TOOL WEAR ANALYSIS IN VARIOUS MACHINING PROCESSES AND
STUDY OF MINIMUM QUANTITY LUBRICATION (MQL)
By
Kyung Hee Park

A DISSERTATION
Submitted to
Michigan State University
in partial fulfillment of the requirements
for the degree of
DOCTOR OF PHILOSOPHY
Mechanical Engineering
2010

ABSTRACT
TOOL WEAR ANALYSIS IN VARIOUS MACHINING PROCESSES AND STUDY OF
MINIMUM QUANTITY LUBRICATION (MQL)
By
Kyung Hee Park

The tool wear analysis on the multilayer coated carbide inserts in turning and milling of
AISI 1045 steels was performed using advanced microscope and image processing techniques. In
turning process, the flank wear evolution, surface roughness and groove sizes on the coating
layers were analyzed to understand the flank wear mechanism(s) involved. The dominant wear
phenomenon was abrasion and, after carbide was exposed, adhesion took over. For flank wear
prediction, 2-body abrasion model was used along the interface conditions from finite element
(FE) model, which provides the temperature on the cutting tool. In a face milling study,
multilayer cutting tools, double (TiN/TiAlN) and triple (TiN/Al2O3/TiCN) layered coated
carbide, processed by physical vapor deposition (PVD) and chemical vapor deposition (CVD)
respectively, were evaluated in terms of various cutting conditions. Similar to the turning case,
abrasion was found to be the most dominant tool wear mechanism in milling. Edge chipping and
micro-fracture were the tool failure modes. Overall, the double layer coating was superior to the
triple layer coating under various cutting conditions due to the benefit coming from the coating
deposition processes themselves.
On the other hand, drilling of carbon fiber reinforced polymer (CFRP)/titanium (Ti)
stacks has been performed using carbide and PCD drills. The dominant wear mechanism for
carbide tools was the abrasion by fibers in CFRP drilling and Ti adhesion, covering entire cutting

edge while the main wear behavior of PCD drill was edge chipping. The adhesion of titanium
seems to be the most important factor in tool wear and micro-chipping because the adhered
titanium seems to be brushed away by the fibers in CFRP when drilling the composites, which
accelerates tool wear in drilling of stacks. PCD drills have more wear resistance than carbide
drills only in terms of the flank wear land. However, due to the inherent brittleness of PCD drills,
chipping was observed at the tool edge when drilling titanium.
Finally, for a better understanding of Minimum Quantity Lubrication (MQL) and its
effective use in practical industrial applications, MQL parameters such as droplet sizes, the
droplet distributions and wetting angles of various types of lubricants were investigated. The
measurement method of droplet size and distribution was proposed using confocal laser scanning
microscopy (CLSM) and wavelet transform. In addition, the empirical droplet size estimation
equation was also introduced for extremely small droplets, which are difficult to be measured.
The distribution of the droplets has been also studied to determine the MQL optimal nozzleworkpiece distance and the nozzle discharge pressure. To extent the applicability of MQL to
more aggressive machining conditions, a potential additive to MQL lubricant, which is a mixture
of exfoliated nano-graphene particles and vegetable oil, have developed. Oil based lubricant,
especially nano-enhanced vegetable oil, showed a better wettablility and tribological behavior.
MQL-ball milling tests with nano-graphene enhanced lubricant were performed to show a
remarkable performance improvement in reducing both central wear and flank wear as well as
edge chipping.

To my wife, Jihyun,
my children, Suin, Jeongin, and Haein,
and my beloved family

iv

ACKNOWLEDGEMENTS

I would like to gratefully and sincerely thank my advisor Dr. Patrick Kwon for his
support, patience, and excellent guidance throughout my graduate studies. His technical and
editorial advice was essential to the completion of this dissertation and has taught me
innumerable lessons and insights on the workings of the research.
My appreciation goes to the committee members Dr. Brian Feeny, Dr. Dashin Liu, and
Dr. Jung-Wuk Hong for providing many valuable comments that improved the contents of this
dissertation.
My thanks also go to Dr. Gyu-Bong Lee for giving me the opportunity to study at
Michigan State University and supporting me to finish this dissertation.
Many thanks go to my colleagues, Jorge Olortegui-Yume, Li Sun, Xin Wang, David
Shrock, Brent Ewald, Matt Lempke, Justin Freeman, and John Tysman for supporting this
research effort.
Last, but not least, I would like to thank my wife, Jihyun Noh, for her understanding and
love during the past few years, along with my three children: Suin (Aileen), Jeongin (Olivia), and
Haein (Eric). Her unbreakable support and encouragement was, in the end, what made this
dissertation possible. My deepest gratitude also goes to my family especially for my parents and
parents-in- law for their dedication and the many years of support during my PhD study.
A special acknowledgement goes to Advanced Cutting Tool Technology Center by NSFI/UCRC program for their support and also to Fraunhofer CCL at MSU for the experiments and
equipments.

v

TABLE OF CONTENTS

LIST OF TABLES ....................................................................................................................... viii
LIST OF FIGURES ....................................................................................................................... ix
Chapter 1 ..........................................................................................................................................1
INTRODUCTION ...........................................................................................................................1
1.1 TOOL WEAR TYPES ...............................................................................................................7
1.1.1 Abrasion ...........................................................................................................................7
1.1.2 Adhesion...........................................................................................................................9
1.1.3 Dissolution .......................................................................................................................9
1.2 IMAGE PROCESSING TECHNIQUE ...................................................................................11
1.3 2D WAVELET TRANSFORM ...............................................................................................13
1.4 FEM SIMULATION ...............................................................................................................15
Chapter 2 ........................................................................................................................................17
FLANK WEAR OF MULTILAYER COATED CARBIDE TOOL AND WEAR PREDICTION
USING 2-BODY ABRASIVE WEAR MODEL ....................................................................17
2.1 FLANK WEAR MODEL AND IMAGE PROCESSING TECHNIQUE ...............................20
2.1.1 Abrasive wear model ......................................................................................................20
2.1.2 Finite element model ......................................................................................................20
2.1.3 Image acquisition and wavelet transform.......................................................................24
2.2 EXPERIMENTAL SETUP......................................................................................................25
2.3 RESULTS AND DISCUSSION ..............................................................................................26
2.3.1 Roughness analysis ........................................................................................................28
2.3.2 Flank wear analysis ........................................................................................................38
2.4 CONCLUSION ........................................................................................................................48
Chapter 3 ........................................................................................................................................50
TOOL WEAR ANALYSIS ON MULTI-LAYERED COATED CARBIDE TOOLS IN FACE
MILLING ................................................................................................................................50
3.1 EXPERIMENTAL SETUP......................................................................................................53
3.2 TOOL WEAR MEASUREMENT...........................................................................................54
3.3 FLANK WEAR ANALYSIS...................................................................................................57
3.4 CONCLUSION ........................................................................................................................68
Chapter 4 ........................................................................................................................................69
TOOL WEAR IN DRILLING OF COMPOSITE/TITANIUM STACKS WITH CARBIDE AND
POLYCRYSTALLINE DIAMOND TOOLS .........................................................................69
4.1 TOOL WEAR IN DRILLING .................................................................................................70
4.1.1 CFRP drilling .................................................................................................................70
4.1.2 Ti drilling........................................................................................................................71
4.1.3 Hybrid structure drilling .................................................................................................73
vi

4.2 EXPERIMENTAL PROCEDURES ........................................................................................74
4.2.1 Workpiece materials .......................................................................................................74
4.2.2 Drilling experiments .......................................................................................................74
4.2.3 Wear evolution analysis .................................................................................................77
4.3 RESULTS AND DISCUSSION ..............................................................................................78
4.3.1 Drilling forces ................................................................................................................78
4.3.2 WC drill wear .................................................................................................................80
4.3.3 PCD drill wear................................................................................................................91
4.4 CONCLUSION ........................................................................................................................97
Chapter 5 ........................................................................................................................................98
A STUDY ON DROPLETS AND THEIR DISTRIBUTION FOR MINIMUM QUANTITY
LUBRICATION (MQL) .........................................................................................................98
5.1 EXPERIMENTAL SETUP....................................................................................................101
5.2 CLSM DATA ACQUISITION AND IMAGE PROCESSING.............................................103
5.3 DROPLET VOLUME CALCULATION ..............................................................................106
5.4 RESULTS AND DISCUSSION ............................................................................................107
5.4.1 Droplet size and distribution ........................................................................................107
5.4.2 Droplet volume calculation by wavelet transform .......................................................111
5.4.3 Small droplet size estimation .......................................................................................121
5.5 CONCLUSION ......................................................................................................................125
Chapter 6 ......................................................................................................................................127
MQL WITH NANOGRAPHENE-ENHANCED LUBRICATES: BALL MILLING
EXPERIMENT......................................................................................................................127
6.1 NANO-GRAPHENE ENHANCED LUBRICANT ..............................................................130
6.2 EXPERIMENTAL SETUP....................................................................................................132
6.2.1 Wetting angle measurement .........................................................................................132
6.2.2 Tribometer test .............................................................................................................133
6.2.3 Ball milling test ............................................................................................................134
6.3 RESULTS AND DISCUSSION ............................................................................................137
6.4 CONCLUSION ......................................................................................................................149
Chapter 7 ......................................................................................................................................150
CONCLUSIONS..........................................................................................................................150
BIBLIOGRAPHY ........................................................................................................................153

vii

LIST OF TABLES

Table 1 Material constants for the JC model .................................................................................22
Table 2 Material properties of tool and work material .................................................................23
Table 3 Thickness of coatings........................................................................................................26
Table 4 Roughness data on the coatings ........................................................................................34
Table 5 Cutting conditions .............................................................................................................54
Table 6 CTE of the tool materials [Wang et al., 1999 and Zhao et al., 2003] ...............................63
Table 7 Drill geometries ................................................................................................................76
Table 8 Drilling experiment conditions .........................................................................................77
Table 9 SEM images for carbide drills at high and low speeds with various holes ......................81
Table 10 SEM images of the PCD tool at low cutting speed.........................................................92
Table 11 Average sizes, droplet count, and area fractions ..........................................................111
Table 12 Droplet size calculation ................................................................................................124
Table 13 Droplet size distribution................................................................................................124
Table 14 Machining conditions ...................................................................................................136

viii

LIST OF FIGURES

Figure 1 Hardness of cementite and coating materials ....................................................................8
Figure 2 Typical configuration of CLSM [Olortegui-Yume and Kwon, 2007] ............................12
Figure 3 Decomposition step of two-dimensional discrete wavelet transform..............................14
Figure 4 Representation of wavelet filter bank (a) decomposition and (b) reconstruction ...........14
Figure 5 Flank wear land (VB) and boundary conditions for FE simulation .................................21
Figure 6 (a) HE and (b) MB images from CLSM at 7min cutting time and (c) raw z-matrix and
(d) wavelet filtered images from HE image ..................................................................25
Figure 7 Evolution of flank wear observed by SEM ....................................................................27
Figure 8 BSE images on flank surface ..........................................................................................27
Figure 9 CLSM image for the flank surface at 4 min (a) and wavelet decoupled components (b)
at Level 3. ......................................................................................................................29
Figure 10 Roughness, Waviness and Form....................................................................................30
Figure 11 Roughness Data on Tool after Cutting for 1 min ..........................................................33
Figure 12 Roughness data and Groove sizes on the tool after cutting of 7 min ............................34
Figure 13 Roughness on the tool after cutting for 23 min after etching of steel ...........................35
Figure 14 Microstructure (a) and a pearlitic colony size (b) of refined AISI 1045 steel. ..............37
Figure 15 3D and cross-sectional profiles of raw data and wavelet-filtered at level 2 ..................40

ix

Figure 16 Flank wear profiles and wear rates. ...............................................................................41
Figure 17 Adhesion on the carbide and coating delamination at cutting time of 23 min. .............42
Figure 18 Notch wear on the tools after 20 and 23 min of cutting ................................................44
Figure 19 Notch wear after etching (a) and scoring marks and notch wear zone (b) ....................45
Figure 20 Temperature profiles on cutting tools from FEM simulations ......................................46
Figure 21 Predicted flank wear based on 2-body wear model. ......................................................47
Figure 22 Schematic of the face milling ........................................................................................53
Figure 23 CLSM images for flank wear land (VB) .......................................................................55
th

Figure 24 HE image of flank wear after 5 layer cutting. ...........................................................56
th

