
unloading contact stiffness, S = dP/dh and the real contact depth, h c in order to
determine E r .
S
dP 2 *
= = E A c
(3.13 in chapter 3)
dh π
2
1 1−ν
1−ν
0
= + (3.3 in chapter 3)
E E
r
E o
2
E r
is called reduced modulus or combined modulus, S=dP/dh s is the
experimentally measured stiffness of the upper portion of the unloading data, which
is the slope of the curve fitted straight line of the initial part of unloading, A is the
projected contact area of the indenter at maximum loading condition, E and ν are
Young’s modulus and Poisson’s ratio for the specimen, and E o and ν o are the same
parameter for the indenter. Using the experimentally determined S and h c , the

57
reduced elastic modulus by micro indentation was calculated and the results are
shown in Fig. 5.10 a) and Fig. 5.10 b) for borided SAE 1020/FeB and SAE 1040/FeB
quality steels, respectively. It is clearly seen from the figures that the extracted
reduced elastic modulus also exhibits a strong peak load dependency as shown in
Table 5.3. According to the result, Young’s modulus values increase with decreasing
applied peak loads in peak load.
Table 5.3 D.U.H analysis of borided SAE 1020 and SAE 1040 quality steels.
Material Type
1020–2h
1020–4h
1020–6h
1040–2h
1040–4h
1040–6h
Load
(mN)
Maximum
Depth (µm)
Residual
Depth (µm)
Hardness
(DHV)
Young’s
Modulus
(GPa)
640 2.126 1.383 775 125
320 1.102 0.641 1001 240
160 0.781 0.499 1033 276
80 0.280 0.038 1621 432
640 1.637 1.099 892 271
320 0.899 0.544 1060 342
160 0.560 0.301 1310 352
80 0.453 0.305 1552 566
640 1.577 1.088 1081 241
320 0.869 0.554 1277 396
160 0.795 0.506 1381 397
80 0.501 0.381 1788 568
640 1.646 1.040 1015 273
320 1.094 0.669 1045 285
160 0.751 0.323 1367 345
80 0.369 0.127 1961 496
640 1.530 1.061 1102 366
320 0.908 0.527 1190 391
160 0.562 0.237 1736 497
80 0.269 0.065 2280 596
640 1.677 1.125 1120 285
320 0.878 0.504 1210 400
160 0.709 0.423 1486 551
80 0.211 0.050 2244 624

58
600
500
1020-2
1020-4
1020-6
400
E (GPa)
300
200
100
0
0 100 200 300 400 500 600 700
Load (mN)
a)
700
600
1040-2
1040-4
1040-6
500
E (GPa)
400
300
200
100
0
0 100 200 300 400 500 600 700
Load (mN)
b)
Figure 5.10 The elastic modulus extracted from the analysis of the load-displacement curves as a
function of the peak loads for; a) SAE 1020/FeB, b) SAE 1040/FeB

59
5.4 Fracture Toughness Measurement
The fracture toughness of a material is of critical importance for use in
mechanical applications. The use of the Vickers indentation method to asses fracture
toughness of brittle materials, particularly glasses and ceramics has been well
developed. This technique can be used on small samples. The specimen preparation
is relatively simple requiring only a flat and polished surface (Ugur Sen et al, 2005).
The Vickers diamond indenter is a standard item used on a dedicated hardness tester
or on a universal testing machine. In many instances the crack lengths can be
measured optically without difficulty. It is also quick and cost-effective (A.H. Üçisik,
C. Bindal, 1997).
The materials used for this study were low-alloy; SAE 1020 and SAE 1040
quality steels. The test pieces prepared a square shape with nominal dimensions of 15
mm x 15 mm x 3 mm. The chemical compositions of base steels used in this study
are given in Table 3.1 at chapter 3. Indentation fracture toughness tests were
performed on the polished cross-sections of the borided layers formed on the nearsurface
regions of low-alloy steels. Fracture toughness was measured with a Vickers
indenter under a load of 10 N. Crack lengths were immediately measured under
SEM.