Figure 25 3D surface profile from HE image after 5 layer cutting .............................................56
Figure 26 EDS images of flank wear. ............................................................................................56
Figure 27 Flank wear (VB) for triple layer coating at various cutting conditions. .......................58
Figure 28 Flank wear (VB) for double layer coating at various cutting speeds.............................59
Figure 29 Flank wear comparison of double (PVD) and triple (CVD) layer coating....................61
Figure 30 Indentation tests of PVD and CVD coatings .................................................................61
Figure 31 Flank wear evolutions of double layer coating..............................................................64
Figure 32 Flank wear evolution of triple layer coating at 300m/min ............................................66
x

Figure 33 Schematics of drilling experimental set up and drilling force data acquisition system 76
Figure 34 Cutting forces vs number of holes drilled by the different tool types and conditions...79
Figure 35 Drill geometry schematics and CLSM images of the WC drill outer corner at high
speed after 40 hole .........................................................................................................81
Figure 36 CLSM images at flank face of WC drill after high speed drilling at 60th hole ............83
Figure 37 Tool flank wear profiles of carbide drills before and after etching at various speeds...86
Figure 38 Flank and edge wear progression of WC drills for various drilling speeds ..................88
Figure 39 Effect of wear progression on drilling forces. ...............................................................90
Figure 40 Micro-fractures observed on flank side of PCD at low speed drilling and at hole 60 .93
Figure 41 CLSM images and flank wear profiles of PCD drill before and after etching at low
cutting speed ..................................................................................................................95
Figure 42 Flank and edge wear progression of the PCD drill in low speed ..................................96
Figure 43 MQL Experimental Set-up ..........................................................................................102
Figure 44 Droplets distribution optical images from CLSM and bitmap images obtained by EDA
at 4psi and 8psi ............................................................................................................104
Figure 45 Droplet images for droplet size measurement .............................................................105
Figure 46 Average mass accumulation, sprayed CLSM images, and corresponding bitmap
images at nozzle distances and air pressures ...............................................................108
Figure 47 Flow chart for the calculation of the droplet size and distribution ..............................112

xi

Figure 48 HE image of the droplet (a), 3D surface profile (b) and 2D droplet profiles in x (c) and
y section (d) at level 3 .................................................................................................113
Figure 49 The Droplet surfaces after wavelet filtering ................................................................116
st

th

Figure 50 Droplet surface after wavelet transform in 1 to 6 level .........................................118
Figure 51 Droplet surface at level 3 (a) and its corresponding estimated droplet volume (b) and
equivalent radius (c) ....................................................................................................120
Figure 52 Droplet volumes corresponding 2D droplet sizes .......................................................123
Figure 53 (a) SEM and (b) TEM images of xGnP .......................................................................131
Figure 54 Suspension stability of xGnP oils (A: 1μm / 0.1wt%, B: 1μm / 1.0wt%, C: 15μm /
0.1wt%, D: 15μm / 1.0wt%) .......................................................................................131
Figure 55 Wetting angle of a droplet ...........................................................................................132
Figure 56 Wetting angle measurement setup ...............................................................................133
Figure 57 Tribometer setup ..........................................................................................................134
Figure 58 MQL and ball milling setup ........................................................................................135
Figure 59 Tool wear measurements of a ball nose insert.............................................................136
Figure 60 Wetting angle test results.............................................................................................137
Figure 61 Tribometer test results for various lubricants ..............................................................139
Figure 62 Comparison of friction coefficients of dry and xGnP coating ....................................140
Figure 63 Comparison of friction coefficients of lubricants........................................................142

xii

th

Figure 64 Central wear at 3500rpm at 8 layer...........................................................................143
th

Figure 65 Flank wear at 4500rpm at 8 layer .............................................................................144
Figure 66 Central wear and flank wear in terms of concentration of xGnP ................................146

xiii

Chapter 1
INTRODUCTION

Metal cutting is a complex process where many factors, such as machine tool, work
material, tool material and machining conditions, have to be considered. Among the factors, tool
wear is one of the most important aspects that affect product quality and cost in metal cutting
process. Many research studies regarding of the tool wear have been carried out over 150 years.
However, no significant work to understand the mechanism behind tool wear has taken place
mainly due to the complexity in metal cutting process. Several tool wear types have been
reported in metal cutting including abrasion [Ramalingam and Wright, 1981], adhesion [Wright
and Bagchi, 1981], diffusion [Trent, 1963] and dissolution [Kramer and Kwon, 1985; Kramer
and Suh, 1980]. In general, abrasive and adhesive wears dominate at low cutting temperatures
while the chemical wear such as diffusion and dissolution are dominant at high cutting
temperatures [Stephenson and Agapiou, 2006]. Thus, the dominant wear mechanism at flank
surface due to relatively low cutting temperature compared to rake surface is abrasion while the
combination of abrasion and dissolution wears takes place at rake surface. In addition, the
abrasion is strongly dependent upon the hardness of a tool material; therefore, a harder coating
can improve the tool life if the abrasive wear is dominant. With the wear mechanisms identified;
the effort to develop physics based tool wear models has been made to predict tool life because
an empirical approach, that has dominated tool wear studies, only relies on experimental tests for
each combination of tool and work material [Rabinowicz, 1964, Wong et al. 2004, Park and
Kwon, 2009, and Olortegui-Yume and Kwon, 2010]. Based on the physics based wear model, it
is possible to describe tool wear behaviors of various combinations of work and tool material,

1

such as various alloys/composites and single/multilayer coated tools, respectively, without a
significant amount of machining experiments.
This study deals with three aspects that directly related to tool wear, which are tool
(coating) material, work material and machining process because understanding these aspects
can eventually improve the productivity.
First, in chapter 2, the tool wear mechanism of multilayer coating is discussed for turning
process. In general, it is known that the multi-layer coated carbide tools can improve the tool life
[Cselle and Barimani, 1995, Olortegui-Yume and Kwon, 2007, and Olortegui-Yume et al., 2008].
However, the reason for the wear deterrence of multi-layered coating is not well established.
Recently, in the case of turning of AISI 1045 steel with multi-layer (TiN/Al2O3/TiCN) coated
carbide tools, the explanation has been made [Olortegui-Yume and Kwon, 2007 and OlorteguiYume et al., 2008]. The reason is that the continuous turning process increases the cutting
temperature high enough to activate the dissolution wear in addition to abrasive wear at rake.
However, the Al2O3 coating is known to be resistant to the crater wear where the dissolution
wear dominates while hard coatings such as TiCN exhibit excellent performance against abrasive
wear. In addition, with the exposure of each coating, the locations of maximum cutting
temperature, which coincide the location of maximum crater depth, change because the friction
condition varies in terms of coating materials. This delays the crater wear evolution. On the other
hand, at flank surface, the dominant wear mechanism is abrasion, which can be interpreted as
only harder coating (TiCN) is required for flank wear [Park and Kwon, 2009]. In addition, it was
commonly believed that the effectiveness of multilayer coating comes from the delay in the
crack propagation [Sadik and Myrtveit, 2009] because the crack can propagate along the coating
interfaces instead of right through the coatings into the substrate. Under the cyclic loading in

2

milling process, enhanced toughness and wear resistance against thermal and mechanical impact
are required and again multilayer coating has been cited to improve the life of cutting tools
[Cselle and Barimani, 1995].
Coatings are deposited onto carbide tools in a single-layer or multi-layer scheme
processed by CVD and/or PVD techniques, which have advantages and disadvantages
[Stephenson and Agapiou, 2006 and Soldergberg et al., 2001]. For example, CVD-Al2O3 in
multi-layered coating provides thermal barrier and good bonding with other coatings. However,
the brittle eta phase can be formed at the coating interface due to high deposition temperature,
which could deteriorate the strength of the substrate and weaken the cutting edge. On the other
hand, PVD coating process produces the sharp cutting edge and free thermal cracks on the
cutting surface. In this regard, the coating materials along with the coating process must be
determined depending on machining applications. Jawaid et al. [2000] found that the CVD
coating tool was superior to the PVD coating tool due to high adhesion strength of the CVD
coating tool despite of the inherent disadvantage from the CVD coating process. Dobranski et al.
[2006] compared CVD and PVD in nitride tool ceramics on a ‘pin-on disk’ method and
concluded that the multilayer coating, processed by CVD, was better due to the adhesion strength.
Therefore, in chapter 3, tool wear of multilayer coating in milling is studied in terms of coating
deposition processes.
In chapter 4, the drilling study of carbon fiber reinforced polymer (CFRP)/titanium (Ti)
stack is presented because CFRP/Ti stack has been increasingly used in aerospace industry.
However, drilling of CFRP/Ti stack is challenging due to the dissimilar machining
characteristics of CFRP and Ti, where each requires the distinct optimal cutting condition. In fact,
it is known that the ideal tooling solution for drilling Ti is not the most cost effective for CFRP

3

drilling while the ideal tooling solution for CFRP cannot cut Ti effectively. The different
behaviors of the materials in the stack in drilling can often lead to severe tool wear, which often
causes increased drilling forces, poor hole quality and large Ti burrs in CFRP-Ti stack drilling
[Kim and Ramulu, 2004]. The CFRP can be a highly abrasive material depending on the
direction of the fiber in the matrix while drilling for titanium causes high cutting temperature due
to its low conductivity. Considering reduced assembly flow time, reduced cost and improved
machined hole quality, tool wear in drilling CFRP/Ti stacks has to be minimized. The possible
wear types on drill bits including flank wear, crater wear, chisel edge wear, margin wear and
chipping wear were characterized [Kanai and Kanda, 1978, Sanjay et. al, 2005]. Rawat and Attia
[2009] investigated the wear mechanism of a tungsten carbide drill in drilling of woven graphite
epoxy composites at high cutting speed. The chipping and abrasion by the fractured graphite
fibers and tungsten carbide grains lead to flank wear evolution which results in delamination and
reduce the hole quality of the work material as well as the cutting forces. PCD tool showed much
less wear rate and higher hole quality compared to HSS and carbide tools when drilling a
graphite/bismalemide composite [Ramulu et al., 1999]. Sharif and Rahim [2007] observed from
Ti drilling experiments with TiAlN-coated and uncoated carbides that TiAlN-coated carbide
tools were superior to uncoated carbide tools in terms of flank wear at various cutting speeds and
both drills showed similar tool failure modes such as chipping and catastrophic failure. A PCD
drill can fracture when drilling Ti at high cutting speeds mainly due to the brittleness of the PCD
tool [Garrick, 2007].
In chapter 5, the minimum quantity lubrication (MQL) machining process is discussed.
MQL has been applied to reduce the cost associated with the maintenance and disposal of the
fluid being used and the conformation to environmental policies [Ueda et al., 2006 and Rahman

4

et al., 2002]. In MQL, the high pressurized air sprays micro-scale lubricant droplets onto the
tool-work interface. Even though it is very promising for many machining applications, several
fundamental issues associated with MQL such as optimal operation conditions have not been
addressed [Park et al., 2008]. The most important point in MQL is that the lubricant must
adequately cover the contact areas, i.e., tool-chip and tool-work interfaces. And another
important variable is the wettability of the lubricant on the tool surface. On the other hand, the
optimal MQL application conditions such as air pressure, nozzle direction, distance, etc., have to
be determined for its effective use prior to machining applications. Therefore, a better
understanding of the application conditions such as droplet size and distribution, air pressure,
nozzle distance, and flow rate is needed to be effective in industrial applications. Thus, a great
interest exists in the machining community in the development of a convenient technique to
measure the droplet size. The smaller droplet, which can provide better penetration into the
cutting surface, are more preferable especially for micro machining applications, where the
cutting surface is relatively small [Jun et al., 2008].
Finally, chapter 6 studies the effect of the MQL lubricant including exfoliated nanographene enhanced vegetable oil by conducting wetting angle measurement, friction tests using a
ball-on-disk set-up and MQL ball milling test. Several works have been published to evaluate the
characteristics of various lubricants for finding optimum lubricants, e.g. oil film on water droplet,
in MQL machining [Heinemann et al., 2006, Itoigawa et al., 2006 and Yoshimura et al., 2005].
In addition, solid lubricants (nano-sized) such as molybdenum disulfide (MoS2), graphite and
boron nitride have been developed as a lubricant itself or as an additive for lubricant [Jianxin et
al., 2009, Krishna and Rao, 2008, Xu et al., 2004 and Reddy and Rao, 2006]. Especially nanographene enhanced lubricants are expected to provide better performance because nano-graphene

5

has not only a higher dissociation temperature but also the high aspect ratio, which naturally
orients the graphene phase for better lubricity when applied through MQL process. Therefore, in
this chapter, the wetting angles for a variety of commercially available lubricants including nanographene enhanced oil were tested to evaluate the wettability of lubricants. And the friction
coefficients of each lubricant and a dry nano-graphene coated insert were measured as a function
of load and speed. Finally, ball milling tests were performed to evaluate the proposed nanographene enhanced lubricant in a practical machining application.
This research work investigates tool wear mechanisms at flank surface in various
machining processes, such as turning, milling and drilling, using the combination technique of
CLSM and image processing technique including wavelet transform. This technique enables us
to characterize 3D topographic profile on flank and ultimately obtain the tool wear evolution in
cutting times or distances. In addition, in case of turning test, 2-body abrasive model along with
FEM simulation predicts the flank wear and compares with experimental results. On the other
hand, a new technique is proposed to measure the droplet sizes and distributions for a typical
MQL process after the droplets have been sprayed onto the polished silicon wafer surface.
Furthermore, MQL parameters such as wetting angle, nozzle distance and air pressure, which
may affect the machining efficiency, are evaluated for optimal MQL conditions. To evaluate the
effect of MQL, ball milling test, based on the MQL parameters determined, is also performed in
both dry and MQL conditions in addition to tribological test.

6

1.1 TOOL WEAR TYPES
1.1.1 Abrasion
Abrasive wear takes place by hard inclusion in work material, which results in the
scoring marks on the tool surface (See Figure 19). In the pearlitic microstructure, 2-body
abrasion occurs because a complex morphology of the second phase is constrained within a
matrix of ferrite while 3-body abrasion can be used because the hard second phase rolls at the
cutting interface face. The 2-body and 3-body abrasive wear models were developed by
Rabinowicz [1964]. For 2-Body abrasion, the wear volume, V B , is determined to be

V

·F·
θ
π·H

B

For 3-Body abrasion, the wear volume,

V
V

B

V

B

B , is determined to be

·F·
θ
·H

·F·
θ H
. ·H
H

B

V

(1)

.

;

H
H

0.8
H
H

; 0.8

·F·
θ H
. ·H
H

;

H
H

1.2

,

(2)

1.25

where x is the sliding length, F is the load between interacting surfaces, θ is the average
roughness angle of the abrasive, H is the hardness of the abraded material, H
7

is the hardness

of the abrasive. For the abrasive model, hot hardness data for coating materials and abrasives
(Fe3C for low alloy steels) are required [Kwon and Kountanya, 1999; Kwon, 2000; and Wong et
al., 2004] as seen in Figure 1.

Figure 1 Hardness of cementite and coating materials
(For interpretation of the references to color in this and all other figures, the
reader is referred to the electronic version of this dissertation.)

8

1.1.2 Adhesion
Adhesive wear takes place when one solid material is sliding over the counteracting
surface. The interaction between two surfaces can be represented by the metallurgical weld or
adhesion joint (See Figure 17 (a)). Adhesive wear takes place by pulling out discrete pieces from
the surfaces as they are sliding and separated [Rabinowicz, 1995]. In addition, Rabinowicz,
[1964] found the quantitative laws of adhesive wear from the experiment as follows; the amount
of wear (1) is directly proportional to the load L between interacting surfaces, (2) proportional to
the distance slid x; and (3) inversely proportional to the hardness p of tool material being worn
away. In metal cutting, adhesion may occur because the work material or the wear debris from
the work material forms strong bonds (junction) with the tool under the high interfacial pressure
and temperature. If the failure strength of the junction formed is larger than the local failure
strength of one of the sliding surfaces, then, the junction will be detached from the surface with
the lower failure strength. Thus, Adhesive wear will occur if the lowest failure strength happens
to be on the tool. Adhesive wear peaks at intermediate temperatures due to low adhesion
tendency at low temperatures and thermal softening at high temperatures [Childs et al., 2000].

1.1.3 Dissolution

Dissolution wear occurs at a rake surface due to high cutting temperature generated by
severe machining condition. In dissolution wear, the species from the tool material forms a solid
solution within the chip. As the chip passes, tool material is constantly removed from the rake
face. Dissolution wear depends on the chemical inertness of the tool material as well as the
chemical affinity between tool component and work material. A quantitative model was
developed by [Kramer 1979, and Kramer and Suh, 1980]. For a given ternary tool material of
9

composition, A

B C

, the chemical solubility was obtained from thermo-dynamic

equilibrium condition,

∆GA

B C

x∆GA

y∆GB

z∆GC ,

(3)

where ∆GA B C is the free energy of formation of the tool material at the interface of chip
and tool, T is the solution temperature, and ∆G (i = A, B, or C) is the relative partial molar
free energy of component i within the solution and can be expressed as

∆G
where ∆G

∆G

RTlnS

,

(4)

is the molar excess free energy of component i in the solution, R is the universal

gas constant, T is the temperature in Kelvin, and S is the solubility concentration of species

i

in mole fraction. From Eq. (4) and data on S at some given T, the excess free energy of
component i into the matrix of a work material can be estimated. Furthermore, ∆G

is taken

as a constant in the low limit of solubility (Henry’ law). For example, a binary compound, TiN,
has only two terms in the evaluation of free energy of formation as

∆GA

B

y∆GB , where A =Ti, x =1, B =N and y =1. Finally, with the definitions of

10

x∆GA

∆G
N

x∆GA
x

y

A B C

y∆GB

z∆GC

z , the chemical solubility,

SA

,M

x · lnx

y · lny

z · lnz and

B C , of a tertiary coating material,

, in a work material (in mole fraction):

SA

B C

exp

∆GA

B C

∆G
NRT

RTM
(5)
,

The dissolution of a tool material is assumed to obey Henry’s law because the amount of
tool material dissolved into chip during the cutting process is small enough to be considered
dilute. Dissolution wear has been successful to predict the tool wear in carbide tools, coated
carbide tools [Wong et al., 2004 and Olortegui-Yume, 2009].

1.2 IMAGE PROCESSING TECHNIQUE

The important characteristic of CLSM is the confocal aperture which eliminates the
reflected light from above and below the focal plane, thus obtaining surface information from a
very thin “optical slice” of the surface is feasible. Surface reconstruction is achieved by storing
and manipulating a stack of optical slices within a set of the vertical travel of the stage. The final
output of the built-in software in confocal microscopes is a z-matrix (z(x,y)) containing the
height information of the surface. For a better handling of the image data, the z-matrix is
encoded in a grayscale, called height encoded (HE) image. Along with HE images, maximum
brightness (MB) images, which are extended depth of field views, can be also obtained. It is
11

important to note that most confocal systems can work in both confocal and optical
(conventional) modes. The basic configuration of CLSM is shown as in Figure 2.

Figure 2 Typical configuration of CLSM [Olortegui-Yume and Kwon, 2007]

12

1.3 2D WAVELET TRANSFORM

Many works have been carried out for surface analysis using wavelet transform [Lee et
al., 1997, Jiang et al., 1999, Raja et al., 2002, and Fu et al., 2003]. The key advantage of the
wavelet transform is the capability of decomposition of the surface profile to various wavelength
regimes, which represent the surface conditions such as roughness, waviness and form, while
maintaining the surface details [Raja et al., 2002, and Olortegui-Yume and Kwon, 2010]. During
the decomposition process, mother wavelet, which can be determined based on characteristic of
the surface texture, is required. Using the mother wavelet chosen, for example, 1D discrete
wavelet transform (DWT) can bring out two coefficients, approximation and detail, at each level
while decomposition process with two channel filter banks, low-pass and high-pass filters in
multi-resolution (MR) scheme. By reverse process of the decomposition, called reconstruction or
inverse discrete wavelet transform (IDWT), the coefficients obtained can be transformed into the
approximation and detail, which are related to the “filtered” surface profile, “form”, (low
frequency) and surface waviness/roughness (medium to high frequency), respectively [Jiang et
al., 1999, Josso et al., 2002, and Yuan et al., 2005]. For 3D topography, in similar way, 2D DWT
analysis decouples raw 3D data to 4 sets of wavelets coefficients, which are called
approximation (cA), horizontal details (cH), diagonal details (cD) and vertical details (cV) as
shown in Figure 3. The surface can be transformed as many levels as possible without losing the
original shape. Then 2D IDWT or 2D wavelet analysis reconstruction for these coefficients are
needed to obtain the approximation (A), horizontal (H), diagonal (D) and vertical (V) details
(See Figure 9 (b)). For the surface roughness analysis, roughness, waviness and form have to be
defined in terms of wavelet decoupled components (A, H, D and V). From the literature
[Olortegui-Yume and Kwon, 2010, and Jiang et al., 1999], the roughness can be defined as the
13

sum of details (H+D+V) up to the cut-off level where the scoring mark is observed. In addition,
the approximation at the cut-off level can be considered as form. In this work, 2D-DWT is used
in a multi-resolution scheme with a two-channel filter bank based on a chosen mother and
scaling wavelet [Rioul and Vetterli, 1991] as in Figure 4.

Figure 3 Decomposition step of two-dimensional discrete wavelet transform

(a) Wavelet decomposition filter bank
Figure 4 Representation of wavelet filter bank (a) decomposition and (b)
reconstruction
14

Figure 4 (cont’d)

(b) Wavelet reconstruction filter bank

1.4 FEM SIMULATION
In the development of flank wear model, the interfacial temperature at the cutting zone is
required since the tool wear is strongly dependent on the cutting temperature. However, the
interfacial condition can be determined neither experimentally nor theoretically. Thus, Finite
Element (FE) simulation of machining process has been widely used for obtaining the
temperature on the cutting surface. In addition, Arbitrary Lagrangian Eulerian (ALE)
formulation in the adaptive mesh technique has been also utilized for maintaining the high
quality mesh throughout the simulation due to the severe mesh deformation associated with the
machining process. The multi-layer coated tool can be modeled as a thermally ‘equivalent’
conductivity of a single composite coating layer on carbide instead of modeling each layer
individually because of the difficulty in modeling thin coating layers. This simplification is
justified as the simulation with the composite coating showed similar results with individual

15

coating layers under the same cutting conditions [Yen et al., 2004, and Olortegui-Yume et al.,
2008].
The J-C constitutive model, suited for the simulation of metals subjected to large strains,
high strain rates and high temperature, is used for AISI 1045 steel. It can be expressed as

σ

A

where is the stress,

(

ε

1

Bε

Clnε

ε is the equivalent strain, ε

1
ε

ε

T

,

(6)

is the dimensionless strain rate

=1.0s-1) and A, B, C, n and m are the material constants from [Ozel and Zeren, 2005].

T
and T

is equal to

T

T

T

T

where T is a reference temperature (25°C)

is the melting temperature of the workpiece.

16

Chapter 2
FLANK WEAR OF MULTILAYER COATED CARBIDE TOOL AND WEAR
PREDICTION USING 2-BODY ABRASIVE WEAR MODEL

Flank wear is one of the most important aspects that affect tool life and product quality in
machining process. However, only few works were published to identify the mechanism behind
flank wear mainly due to the complexity in metal cutting process. At the present time, a
dominant flank wear mechanism is believed to be the abrasion by the hard inclusions in a work
material, which results in the scoring marks on the flank surface [Ramalingam and Wright, 1981].
Recently, 2-body and 3-body abrasion models have been developed for the tool wear prediction
in turning of pearlitic and spherodized steels with various coated carbides [Wong et al., 2004]. It
was observed that the 3-body wear model shows an excellent agreement with the experiment
results of spherodized steels while the 2-body model predicts the flank wear trend for pearlitic
steel with some deviations due to the phase transformation during cutting.
Another important wear mechanism for flank wear is adhesion. This type of wear can
take place when one solid material is sliding over the counteracting surface. The interaction
between two surfaces can be represented by the metallurgical weld or adhesion joint. Adhesive
wear takes place by pulling out discrete pieces from the surfaces as they are sliding and separated
[Rabinowicz, 1964]. On the other hand, excessive localized wear at the depth of cut line on the
flank surface is called notch wear. The main concern for the notch wear is that it often leads to
catastrophic tool failure during the machining. Chandrasekaran and Johansson, [1994]
investigated the mechanisms for notch wear in machining of various austenitic stainless steels
with cemented carbide tools. They concluded that the sequence of the severe shear deformation

17

and lateral extension of the chip leads to excessive abrasion or adhesion wear between the
deformed chip and the tool insert at the notch region.
In the development of flank wear model, the interfacial temperature profile at the cutting
zone is essential since the tool wear is strongly dependent on the cutting temperature. However,
the interfacial condition can be determined neither experimentally nor theoretically. Thus, Finite
Element (FE) simulation of machining process has been widely used for obtaining the
temperature on the cutting surface. In addition, Arbitrary Lagrangian Eulerian (ALE)
formulation in the adaptive mesh technique has been also utilized for maintaining the high
quality mesh throughout the simulation in order to avoid the severe mesh deformation associated
with the machining process. The multi-layer coated tool can be modelled as a thermally
‘equivalent’ conductivity of a single composite coating layer on a carbide tool instead of
modelling each layer individually because of the difficulty in modelling thin coating layers. This
simplification is justified as the simulation with the composite coating showed similar results
with individual coating layers under the same cutting conditions [Yen et al., 2004 and OlorteguiYume et al., 2008].
Olortegui-Yume and Kwon [2010] have conducted a combined experimental and
analytical study to reveal the benefit of multilayer coatings for crater wear based on turning
experiment with multilayer (TiN/Al2O3/TiCN) coated carbide, the followings have been
concluded,
1. The effectiveness of the multilayer comes from the changes in interface conditions as
each layer is exposed. As each layer is exposed, the maximum wear location changes instead of
concentrating on one location. This extends the tool life substantially.