60
a)
b)

61
Figure 5.11 SEM images of Vickers indent in FeB borided layer of SAE 1020 quality steel
with 2h process time; a) Vickers indentation marks, b) left crack length of Vickers indent,
c) right crack length of Vickers indent
c)
For 1020–2 ;
F 98,07
12
H = =
= 0,0138x10
N
2 2
2a
2x
m
−6
( 59,49x10
)
2
K
IC
⎛ E
= 0,016x⎜
⎝ H
⎟
⎠
⎞
1
2
x
F
3
( C ) 2
b
3
⎛ 590x10
⎜
= 0,016x
10
⎜
⎜ 0,0138x10
⎝
−6
12
⎞
⎟
⎟
⎟
⎠
1
2
x
98,07
3
−6
( 107,63x10
) 2
= 0,9188MPa
m

62
a)
b)

63
c)
Figure 5.12 SEM images of Vickers indent in FeB borided layer of SAE 1020 quality steel
with 4h process time; a) Vickers indentation marks, b) left crack length of Vickers indent, c)
right crack length of Vickers indent
For 1020–4;
F 98,07
12
H = =
= 0,0103x10
N
2 2
2a
2x
m
−6
( 68,855x10
)
2
K
IC
⎛ E
= 0,016x⎜
⎝ H
⎟
⎠
⎞
1
2
x
F
3
( C ) 2
b
3
⎛ 590x10
⎜
= 0,016x
10
⎜
⎜ 0,0103x10
⎝
−6
12
⎞
⎟
⎟
⎟
⎠
1
2
x
98,07
3
−6
( 102,6x10
) 2
= 1,1427MPa
m

64
a)
c)

65
c)
Figure 5.13 SEM images of Vickers indent in FeB borided layer of SAE 1020 quality steel
with 6h process time; a) Vickers indentation marks, b) bottom crack length of Vickers
indent, c) upper crack length of Vickers indent
For 1020–6;
F 98,07
12
H = =
= 0,0139x10
N
2 2
2a
2x
m
−6
( 59,205x10
)
2
K
IC
⎛ E
= 0,016x⎜
⎝ H
⎟
⎠
⎞
1
2
x
F
3
( C ) 2
b
3
⎛ 590x10
⎜
= 0,016x
10
⎜
⎜ 0,0139x10
⎝
−6
12
⎞
⎟
⎟
⎟
⎠
1
2
x
98,07
3
−6
( 93,73x10
) 2
= 1,1265MPa
m

66
a)
b)
Figure 5.14 SEM images of Vickers indent in FeB borided layer of SAE 1040 quality steel
with 2h process time; b) bottom crack length of Vickers indent, b) upper crack length of
Vickers indent

67
For 1040–2;
F 98,07
12
H = =
= 0,037x10
N
2 2
2a
2x
m
−6
( 115x10
)
2
K
IC
⎛ E
= 0,016x⎜
⎝ H
⎟
⎠
⎞
1
2
x
F
3
( C ) 2
b
3
⎛ 590x10
⎜
= 0,016x
10
⎜
⎜ 0,037x10
⎝
−6
12
⎞
⎟
⎟
⎟
⎠
1
2
x
98,07
3
−6
( 134,55x10
) 2
= 0,4014MPa
m

68
a)
b)
Figure 5.15 SEM images of Vickers indent in FeB borided layer of SAE 1040 quality steel
with 4h process time; a) upper crack length of Vickers indent, b) bottom crack length of
Vickers indent

69
For 1040-4;
F 98,07
12
H = =
= 0,0074x10
N
2 2
2a
2x
m
−6
( 81,1 x10
)
2
K
IC
⎛ E
= 0,016x⎜
⎝ H
⎟
⎠
⎞
1
2
x
F
3
( C ) 2
b
3
⎛ 590x10
⎜
= 0,016x
10
⎜
⎜ 0,0074x10
⎝
−6
12
⎞
⎟
⎟
⎟
⎠
1
2
x
98,07
3
−6
( 88,29x10
) 2
= 1,6888MPa
m

70
a)
b)
Figure 5.16 SEM images of Vickers indent in FeB borided layer of SAE 1040 quality
steel with 6h process time; a) upper crack length of Vickers indent, b) bottom crack
length of Vickers indent

71
For 1040-6;
F 98,07
12
H = =
= 0,0122x10
N
2 2
2a
2x
m
−6
( 63,38x10
)
2
K
IC
⎛ E
= 0,016x⎜
⎝ H
⎟
⎠
⎞
1
2
x
F
3
( C ) 2
b
3
⎛ 590x10
⎜
= 0,016x
10
⎜
⎜ 0,0122x10
⎝
−6
12
⎞
⎟
⎟
⎟
⎠
1
2
x
98,07
3
−6
( 83,67x10
) 2
= 1,4257MPa
m