18

2. The middle layer, Al2O3, resisted tool wear at relatively high cutting speeds due to the
excellent resistance against dissolution wear while TiCN provided an excellent resistance against
abrasive wear. Therefore, the multilayer coating made of Al2O3 and TiCN provides the
combined effect to resist crater wear whose wear mechanism is a combination of abrasive and
dissolution wear.
Does the multilayer coating deter the progress of flank wear? It was reported that the
benefit of the multilayer coating is realized by the crack propagation through the coating
interfaces, rather than into the cutting tool [Cselle and Barimani, 1995 and Paskvale, 2007]. This
conjecture is proved to be false for the crater wear [Olortegui-Yume and Kwon, 2010]. This
paper explores if this conjecture works for flank wear.
This chapter studies the flank wear of multi-layered coating using the various microscope
techniques, 2D wavelet filtering and FE simulation. The flank wear rate and wear evolution are
analyzed and the relationship between the roughness on the tool surface and flank wear are
investigated. The abrasion model uses hot hardness data based on the temperature profiles from
FEM simulation, which allows us to predict flank wear patterns as a function of cutting time. In
addition, the flank wear predicted with the 2-body abrasion model are compared with the
experimental data for the validation.

19

2.1 FLANK WEAR MODEL AND IMAGE PROCESSING TECHNIQUE
2.1.1 Abrasive wear model

The dominant wear phenomenon on the flank surface is abrasion by hard inclusion,
cementite, which results in groove on the tool surface. In the pearlitic microstructure, 2-body
abrasion occurs because a complex morphology of the second phase is constrained within a
matrix of ferrite. In this work, The 2-body abrasive wear models was used (see Chap. 1.2.1) For
the abrasive model, hot hardness data for coating materials and abrasives, which are shown in
Figure 1, were used [Wong et al., 2004]. In addition, wear constants were determined by
conducting calibration based on cutting experiments in conjunction of FEM temperature profiles.
For validation, the predicted wear evolution was compared with experimental measurements.

2.1.2 Finite element model

FE metal cutting simulation was carried out using ABAQUS/Explicit version 6.6 to
determine the interfacial temperature profiles between flank surface and work material at the
various flank wear land (VB) (from 70μm to 300μm). The flank wear land (VB) was defined as
in Figure 5 (a) in the FE model used. The Johnson-Cook (JC) constitutive model was adapted to
model the work material and ALE adapted meshing technique was used to avoid severe mesh
distortion. In addition, extremely fine meshes were applied around the interface between cutting
tool and work piece. Four node bilinear displacement and temperature quadrilateral elements
(CPE4RT), which has been used in other metal cutting simulation studies [Ozel and Zeren, 2005,
Nasr et al., 2007 and Madhavan and Adibi-Sedeh, 2005], were used. A single composite coating

20

on top of the carbide has the ‘equivalent’ conductivity by combining thermal conductivity of
individual coatings (TiN/Al2O3/TiCN) [Yen et al., 2004]. The cutting tool with the composite
coating layer was assumed as a mechanically rigid and thermally conductive body. Boundary
conditions and initial temperature are shown in Figure 5 (b). The coefficients for JC constitutive
model and material properties of tool and work material are summarized in Tables 1 and 2,
respectively [Olortegui-Yume et al., 2008]. The simulation results, mainly interfacial
temperature, were fed to the abrasive wear prediction models.

(a) Flank wear land (VB) in FEM model
Figure 5 Flank wear land (VB) and boundary conditions for FE simulation

21

Figure 5 (cont’d)

(b) Boundary conditions

Table 1 Material constants for the JC model
A (MPa)

B (MPa)

C

n

M

Tm

553.1

600.8

0.0134

0.234

1

1460

22

Table 2 Material properties of tool and work material
AISI 1045
212 (20°C)
207 (100°C)
Young’s modulus (GPa)
192 (300°C)
175 (500°C)
Poisson’s ratio

WC-Co
612

0.3
7844
50.7 (100°C)
41.9 (400°C)
30.1 (700°C)
26.8 (1000°C)
486 (75°C)
548 (275°C)
649 (475°C)
770 (675°C)
548 (875°C)
11.9 (20°C)
12.5 (100°C)
13.6 (300°C)
14.5 (500°C)
15.1 (700°C)

3

Density (kg/m )
Conductivity (W/m°C)

Specific heat (J/kg°C)

Thermal Expansion (μm/m°C)

0.22
11900
86

334.01
+0.12T(°C)

4.9 (200°C)
5.2 (400°C)
6.5 (600°C)

Coating materials
TiN
Young’s modulus, E (GPa)

250

Poisson’s ratio, ν

0.25

Density , ρ (kg/m )

3

4650

Thermal Expansion, α (μm/m°C)

TiN/Al2O3/TiCN

9.4
22/26/27 (25°C)
23/18/30.5(200°C)
26/8/34 (727°C)
26.5/4/35 (927°C)
27/6/36 (1200°C)
25.9 (25°C)
22.21(200°C)
12.96 (727°C)
7.25 (927°C)
10.33 (1200°C)

Conductivity, K (W/m°C)

‘Equivalent’ Conductivity, (W/m°C)

23

2.1.3 Image acquisition and wavelet transform

A JEOL JSM-6400 SEM, which includes both secondary electrons (SE) and
backscattered electrons (BSE), and a CLSM (Zeiss LSM210) were used for the flank wear
analysis at cutting times. The flank wear images taken by SEM can be mainly used for the
measurement of flank wear land (VB) and the observation of the wear evolution (behavior) for
the coating materials. On the other hand, CLSM can provide the height information (3D surface
profile) of the tool surface with high axial resolution (50nm) by an axial optical slicing process
throughout the surface from top to bottom of the surface height [Yuan et al., 2005 and OlorteguiYume and Kwon, 2010]. For each optical section, the maximum brightness, which is associated
with the surface profile, is detected at a focal plane and stacked up to total number of the sections
so that the HE image can be built up with a series of these optical sections. In this study, for the
3D topography of the flank surface, HE images, corresponding to each maximum bright (MB)
images, were obtained with an objective of 50X as shown in Figure 6 (a) and (b). The 200 optical
sections with a step size of 200nm were used for high accurate surface profile.
In present work, the HE images were obtained from CLSM and wavelet denoising
technique was used to eliminate the noise and artifacts in HE raw image (See Figure 6 (c)).
Finally, the wavelet-filtered wear profiles (See Figure 6 (d)) were obtained for analysis of the
flank wear mechanism. In addition, roughness analysis on the tool surface was performed. The
mother wavelet, Daubechies20 (‘db20’), was chosen for 2D DWT from the reference [OlorteguiYume and Kwon, 2010].

24

(c)

(a)

Z-Matrix raw

HE

y( m)

(d)

(b)

x( m)

Z-Matrix at level 2

x( m)

MB

y( m)

Figure 6 (a) HE and (b) MB images from CLSM at 7min cutting time and (c) raw zmatrix and (d) wavelet filtered images from HE image

2.2 EXPERIMENTAL SETUP

To study the wear mechanisms on the flank surface, turning tests with refined AISI 1045
steel were performed at the cutting time of 1, 2, 3, 4, 7, 10, 12, 20 and 23 min, where each test
was starting from a fresh insert corner. A multilayer coating (SNMA 190612 - ISO) on C6
cemented carbide for substrate (KC9315 grade) was used. For multilayer coating, TiCN coating
layer was deposited on the carbide substrate followed by Al2O3 and TiN layers. The thicknesses

25

of each coating are shown in Table 3. The cutting conditions consist of cutting speed, feed rate
and depth of cut (DOC), which are 250m/min, 0.3175mm/rev and 1.905mm, respectively.

Table 3 Thickness of the coatings
Coating Materials

TiN

Al2O3

TiCN

Thickness

1.940μm

9.160μm

7.584μm

2.3 RESULTS AND DISCUSSION
Figure 7 illustrates the SEM images of flank wear land with the 60X objective at each
cutting time. The flank wear land (VB) was uniformly distributed throughout the cutting edge up
to the cutting time of 12 min. However, the wear increased substantially at the notch region at 20
min and 23 min. On the other hand, BSE was obtained, as shown in Figure 8 to characterize the
wear evolution on the coatings at cutting times. Wear took place mostly on the alumina (Al2O3)
coating, which is shown black in the BSE images, from 1 min to 3 min and then titanium carbon
nitride (TiCN) coating started to be exposed at 4 min. The carbide substrate was exposed after 12
min and the carbide layer was entirely covered by the steel (see 20 and 23 min in Figure 8).

26

1 min

2 min

3 min

4 min

7 min

10 min
Notch wear

12 min

20 min

23 min

Figure 7 Evolution of flank wear observed by SEM

1 min

3 min

2 min
TiCN
Al2O3

TiN

4 min

7 min

10 min

WC
Steel

12 min

20 min

23 min

Figure 8 BSE images on flank surface
27

2.3.1 Roughness analysis
Figure 9 (a) shows the confocal image obtained from CLSM while Figure 9 (b) shows the
approximation (A) and details (H, D and V) at the level 3 of decomposition, corresponding of
Figure 9 (a), after the wavelet decomposition and reconstruction process. The decomposition
level has to be determined to obtain more accurate data for surface analysis. Thus, by comparing
the approximations at various levels, the maximum decomposition level was selected to be A6
since the ‘approximation’ surface at the level 7, A7, has been significantly distorted and the
directionality of the scoring mark has been lost for details (H, V and D).
The details (H, D and V) obtained from the 2D DTW can be used to identify the small
features on the flank surface associated with surface roughness, waviness as shown in Figure 10.
The roughness can be regarded as high frequency components of tool surface which can be
correlated with the scoring marks, while the waviness from the medium frequency, which has
greater spacing than the roughness, represents irregularities from a mean line. Waviness may
come from vibrations of machine tool. The form can be considered to be the low frequency
components [Raja et al., 2002 and Whitehouse, 2002]. These features can be obtained by adding
wavelet-decoupled details obtained at different levels of the decompositions up to the criterion
(level 6) determined above [Olortegui-Yume and Kwon, 2010, and Jiang et al., 1999]. The
roughness and waviness are defined as in Eqs. (7) and (8), respectively.

28

(a) CLSM Image

(b) Approximation (A) and Details (H, D and V)
Figure 9 CLSM image for the flank surface at 4 min (a) and wavelet decoupled
components (b) at Level 3

29

Roughness

HL

VL

DL

L

(7)

Waviness

HL

VL

DL

L

(8)

The roughness image obtained by adding the details (H, D and V) up to the level 6 shows
the scoring marks on the flank wear with the directionality to the abrasion path on the wear land
as seen in Figure 10 (b). For the surface roughness statistical analysis, the root mean square (Rq)
of the profiles was used rather than average surface roughness (Ra) because with Ra detailed
information is lost [Gadelmawla et al., 2002].

(a) Surface features on flank surface
Figure 10 Roughness, Waviness and Form
30

Figure 10 (cont’d)

(b) Roughness

On the tool after cutting 1 min, surface roughness (Rq) as a function of distance from the
edge was measured on Al2O3 coating at three locations indicated by three short lines in Figure11
(a). The green line on Figure 11 (a) is closest from the trailing edge or furthest away from the
cutting edge. As seen in Figure 11 (b), the roughness deviation has been reduced as we assumed
that larger cementite phases may be broken down or embedded into the matrix of softer ferrite.
Thus, the abrasion damage becomes more uniform along the flank wear land. Furthermore, to
observe the roughness change at cutting times on Al2O3 coating at the trailing edge, the
roughness on Al2O3 coating was measured at cutting times up to 12 min as summarized in Table
4. The roughness value with the cutting time of 12, 20 and 23min was disregarded since the
roughness data were distorted due to the adhesion on the tool surface. The items on the first
column in table 4 represent each event during the wear evolution as each layer of the multi-layer
31

coating is exposed. Initially, the roughness increases with cutting time up to 3min at which only
Al2O3 coating was exposed but after TiCN coating was exposed (between 4 and 10 min), the
roughness at the trailing edge (on the Al2O3 coating) slightly decreased with cutting time. This
might be because the abrasive, cementite, in the work material may be ‘driven’ into the softer
matrix by the hard TiCN layer, while traversing across the flank land.
Figure 12 (a) and (b) present the roughness and groove size of the flank surface at the
cutting time of 7 min. The tool after cutting 7 min was chosen for this observation, as the steel
was not adhered to the tool. As was done in Figure 11, the locations, where the roughness and
groove size were measured, are presented with the short colored lines in Figure 12 (a). The
measurements at four locations are presented in Figure 12 (b) with the corresponding coating
materials. The Al2O3 coating being the softest had the largest roughness and groove size
demonstrating the consequence of abrasive wear while the TiCN coating being the hardest had
the least roughness and groove size. This trend became more obvious at 23 min as seen in Figure
13 (b). Thus, it was confirmed that the hardest coating TiCN among the multi-layer coatings has
the most significant role against abrasion wear while alumina coating has a great resistance
against thermal shock and chemical wear [Olortegui-Yume et al., 2008].

32

(a) Locations measured on Al2O3 coating

(b) Roughness at the locations in (a)
Figure 11 Roughness Data on Tool after Cutting for 1 min
33

Table 4 Roughness data on the coatings
Time (min)

0.6717

0.7225

2

0.7707

0.6709

3

0.9100

1.7413

4

0.7814

1.0473

7

0.5299

0.6533

10

The exposure of TiCN

TiN

1
The exposure of Al2O3

Al2O3

0.4777

0.5480

(a) Locations of roughness measured
Figure 12 Roughness data and Groove sizes on the tool after cutting of 7 min

34

Figure 12 (cont’d)

(b) Groove sizes at the locations in (a)

50
100
150
200
250
50

100

150

200

250

x(pixels)

(a) Locations measured
Figure 13 Roughness on the tool after cutting for 23 min after etching of steel
35

Figure 13 (cont’d)

1.9um

2.1um

3um

3.9um

3.1um

(b) Roughness at the locations in (a)

Figure 14 (a) shows the microstructure of the work material used, refined AISI 1045 steel.
The groove size should be related to the size of the cementite due to the fact that abrasion wear
takes place by the hard cementite phase. However, the size of each cementite in pearlite colony is
too small (~0.5μm) to be matched with the size of scoring mark (3~10μm) observed.
Nonetheless, if the angle of the cementite orientation in pearlitic phase to the tool surface and the
size of pearlite conlony are considered, the groove size can be identified. For example, if the
angle is 90º as in Figure 14 (b), very small groove size should be detected. However, the hard
cementite in pearlitic phase can attack in any orientation, which can attribute to the observed
36

groove size. On the other hand, the size of the pearlite phase shown in Figure 14 (a) is between
20μm and 35μm. Despite of the fact that a larger size of pearlite phase observed compared to the
groove size, it should be considered that only small segments of pearlite colony might be
involved in abrading action as seen in Figure 14 (b). In addition, the groove size can be also
dependent upon the penetration depth, D. This is why groove depth observed on TiCN was
smaller than other coatings as seen in Figure 12 (b) and Figure 13 (b). Another possible scinerio
is that pearlite can be broken down to smaller pieces due to the severe shear deformation in the
cutting process so that only smaller fractured pearlites generate the scoring mark. In both cases,
the abrasive wear takes place by the pearlite grains rather than individual cementite lamella.

(a) Microstructure of refined AISI 1045 steel
Figure 14 Microstructure (a) and a pearlitic colony size (b) of refined AISI 1045
steel
37

Figure 14 (cont’d)

(b) Groove size by a pearlite colony

2.3.2 Flank wear analysis

To study the flank wear evolution, the surface profiles on the flank were extracted from
the HE images obtained by CLSM using wavelet-denoising and 2D DWT process. The
decomposition level 2 was selected for 2D DWT since the wavelet-filtered profile
(approximation) was too smoothed after the level of 2. Figure 15 shows both 3D and 2D crosssectional profiles of the flank wear for the raw image and the wavelet filtered image (A2) at the
decomposition level 2. The cross-sectional profiles at the cutting times were collected and finally
merged on to the mask of the coating materials in order to analyze the flank wear evolution as
presented in Figure 16. Up to 3 min., TiN and Al2O3 coatings were worn down quickly and no
TiCN coating layer was exposed. However, after cutting 4 min., TiCN layer was exposed and
38

stayed exposed until the cutting time of 12 min. It indicates that TiCN did not wear at such high
rates between 4 min and 12 min. due to its high hot hardness. Subsequently, the carbide was
substantially worn down at 23 min due to the fact that the carbide substrate has a lower wear
resistance associated with the lower hot hardness compared to the coating materials. This is
easily depicted in Figure 15. Interestingly, the flank wear rate slightly decreased from 12 min to
23 min while the tool got close to the tool failure as in Figure 16 (b). Actually, the cutting
experiments had to be repeated few times due to the catastrophic fracture of the cutting tool right
before the cutting time reached 23 min. However, as shown in Figure 16 (c), the volume wear
rate, indeed, significantly increased after the cutting time of 12 min. It means that once the
carbide is exposed, the wear front will attack mostly the softer carbide without extending the
flank wear land. Thus, the standard flank wear, VB, does not adequately describe the state of tool
wear and the wear volume and profiles may be considered for the flank wear analysis.
The accelerated wear rate of the carbide comes also from the adhesion. Figure 17 (a)
shows the steel adhesion on the carbide at cutting time of 23 min. Once the carbide substrate was
exposed, the steel adhesion can take place due to the excellent affinity between steel and carbide.
The adhered steel was pulled out the carbide grain when the steel detached from the carbide
substrate. To identify the amount of adhesion, the cutting insert was etched ultrasonically with
hydrochloric acid for 20 min. and compared the surface profiles before and after etching at 23
min. Finally, it was found that the adhesion thickness was around 20μm as shown in Figure 16
(a). In addition, the deterioration of TiCN coating due to the accelerated wear of carbide
substrate was observed in Figure 17 (b). On the other hand, it is commonly believed that
multilayer coating can enhance the resistance of coatings against crack propagation [Cselle and

39

Barimani, 1995 and Paskvale, 2007]. However, the argument does not agree with our
experimental data as shown in Figure 16 (a).

Figure 15 3D and cross-sectional profiles of raw data and wavelet-filtered at level
2

40

(a) 2D Flank wear profiles at cutting times

(b) Flank (VB) wear rates
Figure 16 Flank wear profiles and wear rates
41

Figure 16 (cont’d)

(c) Volume wear rates

(a) Steel adhesion on the carbide
Figure 17 Adhesion on the carbide and coating delamination at cutting time of 23
min

42

Figure 17 (cont’d)

(b) Coating delamination

Figure 18 presents the notch wear on the tool at depth of cut line (DCL) at cutting time of
20 min. and 23 min. The chip material suffers the severe shear deformations and the lateral
extension of the chip at DCL was reported [Chandrasekaran and Johansson, 1994]. The possible
notch wear mechanism was proposed as the chip at the DCL region with relatively high micro
hardness can plough or adhere to the tool surface depending upon the affinity between the
materials. In our cutting test, notch wear become significant between 20 and 23 min. at which
carbide was exposed. Thus, the ‘extended chip’ abrades the tool surface until carbide was
exposed and the adhesion wear takes over as soon as the carbide was exposed. Note that the
notch wear was not detectable up to 12 min. due to the high wear resistance of the ceramic
coatings against the abrasion. The intermittent and repeating process of the adhesion/abrasion
between carbide and work material produces the severe notch wear, which eventually contributes
to the catastrophic tool failure around 23 min. Notch wear can be observed more clearly after
43

etching away the steel as seen in Figure 19 (a) and the size of scoring mark on the alumina
coating in notch wear zone is much bigger than that in the uniform flank wear zone as in Figure
19 (b). This is because the blunt edge at notch region cannot cut the pearlite colony sharply so
the size of pearlite becomes bigger and pearlite surface gets rougher so that the roughness on the
coating became much larger at notch wear zone.

Carbide

Steel
TiCN
Al2O3
TiN
(a) At 20 min

(b) At 23 min
Figure 18 Notch wear on the tools after 20 and 23 min of cutting

44

200μm
(a) After etching of Figure 18 (b)

(b) Scoring marks on alumina coating at 23 min in (a)
Figure 19 Notch wear after etching (a) and scoring marks and notch wear zone (b)
The temperature distributions from FEM simulations at the flank surface were obtained
as a function of the flank wear land as shown in Figure 20. The temperature change due to the
flank wear shows the continuously decreasing trend with the flank wear land up to 70μm.
However, the temperature profiles start to have parabolic shapes after the flank wear land (VB)
45

reaches 120μm. At that point, the maximum temperature occurs at the trailing edge of the flank
wear land. Until then, the temperature decreases along the flank surface. However, as the flank
wear land increases further, the temperature increases again near the trailing edge and reaches
the local maximum. At the same time, the result of the FEM simulation shows that the cutting
temperature also slightly decreases as the flank wear land (VB) increases as seen in Figure 20.
This can be explained by the 2-body abrasion model in which the wear rate increases with
temperature. In addition, the phase transformation of the work material due to the high interface
temperature and pressure, may be affecting the wear evolution [Kwon, 2000]. However, the
volume wear rate increases once the carbide is exposed after 12min in Figure 16 (a). This is
possibly because adhesive wear might be gaining its momentum over abrasive wear.

Figure 20 Temperature profiles on cutting tools from FEM simulations

46

From the experiment measurements and FEM simulation results, the wear calibration
constants were determined and temperature profiles on the cutting surface were obtained
respectively. The temperature profiles and the wear constants were used for the 2-body abrasive
prediction model presented in Eq. (1) programmed by MATLAB software. The predicted flank
wear profiles are shown in Figure 21. Comparing to Figure 16 (a), the 2-body abrasive model
with hot hardness data of the coating materials predicted quite well, as the assumption that 2body abrasion wear takes place when the pearlitic steel is used because of the constrained
microstructure of cementite phase in the form of peralite [Wong et al., 2004]. Again, based on
the predicted results, the wear progress on TiCN coating, that is between 7 and 12 min, was so
small due to its high hot hardness, which agrees well with the experiment results.

Figure 21 Predicted flank wear based on 2-body wear model

47

2.4 CONCLUSION

The following conclusions can be drawn:
The combination of CLSM and 2D wavelet transformation techniques is useful in
obtaining the flank wear profiles and the surface features such as groove size and
roughness on the tool, which can be used to study the cause and evolution of flank wear.
Finite element simulation is able to provide the interfacial cutting temperature, which are
needed to predict the flank wear profiles. The interfacial pressure was assumed constant
as the FE simulation cannot maintain the intimate contact.
TiCN coating shows high wear resistance against the abrasion while the carbide has the
least resistance in wear. After carbide was exposed, the work material adhered to the
carbide substrate and the cutting tools were fractured due to the crater degradation on the
carbide surface. Because of the larger grains in carbide, the wear mechanism can be
quite different as the carbide grains can pull out.
Notch wear becomes significant at 20 and 23 min by the intermittent repeated process of
adhesion/abrasion. This may come from the stick-slip interfacial condition. The size of
the scoring mark on the Al2O3 coating at notch wear zone is much bigger than that at
uniform flank wear zone.
The groove size was measured and showed that the softer coating is damaged much
more. As soon as the hard TiCN coating is exposed, the abrasive cementite cannot
affectively abrade the hard surface.

48

The correlation between groove size and pearlite grain size was investigated. Only a
small segment of the pearlite colony and/or fractured pearlite may abrade the tool
surface producing the scoring marks.
Unlikely for the crater wear in which the complex wear mechanism is involved, for the
flank wear the multilayer coating did not show much beneficial effect and only hard
coating that has high hot hardness is required.
2-body abrasive model in conjunction with FEM simulation results shows very good
prediction of flank wear.