72
5.5 FEM Model
In order to improve the calculation accuracy in the continuous FEM simulation of
the nanoindentation, an axisymmetric FEM model of the semi-infinite layered half
space was built (Swanson Analysis system, 1995). To fulfill this target it was
necessary to replace the Vickers pyramid through an equivalent cone. This
replacement increases the calculation accuracy, since it enables the description of a
three-dimensional problem through the application of a plane axisymmetric model.
The lack of edge regions of the pyramid indentor negligibly affects the penetration
procedure, because these regions are limited in comparison to the whole contact
indentor-specimen area. The applied Vickers pyramid and the corresponding defined
equivalent cone are demonstrated in the upper part of Fig. 5.17 (K.-D. Bouzakis,
2001).
Figure 5.17 Determination of an equivalent cone to the
Vickers pyramid nanoindentor, used in the developed FEM
simulation of the nanoindentation.

73
The criterion that governs this replacement is that the cross-section areas A and B
of the pyramid and conical indentor, respectively, at the same penetration depth h,
are equal. The equivalent cone data are defined equalizing the rectangle area of
section A, to a circular area of section B. Thus, the equivalent cone cross-section
radius r egv at the penetration depth h, is calculated by means of the equation:
a
r egv
=
π
Where a is the Vickers pyramid rectangle side length. At the bottom part of the
figure, the equivalent cone, penetrating the coated specimen is illustrated. In the case
of a Berkovich indentor, a triangular cross-section shape is considered. Taking into
account the aforementioned assumptions, a deformable diamond equivalent cone was
used to establish the FEM model, simulating the nanoindentation procedure.
In order to achieve a flexible and reproducible model, the indentor, the coating
and the substrate material properties as well as the penetration depth are variable and
changeable parameters. The simulation of the nanoindentation test has been
performed considering two load steps. The first load step, the so-called loading stage,
represents the indentation phase into the coating. During the second load step, the socalled
relaxation stage, the indentor cone is removed, leading to a material elasticplastic
recovery (K.-D. Bouzakis, 2001).
FEM was performed with the commercial software package ABAQUS 6.5-1. The
model was constructed with axial symmetry geometry as illustrated Figure 5.15. The
indenter had a conical tip with semi-vertical angle of 70.3, which gives the same
area-to-depth function as Berkovich and Vickers indenters. At the very tip of the
indenter, a spherical rounding with a radius of 0.5 mm was constructed because of
the fact that no real indenter can be ideally sharp. The indenter had a cylindrical body
which was large enough to uniformly transfer the load from the top surface to the
contact area. The material of the indenter was taken as diamond and assumed to be
elastic with Young’s modulus of E s = 1140 GPa and Poisson’s ratio=0.04 (C.A.
Brookes, 1979). Figure 5.18 shows the schematic of the finite element model used in
this work.

74
The indenter was meshed by approximately 1500 four-node and 18500 four-node
and eight nodes for specimen. Elements were finest and in the central contact area
and became coarser outwards and CAX4R and Quad-dominated element types were
used. The interaction between the diamond indenter and specimen was modeled by
without contact element with no friction.
Figure 5.18 Schematic of the finite element model used in this work
In this study, the FeB layers were modeled as elastic, perfect-plastic materials
(Young’s modulus E= 595 GPa, Poisson’s ratio, ν= 0.2). Yield strength was assumed
to be between 5 and 7 GPa to determine the same penetration depth with
experimental results. The substrate was chosen to be commercial steel and was
modeled as elastic material with E= 205 GPa and ν=0.3 and yield strength= 280
MPa, ultimate strength= 600 and work hardening exponent= 0.2 (J.L.He, S. Veprek,
2003).
Figure 5.19 and figure 5.20 show mesh design of the entire model and magnified
view of mesh design under Vickers indenter (with 70.3 o equivalent angle). Numerical
analysis steps include; loading, holding and unloading parts. Figure 5.21 a) and b)
and Figure 5.22 show loading and unloading step modules with Von Misses stress
distribution at contact region of indenter and layer, respectively.