49

Chapter 3
TOOL WEAR ANALYSIS ON MULTI-LAYERED COATED CARBIDE TOOLS IN
FACE MILLING

Recent coating technology has advanced to extend the life of carbide tools effectively in
metal cutting process. In general, the coating can be deposited onto the carbide tools in a singlelayer or multi-layer scheme processed by CVD and/or PVD methods. The coating materials
along with the coating process must be determined depending on the machining process and the
work material. Each coating process has its advantages and disadvantages. For example, CVDAl2O3 coating provides a thermal barrier and good bonding with other coatings. However, the
brittle eta phase can be formed at the coating interface at high deposition temperature, which
reduces the transverse rupture strength (TRS) of the substrate and weakens the cutting edge
[Stephenson and Agapiou, 2006]. On the other hand, the PVD coating process produces the sharp
cutting edge and free thermal cracks on the cutting surface [Sodergberg et al., 2001]. The
performances of PVD and CVD coated carbide tools were investigated in many recent works
[Jawaid et al., 2000 and Dobrzanski et al., 2006]. Jawaid et al., [2000] found that the CVD
coating tool was superior to the PVD coating tool due to high adhesion strength of the CVD
coating tool despite of inherent disadvantage from the CVD coating process. Dobranski et al.,
[2006] compared CVD and PVD in nitride tool ceramics on a ‘pin-on-disk’ method and
concluded that the multilayer coating, processed by CVD, was better due to the adhesion
strength.
Several tool wear mechanisms have been reported in metal cutting processes [Stephenson
and Agapiou, 2006, Melo et al., 2006, Olortegui-Yume and Kwon, 2007(1) and Gu et al., 1999].
Abrasive wear dominates the wear process when the hard inclusions in work material abrade the

50

tool surface. Thus, the abrasive wear strongly depends on the hardness of tool materials. In
addition, adhesive wear can occur when a high affinity of work material to a particular tool
material exists. At low cutting speeds, build up edge (BUE) can be formed and detached. During
the detachment of BUE, small volumes of tool material can be separated away. When the cutting
temperature is high as in continuous turning, diffusion or dissolution wear is dominant at the rake
surface [Olortegui-Yume and Kwon, 2007(1)]. In milling, in addition to abrasive wear, thermomechanical fatigue wear and crack can take place mainly due to the nature of the interrupted
process, which generates the thermo-mechanical cycles.
In general, the multi-layer coated tools are known to deter tool wear of carbide tools
[Olortegui-Yume and Kwon, 2007(2), Olortegui-Yume et al., 2008, Cselle and Barimani, 1995].
However the reason for the wear deterrence is not well established. In the case of turning AISI
1045 steel with multilayer (TiN/Al2O3/TiCN) coated carbide tools, the reason for extending tool
life becomes clear recently [Olortegui-Yume and Kwon, 2007(2), Olortegui-Yume et al., 2008].
At the rake, crater wear occurs in combination of abrasive and dissolution wear mechanisms. At
high cutting speed, the continuous turning process increases the interfacial temperature high
enough to activate dissolution wear. At the flank, abrasive wear is the main dominant wear
mechanism. The success of the multilayer coated tools comes from the fact that the Al2O3
coating shows outstanding resistant to dissolution wear which is the important wear mechanism
for crater wear especially at high cutting temperature. On the other hand, hard coating, TiCN,
exhibits excellent performance against the abrasive wear while TiN coating provides low friction
at the cutting interface. Therefore, the multilayer scheme works because the tribological
condition at the interface changes as each coating exposes during machining. With the change in
the tribological condition, the location of maximum cutting temperature changes, which in turn
51

changes the location of maximum crater depth. Thus, the multilayer scheme extends the tool life
by changing the maximum wear locations as each layer is exposed [Olortegui-Yume et al., 2008].
On the contrary, it is commonly believed that the effectiveness of multilayer coating
comes from the delay in the crack propagation [Sadik and Myrtveit, 2009] because the crack can
propagate along the coating interfaces instead of right through the coatings into the substrate.
Under the cyclic loading in milling process, enhanced toughness and wear resistance against
thermal and mechanical impact are required; and, again, multilayer coating has been cited to
improve the life of cutting tools [Cselle and Barimani, 1995].
This chapter reexamined these issues by analyzing the tool wear evolution of multi-layer
coated inserts. Wear surface profiles was obtained using the combined technique of confocal
laser scanning microscopy (CLSM) and wavelet transform as presented in [Olortegui-Yume and
Kwon, 2010]. The 3D surface profiles as a function of cutting times were obtained to observe the
flank wear evolution of the multi-layer coated inserts after face milling. Two multi-layered
coated inserts processed by CVD and PVD respectively were evaluated in terms of flank wear
(VB) and edge chipping as well as adhesion test.

52

3.1 EXPERIMENTAL SETUP

Face milling tests were performed without any lubrication on an AISI 1045 steel block
(8”X5”X6”) with the coated carbide inserts (SEEW 42 A7A) provided by Valenite Inc. (Madison
Heights, Michigan). The square type inserts were coated by two different types of coating
processes, CVD for triple layer (TiN/ Al2O3/TiCN) and PVD for double layer (TiN/TiAlN). Only
one insert was used in a three-flute tool holder for the experiment to reduce variables in tool
wear. Figure 22 shows the face milling setup. The face milling started at one corner of the steel
and completed one path. After cutting 5 paths, one layer was completely removed. The cutting
th

was performed up to the 8 layer unless edge chipping occurred. The flank wear was observed
with CLSM after each layer has been removed. The cutting conditions are summarized in Table
5.
An indentation test was also performed using a Wilson Rockwell Hardness Tester in order
to evaluate the adhesion quality of the double and triple layer coatings deposited on to the
carbide. Maximum applied force and maximum radius of indentation were set at 1472N and
0.2mm respectively.

Figure 22 Schematic of the face milling
53

Table 5 Cutting conditions
Cutting Conditions
Work material

AISI 1045

Cutting speed (m/min)

200, 300 & 400

Feed rate (mm/min)

0.15

Axial DOC, ADOC, (mm)

0.5, 1 & 2

Radial DOC, RDOC, (mm)

26.1

3.2 TOOL WEAR MEASUREMENT

For CLSM, a Zeiss LSM210 was used for flank wear width (VB) measurement and 3D
flank wear profiles. Additionally, a scanning electron microscope (SEM) (JEOL JSM-6400) for
backscattered electron image (BSE) and an energy dispersive X-ray Spectroscopy image (EDS)
was also utilized for the observation of the extent of the adhesion. Figure 23 illustrates the
CLSM images of flank wear land (VB) with an objective of 20X for measurement. In addition, to
characterize the flank wear evolution on the insert after removing each layer, height encoded
(HE) images were obtained from the CLSM with an objective of 50X as shown in Figure 24.
Ultimately, these HE images were converted into 3D surface profiles using an image processing
technique including wavelet transform, which are represented in Figure 25. For example, once
the HE image in Figure 24 (a) is obtained by the CLSM, wavelet de-noising with the mother
wavelet, Daubechies20 (‘db20’), is performed using Matlab software to eliminate the noise and
spikes in a HE raw image. And then the 2D wavelet transform is applied in order to obtain the
tool wear profiles by decoupling the small features on the surface such as roughness. The 2Dwavelet analysis decouples the raw 3D data to 4 sets of wavelets coefficients, called
approximation, horizontal details, diagonal details and vertical details at levels. The 3D data can
54

be transformed up to the level where the surface profiles maintain the original shape without
distortion. Then inverse wavelet transform for these coefficients are required to obtain an
approximation, horizontal, diagonal and vertical details. The approximation can be used as the
tool wear profile while the details can be considered as surface roughness. The decomposition
level has to be determined to obtain more accurate data for surface analysis. Thus, by comparing
the approximations at various levels, the maximum decomposition level can be selected. In this
study, the approximations at zero level are used to obtain the tool wear profiles at cutting
distances (layers). The detailed information for the image processing technique and wavelet
transform can be found in [Olortegui-Yume and Kwon, 2010 and Park and Kwon, 2009]. On the
other hand, EDS images of the worn tool were obtained to map the elements on the tool and
observe the wear progress of coating, as shown in Figure 26.

VB
VB

(a) After 5

th

th

layer cutting

(b) After 8 layer cutting

Figure 23 CLSM images for flank wear land (VB)

55

th

Figure 24 HE image of flank wear after 5

layer cutting

th

Figure 25 3D surface profile from HE image after 5

th

(a) After 5

layer cutting

Figure 26 EDS images of flank wear
56

layer cutting

Figure 26 (cont’d)

th

(b) After 8

layer cutting

3.3 FLANK WEAR ANALYSIS

The flank wear land (VB) was uniformly distributed throughout the cutting edge until the
chipping occurs at the cutting edge. And no significant crater wear was observed regardless of
the cutting conditions. As shown in Figure 26, steel adhesion was observed on carbide substrate.
Figure 27 represents the flank wear land (VB) for triple layer coating at various cutting
conditions. It was observed that the higher cutting speed caused the more flank wear coating
carbide, as expected.
The effect of depth of cut (DOC) was minimal for the cutting speeds of 200m/min and
rd

300m/min. The considerable difference of the flank wear land (VB) for the 3 layer cutting at
300m/min was observed but that is mainly because catastrophic edge chipping occurred. It was
also found that at high cutting speed, the edge chipping was the main tool failure mode.
57

st

Especially, at the cutting speed of 400m/min, the tool was fractured only after the 1 layer
cutting. While the abrasive wear was a dominant tool wear mechanism, the adhesion was also
observed after the carbide layer was exposed. Once the carbide substrate has been exposed, the
wear rate became significant due to the relatively low hardness of the carbide and the edge
chipping during the removal of the adhesion layer. For comparison the flank wear of double layer
coating was also measured at DOC of 0.5mm at various cutting speeds as seen in Figure 28.
Similar to the case of the triple layer coating, the higher cutting speed reduced the tool life of the
double layered tools.

Figure 27 Flank wear (VB) for triple layer coating at various cutting conditions

58

Figure 27 (cont’d)

Figure 28 Flank wear (VB) for double layer coating at various cutting speeds

Figure 29 represents the flank wear land (VB) for both double and triple layer coating at
cutting speeds where the solid lines and dot lines indicate the double and triple coatings
respectively. It shows that the double layer coating outperformed the triple layer coating
regardless of the cutting speeds. Especially at 200m/min, the flank wear rate for the double layer
is much smaller than that for the triple layer. The chipping occurred at the entire cutting edge at

59

the triple layer coating while no significant edge chipping was observed for the double layer
coating. To explain the reason for the double layer tool outperforming the triple layer tool
observed in the experiment, the hardness of the coatings has to be taken into account first due to
the fact that the dominant wear mechanism at flank is the abrasion. In this work, the hot hardness
of the CVD and PVD coatings could not be directly measured. However, it was reported that
TiCN coating is slightly harder than TiAlN coating below 800ºC [Cung-Chen and Hong, 2002]
and the hardness of PVD TiCN coatings was greater than that of the CVD TiCN coatings [15]. In
this regard, it can be inferred that the hardness of the double layer (TiN/TiAlN-PVD) coating
could be similar or slightly higher than the triple layer (TiN/Al2O3/TiCN-CVD) coating.
Another factor to be considered is the quality of the coating adhesion to the carbide. From
the indentation test, it turned out that both coatings (Figure 30 (a) and (b)) showed acceptable
adhesion strength quality between HF3 and HF4 as defined in Figure 30 (c) where HF1-HF2 and
HF3-HF4 were defined as “preferred” and “acceptable” adhesion quality respectively; whereas
HF5-HF6 were classified as “not acceptable” or “insufficient” adhesion [Heinke et al., 1995].
From the analysis of the hardness and adhesion, it cannot be explained conclusively that double
(PVD) coating was superior to the triple (CVD) coating.

60

Figure 29 Flank wear comparison of double (PVD) and triple (CVD) layer coating

(a) Indentation test for the double CVD coating

(b) Indentation test for the triple PVD coating
Figure 30 Indentation tests of PVD and CVD coatings

61

Figure 30 (cont’d)

(c) Adhesion quality of coating [Heinke et al., 1995]
However, the inferior performance of the triple layer (CVD) coating can be explained by
the reduction of the TRS by formation of the brittle eta phase at the cutting edge by the CVD
coating process [Stephenson and Agapiou, 2006 and Jawaid et al., 2000]. In addition, the CVD
coating process causes the tensile residual stresses at the cutting surface generated by high
deposition temperature in the coating process. As in Table 6, the coefficients of thermal
expansion (CTE) of the coating materials are higher than that of the carbide so the CTE
mismatch causes the tensile stress at the cutting surface when the tool cools down from a high
deposition temperature around 1000ºC. Even the deposition temperature of a medium
temperature (MT) CVD method is around 800ºC. Thus, the tensile stress is inherently higher
with the CVD process. In addition to the additional tensile stresses associated with the CVD
process, the thermo-mechanical loading cycle from the interrupted process can generate and
62

accelerate the micro-cracks on the cutting tool, which results in the edge chipping and tool
failure. On the other hand, the deposition temperature of the PVD process is around 500ºC,
which is substantially low compared to that of CVD. Consequently, the PVD coating generally
provides free of the thermal crack. In this regard, the PVD coating can be superior to the CVD
coating in milling process.
Table 6 CTE of the tool materials [Wang et al., 1999 and Zhao et al., 2003]
-6

Coating Materials

CTE (10 /K)

TiN

9.4

Al2O3

7.8-8.4

TiCN

8.4

TiAlN

7.5

WC/Co

5.0

To investigate the wear pattern of the cutting tool, 2D cross-sectional profiles were
obtained at cutting distances. Figure 31 shows the flank wear evolutions of the double layer
coating at the cutting speed of 200m/min, 300m/min and 400m/min. As seen in Figure 31 (a), the
slopes of the profiles on the TiAlN coating are lower than that on the carbide. This means that
the high hardness of TiAlN coating reduced tool wear. Again, in Figure 31 (a) the cross sectional
rd

profiles up to 3 layer cutting, green line, still stayed on the TiAlN coating but once the carbide
was exposed, the wear rate was substantially increased. In addition, the adhesion on the carbide
was also observed which might accelerate the edge chipping due to the thermo-mechanical
impact during the cutting. This trend becomes more obvious at 300m/min and 400m/min as
th

shown in Figure 31 (b) and (c) respectively. The tool was fractured after the 4 layer cutting at
400m/min.

63

(a) 200m/min

(b) 300m/min
Figure 31 Flank wear evolutions of double layer coating

64

Figure 31 (cont’d)

(c) 400m/min

Figure 32 shows flank wear evolution of the triple layer at cutting distances at 300m/min
for comparison with the double layer coating. The triple layer coated tools showed the same
trend as the double layer coating that flank wear slightly increased up to the first 2 layers cutting
rd

due to the hard TiCN coating. However, after the 3 layer cutting, the flank wear substantially
increased and the edge chipping was observed. However, the flank wear was much more severe
for the triple layer, not only the flank wear land (VB) but also the wear volume. Even the tool
rd

was fractured only after the 3 layer cutting for the triple layer coating (See Figure 31 (b) and
Figure 32).

65

Figure 32 Flank wear evolution of triple layer coating at 300m/min

From the observation of the results, the abrasive wear is the dominant wear mechanism at
the flank. Thus, the hardness of the coating is of primary importance for flank wear. In fact, the
Al2O3 coating layer with the high dissolution resistance is not needed. In this regard, the harder
coating is more preferred to improve the wear resistance of the tool regardless of a coating layer
scheme, a single-layer or multi-layer. This argument agrees with other work that the multi-layer
coating did not provide better performance in [Sadik and Myrtveit, 2009]. However, in some
work, the multilayer can improve the hardness and adhesion strength of coating so that it can
perform better than a single-layer coating [Chen et al., 2008]. On top of that, a multi-layer
coating can also provide increased thickness of coating, which can reduce the propensity for
crack propagation [Cselle and Barimani, 1995]. On the other hand, it is commonly believed that
66

the superior performance of the multilayer coating comes from the crack advancement
deteriorated by propagating along the coating interface. However, as shown in Figs. 10 and 11,
no cross-sectional profiles show that the crack propagates along the coating interface. Thus, it
can be concluded that a multilayer can improve the coating strength and hardness [Paldey and
Deevi, 2003] so that the wear rate can be reduced when the abrasive wear takes place. However,
it was observed that the benefit of multilayered cutting tools did not come from the wear
resistance enhanced by promoting the interfacial cracks between the coatings. The only benefit
may come from the improvement in the coating strength and hardness [Paldey and Deevi, 2003].

67

3.4 CONCLUSION

The following conclusions can be drawn from the present chapter:

• Double layer (PVD) coating outperformed triple layer coating (CVD) in face milling of
AISI 1045 steel due to inherent advantages of the PVD coating deposition process.
• The 2D cross-sectional profiles of flank surface at cutting distances were obtained using the
combination technique of CLSM and 2D wavelet transform in order to investigate the flank
wear patterns and evolution under various cutting conditions.
• Both TiCN and TiAlN coatings show high wear resistance against the abrasion while the
carbide has the least resistance in wear. After the carbide was exposed, the work material
adhered to the carbide layer, which could cause edge chipping.
• Only the abrasive wear is dominant until the carbide was exposed. Mainly, the flank wear
rate was dependent upon the hardness of the coating materials. Thus, to resist flank wear, a
harder coating, not necessary in a multi-layer scheme, is essential.
• A multilayer scheme did not provide a benefit by propagating the crack along the interfaces
since the hardness of coating is more of importance to resist flank wear. However, the
multilayer scheme may improve the coating strength and hardness in some applications and
enhance the wear resistance.

68

Chapter 4
TOOL WEAR IN DRILLING OF COMPOSITE/TITANIUM STACKS WITH CARBIDE
AND POLYCRYSTALLINE DIAMOND TOOLS

Hybrid structures provide structural advantages to components made from different
materials. In hybrid structure, various materials may work together or independently, but are
always better than a single material. In aerospace structural applications, the hybrid structure of
carbon fiber reinforced plastics (CFRP) composite and titanium (Ti) stack has been widely
utilized due to its corrosion resistance and unique ability to withstand high stress in service
[Zitoune et al., 2010, Brinksmeier and Janssen, 2002, Lantrip, 2008, Kim and Ramulu, 2007 and
Ramulu et al., 2001]. Drilling is essential to the stacked structures when conducting riveted
assemblies and structural repairs.
In the present work, tool wear mechanisms when drilling of CFRP-Ti stacks are
investigated. The drilling experiments were conducted with the tungsten carbide (WC) and PCD
tools with two different drilling speeds. The tool wear mechanism on the drills was characterized
through the microscope techniques such as scanning electron microscopy (SEM) and confocal
laser scanning microscopy (CLSM). The flank wear progression was also obtained by height
encoded (HE) image by CLSM and image processing program in Matlab Software. The relations
between drilling forces and tool wear measured will be discussed and the tool wear mechanisms
for WC and PCD tools in drilling CFRP-Ti stacks will be summarized.

69

4.1 TOOL WEAR IN DRILLING
4.1.1 CFRP drilling

A large part of research in the area of drilling CFRPs has been concerned with the hole
quality issues during the process. Since König and Grass [Konig, et al., 1985] studied the hole
quality parameters in drilling CFRPs and proposed the concept of the 'critical thrust force'
causing exit-ply delamination [Konig, et al., 1985]; composite damage or delamination, in
particular, during the drilling process has been extensively studied. Hocheng and Dharan, [1990]
proposed a mathematical model of exit hole delamination with respect to the critical thrust force.
Other researchers [Jain and Yang, 1993, Hocheng and Tsao, 2003, Lachuad et al., 2001,
Upadhyay and Lyons, 1999, Zhang et al., 2001 and Gururaja and Ramulu, 2009] developed
analytical models to predict the critical thrust force for delamination free drilling. Several
research works have also been performed on tool material and wear mechanism in the drilling of
CFRPs. CFRP is a highly abrasive material, depending on direction of the fiber in the matrix,
which produces severe tool wear. Unlike metals, CFRP cutting involves not only tool edge
chipping but also excessive abrasive wear due to the hard carbon fibers [Faraz et al., 2009,
Rawat and Attia, 2009, Lin and Chen, 1996 and Chen, 1997]. It has been commonly found that
the wear on the WC tools is generally smooth and uniform along the entire cutting edge of a tool.
The dominant wear mechanism on the WC tool in high speed drilling was edge micro-chipping
and abrasive wear by the hard fractured graphite fibers and carbide grains. Furthermore, the
carbon fibers can attack Co binder, which is relatively soft, causing to lower the binding force of
the carbide grains, which can accelerate tool wear and crack propagation [Rawat and Attia,
2009]. The flank wear on the WC tool, when drilling CFRP, was significantly affected by cutting

70

speed. The highest cutting speed occurs at the outer corner of the cutting edge, the round area of
the outer corner of the cutting edge grows as the drill penetrated further into the work material.
At higher drilling speeds, the rounded area at the outer corner of cutting edge grows faster than at
lower drilling speed [Lin and Chen, 1996]. Several empirical models were proposed by the
researchers [Tsao, 2008, Lin and Ting, 1995 and Iliescu et al., 2010] to relate the tool wear and
the drilling process parameters such as the feed rate and the cutting speed. The PCD tool is
commonly used for cutting composites because PCD tools have high wear resistance against
abrasion due to its high hardness [Klocke and Wurtz, 1999, Davim and Antonio, 2001, Nabhani,
2001, Ramulu et al., 1999 and Teti, 2002]. PCD tools reduced wear rate and improved hole
quality when compared to HSS and WC tools in drilling of a CFRP composite. It was found that
a small amount of micro-chipping was observed when drilling CFRP composites using a PCD
drill [Ramulu et al., 1999]. Due to its superior wear resistance, PCD has been the preferred tool
for drilling CFRP [Garrick, 2007].

4.1.2 Ti drilling

Tool wear mechanisms in drilling Ti are totally different from those in drilling CFRP. In
the case of Ti drilling, the distinctive material characteristic, such as low thermal conductivity
and strong chemical affinity to tool materials, leads to the reduction in tool life and might cause
catastrophic tool failure [Zhang et al., 2008 and Sharif and Rahim, 2007]. In addition, in Ti
machining, the contact area between chip and tool is shorter than that in steel machining. Smaller
contact area results in stress concentration at tool edge where the maximum cutting temperature
is reached. This can cause edge chipping or tool fracture [Hughes et al., 2004 and Stephenson
and Agapiou, 2006]. Coated drills suffered less damage, which suggested that the hard coating
71

material protected the tool and substantially reduced the wear rate of the substrate [Sharif and
Rahim, 2007]. When drilling Ti, Ti chips can be easily welded to the cutting edges of the drill
and formed built-up-edge (BUE), which leads to chipping and premature failure of the tool. A
PCD drill is one of the hardest materials, showing high wear resistance, and also has low friction
characteristic and thermal conductivity. However, due to the brittleness nature, the PCD tool can
fracture in drilling Ti mainly at high cutting speeds [Sharif and Rahim, 2007 and Garrick, 2007].
Therefore, it is recommended that PCD drill needs to be operated at relatively lower cutting
speed than that of WC tools. The machining conditions in drilling play an important role on tool
wear, tool life and hole quality.
Tool wear in Ti drilling is very sensitive to the drilling process parameters such as feed
rate, cutting speed, and coolant as well as the drill geometry and material [Zhang et al., 2008].
Increasing feed rate and/or cutting speed increased tool wear. In particular, a small change in
cutting speed could cause a very large change in tool wear rate [Yang and Liu, 1999]. A list of
optimal drill geometry (such as drill point angle, chisel angle, lip relief angle, splitting angle, and
the angle of notch) for drilling Ti was proposed by Barish, [1988]. Zhu and Wang, [2006] listed
the order of these factors affecting tool wear (from greater to smaller) as: relief angle, rake angle,
eccentricity of drill point, and side cutting edge angle. It was also reported that a drill material
with higher hardness and higher density was more wear-resistant [Zhu and Wang, 2006].

72

4.1.3 Hybrid structure drilling

Drilling of hybrid composite structures including CFRP and metal stacks has been
studied for about a decade [Zitoune et al., 2010, Brinksmeier and Janssen, 2002, Kim and
Ramulu, 2007, Ramulu et al., 2001, Kim and Ramulu, 2004 and Park et al., 2010]. Drilling of
CFRP-metal stacks at once poses several challenges due to different machinability of two or
more distinct materials. Different machining properties of the materials often lead to severe tool
wear, which often causes increased drilling forces, poor hole quality and large metallic burrs in
CFRP-metal stack drilling [Zitoune et al., 2010, Brinksmeier and Janssen, 2002, Kim and
Ramulu, 2007, Ramulu et al., 2001, Kim and Ramulu, 2004 and Park et al., 2010]. Stack
drilling, also, requires different optimal machining conditions for each material [Kim and
Ramulu, 2004]. It means that the ideal tooling solution for metal is not the most cost effective for
CFRP and the ideal tooling solution for CFRP cannot cut metal effectively [Kim and Ramulu,
2004]. For example, PCD has long been the preferred method of drilling CFRP, while WC drill
bits have been the preferred tool for drilling Ti. It was reported that the tool selection in drilling
CFRP-Ti stacks should match well with the Ti drilling condition rather than the CFRP drilling
condition. Excessive heating caused by Ti machining induced a number of hole quality problems
in CFRP [Ramulu et al., 2001, Kim and Ramulu, 2004].