75
Figure 5.19 Mesh design of in the entire model
Figure 5.20 Magnified view of mesh design under Vickers indenter

76
a)
b)
Figure 5.21 Magnified view of model after loading step under Vickers indenter
Figure 5.23 a) and b) shows numerical solution and experimental test results.
Yield strength was assumed to be between 5 and 7 GPa to determine the same

77
penetration depth with experimental results. With the yield strength of FeB layer is 5
GPa, Abaqus solution shows the same penetration depth with experimental result
under 640 mN applied peak load as seen in figure 5.23 b).
Figure 5.22 Magnified view of model after unloading step under Vickers indenter

78
720
640
560
Yield Strength=7000 MPa
Yield Strength=6500 MPa
Yield Strength=6000 MPa
Yield Strength=5000 MPa
Force (mN)
480
400
320
240
160
80
0
0 0,2 0,4 0,6 0,8 1 1,2 1,4 1,6 1,8 2 2,2 2,4
Depth "h" (micrometer)
a)
720
640
Yield Strength=5000 MPa-Abaqus Analysis
DUH analysis
560
480
Force (mN)
400
320
240
160
80
0
0 0,2 0,4 0,6 0,8 1 1,2 1,4 1,6 1,8 2 2,2 2,4
Depth "h" (micrometer)
b)
Figure 5.23 Comparison of numeric and experimental results with 640mN applied
peak load; a) FeB layer’s yield strength between 5 GPa to 7 GPa, b) Details of
numeric and experimental results under 640 mN applied peak load.

79
The load displacement is nearly a straight line for all numerical results. The
loading curves show a larger deviation than unloading curves. This may be due to
differences in the yield strength or due to use of constitutive model of materials like
perfect elastic-plastic one without work hardening rate (M. Toparli, N.S. Koksal,
2005).

80
CHAPTER SIX
CONCLUSION
6.1 General Results
In this study, characterization of FeB and Fe 2 B boride layer were investigated and,
mechanical properties such as young’s modulus, hardness and yield strength of upper
layer (FeB) were determined. The following conclusions are obtained from the
experimental works:
1- XRD peaks demonstrated that the boron rich phase (FeB) with (111), (210),
and (101) planes was formed on surface of both SAE 1020 and SAE 1040
quality steels.
2- SEM cross-sectional investigations of borided steels show that double phase
layer forms from surface to inside of substrate. The structural compositions of
layers consist of boron rich phase (FeB) and iron rich phase (Fe 2 B),
respectively. From SEM-BEC photographs can be seen that films consist of
two different diffusion layers. According to the saw-tooth like micro structure
graphs; dark region symbolizes FeB diffusion layer while bright areas
indicate Fe 2 B diffusion layer. As FeB and Fe 2 B layers formed, tensile and
compressive residual stress come out, respectively, crack formation is often
observed at or in the neighborhood of the FeB / Fe 2 B interface of doublephase
layer.
3- Average thickness measurements of FeB and Fe 2 B layer on SAE 1020 and
SAE 1040 were determined from SEM graphs. Increase in time from 2h to 6
h, FeB and Fe 2 B layers thicknesses increase from 23.6 µm to 42.5 µm and
from 27.3 µm to 90.8 µm, respectively; as expected, for SAE 1020 quality
substrate. Like borided SAE 1020 quality steel substrate, FeB and Fe 2 B layers
thicknesses increase for same process time from 29.8 µm to 49.8 µm and
from 56.8 µm to 69.7 µm, respectively for SAE 1040 quality steel substrate.
4- Surface roughness of upper layer (FeB) was measured before and after
polishing process. Because micro indentation test results are very sensitive to
80

81
surface roughness, polishing process was applied to whole surface to
minimize Ra value of samples. After polishing process, Ra value of FeB
decreased from 0.90 µm to 0.09 µm, from 0.95 µm to 0.08 µm and from 1
µm to 0.010 µm, for borided SAE 1020 quality steel substrate for 2h, 4h and
6h boriding time, respectively. Like borided SAE 1020 quality steel substrate,
Ra value of FeB decreased form 1.2 µm to 0.09 µm, from 0.95 µm to 0.07
µm and from 0.89 µm to 0.05 µm for borided SAE 1040 quality steel
substrate for 2h, 4h and 6h, respectively.
5- Hardness determinations of FeB showed that increase in boriding process
time 2h to 6h, FeB layer hardness’s increase from 775 DHV to 1081 DHV,
from 1001 DHV to1277 DHV, from 1033 DHV to 1381 DHV and finally,
1621 DHV to 1788 DHV with decreasing applied peak loads from 640 mN,
320 mN, 160mN and 80 mN for borided SAE 1020 quality steel substrate.
Like borided SAE 1020 quality steel, increase in boriding process time 2h to
6h, FeB layer hardness’s increase from 1015 DHV to1120 DHV, 1045 DHV
to 1210 DHV, from 1367 DHV to 1486 DHV and lastly 1961 DHV to 2244
DHV with decreasing applied four different peak loads from 640 mN to 80
mN for borided SAE 1040 quality steel substrate.
6- Calculations of Young’s modulus of FeB with increase in boriding process
time and decrease in applied peak load, indicated that calculated young’s
modulus increased. The elastic modulus results showed that increase in
process time from 2h to 6h causes elastic modulus increase from 125 GPa to
241 GPa, from 240 GPa to 396 GPa, from 276 GPa to 397 and finally from
432 GPa to 568 GPa with decreasing applied peak loads from 640 mN to 80
mN for borided SAE 1020 quality steel. Similar to borided SAE 1020 quality
steel, increase in process time from 2h to 6h causes, elastic modulus increase
from 273 GPa to 285 GPa, from 285 GPa to 400 GPa, from 345 GPa to 551
GPa, from 496 GPa to 624 GPa with decreasing applied peak loads from 640
mN to 80 mN for borided SAE 1040 quality steel substrates.
7- Fracture toughness of upper layer FeB increases from 0.9188 MPam 1/2 to
1.1265 MPam 1/2 with the increasing process time from 2h to 6h for SAE 1020
quality steel. Like borided SAE 1020 quality steels, fracture toughness of

82
FeB layer increase from 0.4014 0.9188 MPam 1/2 to 1.4257 MPam 1/2 with the
increasing process time from 2h to 6h for SAE 1040 quality steel.
8- FEM study results show that Von Misses Stress and equivalent strain
distribution at contact region between indenter and simulated FeB layer. By
changing yield stress of FeB from 5 GPa to7 GPa in ABAQUS, penetration
depth values were determined under 640 mN peak load. When the yield stress
of FeB layer is chosen 5 GPa under 640 mN applied peak load, the same
penetration depth of 2.13 µm determined in experimental study.
6.2 Future Plan
Is is recommended that further study should be carried out in order to obtain more
desirable single phase boride layer (Fe 2 B). With this aim, pack boriding process
variations such as process time and temperature can be changed to get hold of
Fe 2 B layers on SAE 1020 and SAE 1040 quality steels. In addition, tensile
and compressive residual stresses of layer should be calculated. Because load-unload
curves of layers can change under this parameter.

83
REFERENCES
A.C. Fischer-Cripps, 2001. Use of Combined Elastic Modulus in the Analysis of
Depth-Sensing Indentation Data. Journal of Materials Research, 16, 3050-3052.
A.E. Giannakopoulos and S. Suresh, 1999. Determination of Elastoplastic Properties
by Instrumented Sharp Indentation. Scripta Materialia, vol. 40, 1191-1198.
A. Galibois, O. Boutenko, and B. Voyzelle, 1980. Vol: 28, 1753-1763, 1765-1771.
A. Graf von Matuschka, 1980, Boronizing, Hanser.
A.H. Üçisik, C. Bindal, 1997. Fracture toughness of boride formed on low alloy
steels. Surface and Coatings Tech., 94-95, 561-565.
A. Bolshakov, G.M. Pharr, 1998. Influence of pileup on the Measurement of
Mechanical Properties by Load and Depth Sensing Indentation Techniques.
Journal of Materials Research, vol. 13, 1049-1058.
B.R. Chatterjee-Fischer, and O. Schaaber, 1976. Proceeding of Heat Treatment, The
Metal Society, 27-30.
C.A. Brookes, 1979. Properties of Diamond, Academic Press, New York.
D. Munz and t. Fett, 1999. Ceramics (Mechanical properties, failure behaviour,
material selection). Institute of Materilas Research, Karlsruhe, Germany.
G.M. Pharr, W.C. Oliver, and F.R. Brotzen, 1992. On the Generality of the
realationship Among Contact Stiffness, Contact Area, and Elastic Modulus
During Indentation. Journal of Materials Research, vol. 7 613-617.

84
H.C. Child, 1981. Metallurgy and Materials Technology, Vol: 13, 303-309.
Ian N. Snedon, 1965. The relation between Load and Penetration in The Axisymetric
Boussinesq Problem for a Punch of Arbitrary Profile. Int. J. Engng Sci, vol.3 47-
57.
Ibrahim Ozbek, Cuma Bindal, 2002. Mechanical properties of boronized AISI W4
steel. Surface and Coatings Tech., 154, 14-20.
J. Alcala, A.E. Giannakopoulos, and S. Suresh, 1998. Continuous Measurements of
Load-Penetration Curves with Spherical Microindenters and the estimation of
Mechanical Properties. Journal of Materials Research, vol. 13, 1390-1400.
J.L. He, S. Veprek, 2003. Finite Element Modeling of indentation into Superhard
coatings. Surface and Coatings Tech. 163-164, 374-379.
J.L. Loubet, J.M. Georges, J.M. Marchhesini, and G. Meille, 1984. Vickers
Indentation Curves of Magnesium Oxide, J. Tribology, 106, 43.
K.-D. Bouzakis, N. Michailidis and, G. Erkens, 2001. Thin hard Coatings stressstrain
Curve Determination through a FEM Eupported Evaluation of
Nanoindentation Test Results. Surface and Coatings Tech. 142-144, 102-109.
K.L. Johnson, 1970. The correlation of Indentation Experiments, J. Mech. Phys.
Solids, vol. 18, 115-126.
K.L. Johnson, 1985. Contact Mechanics, Cambridge University Press.
Kurt E. Peterson and C.R. Guarnieri, 1979. Young’s Modulus Measurement of Thin
Films Using Micromechanics. Journal of Applied Physics, vol. 50, 6761-6766.

85
L. De Fazio, S. Syngellakis, R.J.K. Wood, F.M. Fugiuele, G. Sciume, 2001.
Nanoindentation of CVD Diamond: Comparison of an FE Model With Analytical
And Experimental Data. Diamond and related Materials,765-769.
M. Toparli, N.S. Koksal, 2005. Hardness and yield strength of dentin from simulated
nanoindentation tests. Computer Methods and Programs in Biomedicine, 77,
253-257.
M.J.Lu, 1983. Härt Tech. Mitt. Vol: 38, 156-159.
Orhan Uzun, Uğur Kölemen, Selahattin Çelebi and Nusret Güçlü, 2005. Modulus
and hardness evaluation of polycrystalline superconductors by dynamic
microindentation technique. Journal of European Seramic Society, 25, 969-977.
P. Dearnly and T. Bell, Surface Engineering, 1985. Surface Engineering, Vol:1
(no:3), 203-217
P.A. Dearnly, T. Farrell, J. Mater and T. Bell, 1986. Energy System, Vol:8, 128-131.
R. Chatterjee-Ficher, Härt Tech. Mitt., 1981, Vol.36 (no:5), 248-254
R. Chatterjee-Fischer, 1977. Powder Metallurgy, Vol:20, 96-99.
R. Chatterjee-Fischer, 1989. Surface Modification Technologies, T.S. Sudarshan,
Ed., Marcel Dekker, 567-609.
T.B. Cameron and J.E. Morral, 1986. Met. Trans. A, Vol:17A 1481-1483.
T.B. Massalski, 1986. Binary Alloy Phase Diagrams, American Society for Metals.

86
Ugur Sen, Saduman Sen, Sakip Koksal, Fevzi Yılmaz, 2005. Fracture toughness of
borides formed on boronized ductile iron. Materials and Design, vol. 26, 175-
179.
W.C. Oliver, G.M. Pharr, 1992. An Improved Techniques for Determining Hardness
and Elastic Modulus Using Load and Displacement Sensing Indentation. Journal
of Materials Research, vol. 7, 1564- 1583.
W.J.G. Fichtl, 1974. Härt Tech. Mitt., Vol: 29, 113-119.
W.J.G. Ficthl, “Saving Energy and Money by Boronizing,” 1988. Paper presented at
the meeting of the Japan Heat Treating Association, Tokyo.
West Conshohocken, 1987. Standard Test Method for Vickers Hardness of Metallic
Materials. American Society for Testing and Materials.
Xi Chen and Joost J. Vlassak, 2001. Numerical Study on the Measurement of Thin
Film Mechanical Properties by Means of Nanoindentation. Journal of Materials
Research, vol. 16, 2974-2982.
Yang-tse Cheng, che-Min Cheng, 1999. can Stress-Strain relationships be obtained
from Indentation Curves Using Conical abd Pyramidal Indenters? Journal of
Materials Research, vol. 14, 3493-3496.
Ziheng Yao, 2005. Development of an Indentation method for material surface
mechanical properties measurement. Department of Mechanical and Aerospace
Engineering, Morgantown, Virginia.