73

4.2 EXPERIMENTAL PROCEDURES
4.2.1 Workpiece materials

The CFRP laminates acquired from The Boeing Company were multidirectional graphite
epoxy composites consisting of graphite fibers and an epoxy matrix with a quasi-isotropic ply
orientation of [(90o/-45o/0o/45o)4/0o/90o/0o/90o]s. The CFRP laminate had a thickness of 7.54mm
with an average ply thickness of 0.1141mm. The titanium plate material used is Ti-6Al-4V with a
thickness of 6.73mm. To obtain the grain size, a sample of the Ti was etched using Krolls
etchant. Using ASTM E112 Section 11 standard, the ASTM grain size number was calculated to
be 10.95. This number corresponds with an average α-grain size of 8µm.

4.2.2 Drilling experiments

Two types of twist drill materials were selected for this investigation based on their
availability and widespread usage in industry: WC and PCD. Both drills have a diameter of
9.525 mm and point angle of 135 o. The PCD drill has slightly larger dimension for the flute
length and the web thickness. The larger web thickness gives the drill a more ridged body, but
also can translate into a slightly higher thrust force. Table 7 summarizes the geometry
dimensions of the drills used in the experiment.
Drilling experiments were performed on a commercially available 3-axis CNC vertical
mill (MiniMill, HAAS, USA). The stack sequence of CFRP on top of Ti is typical stacking
sequence in aircraft structure assembly. The dynamometer (TRS-1K-OPT-THR, Transducer
Techniques, USA) had a fixture mounted on it to hold the stacks and ensure accurate force data.

74

The stacked materials were clamped to the fixture to enable any movement while drilling, this
way the force experienced by the materials will directly translate to the dynamometer. While
drilling, the thrust and torque forces were transmitted from the dynamometer to signal amplifiers,
A/D board (NI USB-6251, NI, USA) and recorded on a personal computer (PC) with data
acquisition software (LabView 7.1, NI, USA). The entire experimental step is shown below in
Figure 33.
Drilling experiment conditions were summarized and shown in Table 8. In order to
investigate how drilling speed effects WC tool wear, two different speeds (high speed and low
speed) were selected. CFRP and Ti are two materials that have very different optimal machining
settings. It is because of their dissimilar properties that the CNC mill was programmed to change
speed and feed depending on the material type. For each high speed or low speed drill, two
speeds are given depending on the material type. For example, the high speed for CFRP is set at
6000rpm while the high speed for Ti is set at 800rpm. The low speed WC tool drilled CFRP and
Ti at 2000rpm and 300rpm, respectfully. The drilling feed rate (mm/rev) also changes with the
material type and is shown in Table 8. CFRP is best machined at a higher speed and higher feed.
Titanium’s thermal properties limit the speed and feed. That is the reason that the speed and feed
of titanium is lower than that of the CFRP. By changing the speed and feed like this, each
material can be drilled at its optimum machining settings which will translate into a longer tool
life and a lower cycle time. A water-soluble cutting fluid was used as a mist coolant throughout
the operation with a constant flow rate of 16mL/min.

75

Table 7 Drill geometries
WC (9% Co ultra-fine grain)

PCD (Bimodal grade)

9.525
49.15
100.3
0.853
5.115

9.525
50.04
100.6
0.953
5.115

Point angle (deg.)
Helix angle (deg.)
Lip relief angle (deg.)

135
28
72

135
28
72

Chisel edge angle (deg.)

45

45

Drill diameter (mm)
Flute length (mm)
Overall length (mm)
Web thickness (mm)
Cutting lip length (mm)

Figure 33 Schematics of drilling experimental set up and drilling force data
acquisition system
76

Table 8 Drilling experiment conditions
Tool
RPM in CFRP
High speed
WC
6000
WC
2000
Low speed
PCD
2000

RPM in Ti
800
400
300

Feed: 0.0762mm/rev (CFRP), 0.0508mm/rev (Ti)
Coolant: Water-soluble cutting fluid, Mist coolant flow rate at 16 mL/min.

4.2.3 Wear evolution analysis

Tool wear was has been measured typically using a number of instruments and
techniques that include a tool-maker’s microscope, profilometer, white-light interferometry and
SEM. Most of these techniques cannot measure full 3 dimensional (3D) images required for
detailed analyses of tool wear. We have developed a new method of measuring tool wear in 3D
using CLSM [Park and Kwon, 2009 and Olortegui-Yume and Kwon, 2010]. The image
processing program, manly using wavelet filtering from CLSM, has been developed to eliminate
the noises and artifacts inherent to the height encoded (HE) image obtained by the CLSM. The
techniques introduced in [Park and Kwon, 2009 and Olortegui-Yume and Kwon, 2010] have
enable us to determine where the smooth contour due to dissolution or the scoring marks due to
abrasion has occurred and what inclusion has caused the abrasive wear. The wear evolution of
each drill was constructed by interrupting the drilling process after drilling a specified number of
holes (20 holes) on CFRP/Ti stacked materials. This enables us to examine and analyze the tool
wear mechanisms as the wear progresses into the drill. In this paper, because of the pronounced
adhesion layer of Ti on the cutting edge, we were able to use the same technique to estimate the
amount of the adhesion layers on various drills.

77

4.3 RESULTS AND DISCUSSION
4.3.1 Drilling forces

Figure 34 shows the drilling force change over the number of holes drilled. Given that the
stack structure is two different materials (CFRP on the top and Ti on the bottom), maximum
thrust and torque values were indicated separately for CFRP and Ti. When the drills were new,
both the drilling thrust and torque of the WC drill were 40% to 60% lower than those of the PCD
drill mainly due to the difference in tool geometries. As shown in Table 7, the web thickness of
the PCD drill was approximately 40% larger than that of the WC drill. The results agree well
with the previous experimental study results, which showed the drilling forces increase with
increasing web thickness [Chen, 1997]. As the number of holes drilled increase, the drilling
forces of both WC and PCD drills gradually increase in both CFRP and Ti. This increase is
mainly due to tool wear. Note that the drilling forces produced by WC drill at high speed
increased much faster than those at low speed. For the PCD drill, the thrust force did not increase
much over the number of holes drilled. In addition, the torque produced by PCD did not show
any significant overall increase over the number of holes drilled.

78

2500

WC - high speed
WC - low speeed

2000

Max. Thrust (N)

PCD - low speed
1500

1000

500

0
0

20

40

60

80

100

80

100

Hole number

(a) Thrust force
15

WC - high speed

Max. Torque (Nm)

WC - low speeed
PCD - low speed
10

5

0
0

20

40

60

Hole number

(b) Torque
Figure 34 Cutting forces vs number of holes drilled by the different tool types and
conditions
79

4.3.2 WC drill wear

Table 9 shows the SEM images of WC drills up to hole 80. The images mostly cover
flank face, cutting lip or cutting edge, and outer corner. Given that the stacking sequences of the
stacks are CFRP on top of the Ti, Ti adhesion is observed on all the drills after drilling a set of 20
holes. Adhesion is a well-known characteristic in Ti drilling due to the high chemical affinity to
typical tool materials [Park et al., 2010]. The adhesion leads to build-up edge (BUE) formation,
which is an accumulation of work material on the cutting edge best shown Table 9 for the WC
drill after drilling 20 holes. The BUE deteriorates the hole quality and leads to the tool chipping
or fracture when the BUE breaks off from the cutting edge. The adhesion was formed as soon as
Ti drilling started, and covered a large part of the flank surface especially at the high cutting
speed as shown in the CLSM images (Figure 36). More Ti adhesion was observed for the high
cutting speed rather than with the low cutting speed. The amount of Ti adhesion is directly
related to the chemical reactivity of Ti, which becomes higher at high temperatures.
In order to verify the tool wear geometries and wear progression without Ti adhesion, the
WC drills were etched using a Ti wet etchant (20 parts of H2O, 1 part of 30% H2O2, and 1 part
of 49% HF by volume). Micro-fracture or micro-chipping, with the approximate size of 30
microns, was observed at the outer edge on WC drills drilling at high speed after 60 holes. In
addition, as shown in Figure 35, the micro-chipping was observed at outer corner edge at hole 40
due to the combination of hard inclusion in Ti and adhesion. The hard inclusion and adhesion of
Ti attacked the sharp edge in the direction of scoring marks when Ti adhesion was detached. The
micro-chipping was only observed on the WC drills at high speed (large adhesion) while no such
chipping was observed on the WC drills at low speed and PCD tools (small adhesion).

80

Table 9 SEM images for carbide drills at high and low speeds with various holes

(a) Schematics of drill geometry
Figure 35 Drill geometry schematics and CLSM images of the WC drill outer
corner at high speed after 40 hole

81

Figure 35 (cont’d)

(b) Before etching (10X)

(c) After etching (10X)

82

To characterize the flank wear and quantity of adhesion for the WC tools, the CLSM was
used to take the HE images at every 20 holes with the objective of 50X before and after etching
as seen in the Figure 36 (a) and (c). Using the image processing program with wavelets, the noise
filtered 3D topographic profiles were obtained as shown in Figure 36 (b) and (d). Figure 36 (a)
and (b) illustrates the HE image of titanium adhesion on the cutting edge and corresponding 3D
topography image, respectively. On the other hand, Figure 36 (c) and (d) shows HE images and
3D topographic profile of the same tool after the Ti adhesion layer was etched away.

(a) HE image before etching (50X)

(b) 3D topographic profile before etching
Figure 36 CLSM images at flank face of WC drill after high speed drilling at 60th
hole
83

Figure 36 (cont’d)

(c) HE image after etching (50X)

(d) 3D topographic profile after etching
Figure 37 shows the cross-sectional profiles of carbide tool at various hole numbers
obtained from 3D topography profiles (Figure 36). The drilled tools have Ti adhesion covering
the WC tool surface regardless of drilling speeds (Figure 37 (a)). After Ti removal through
etching, the WC tool surfaces can be observed (Figure 37 (b)). From the observation, it was
found that wear first progressed on the cutting edge and flank surface at a simalar rate. In this
study, therefore, the two major wear types were observed to be flank wear (VB) and edge wear.
Edge wear has been identified as a major wear type in machining composite material, where the
84

cutting edge becomes rounded [Faraz et al., 2009]. Both types of wear were measured after
drilling sets of 20 holes as shown in Figure 38. The two wear types were caused by different
wear mechanisms due to dissimilar characteristics of the materials in the stacked structure. Flank
wear is the main mechanism from abrasion. The complex hard phase within titanium abrades the
flank surface after shearing the 3D microstructure. Edge wear is dominant as the fibers are being
cut by the cutting edge and the broken side of the fibers are scraping both flank and crater
surfaces of the drills.
At high speed drilling, surface profiles evolved as the flank wear land grew, especially
after drilling 40 holes, as in Figure 37 (a). At low speed, flank wear and edge wear occurred
almost equally as seen in Figure 37 (b). These figures shows the effect of drilling speed on tool
wear; much less flank wear was observed at low cutting speed than at high cutting speed while
the difference on edge wear was not significant and edge wear rate became steady regardless of
drilling speeds. Due to fast evolving flank wear at high speed, thrust force increase more rapidly
at high speed over low speed, as shown in Figure 35 (a). On the other hand, the difference in
drilling speed did not affect much on edge wear in CFRP cutting but when drilling Ti, especially
at high cutting speed, due to low thermal conductivity the cutting temperature became high
enough to activate high affinity of adhesion and to soften the WC drill. Softening of the WC tool
translates into the softening of the cobalt (Co) binder that holds the carbide grains together. The
flank wear is therefore extended by the mechanisms of carbide grains pulled out of the Co binder
as the Ti adhesion is removed by the next holes CFRP drilling, in addition to the abrasion by the
hard phase in the Ti alloy. In metal drilling, the cutting edge can be partially protected due to the
stagnant zone of the cutting edge maintaining the edge more consistently. This is why Ti drilling
had greater effect on flank wear rather than edge wear. On the other hand, when drilling CFRP

85

no stagnant zone is present. Without a stagnant zone the cutting edge becomes rounded and
dulled due to the abrasiveness of the fibers. Therefore, it can be said that flank wear land is
elongated by the Ti drilling and the edge wear is predominately due to the abrasiveness of the
hard carbon fibers when drilling CFRP.

(a) 2D cross-sectional profiles at high cutting speed
Figure 37 Tool flank wear profiles of carbide drills before and after etching at
various speeds

86

Figure 37 (cont’d)

(b) 2D cross-sectional profiles at low cutting speed
When drilling Ti at the low speed, the temperatures of the drills do not increase as much
as that of the drills at the high speed. Therefore, low speed conditions do not produce as much
flank wear. It was found that this amount of wear is small and similar to the amount of edge wear
done by the CFRP producing an evenly rounded tool edge (Figure 37 (d)). It is clearly identified
in Figure 6 (c) that the rate of flank wear and edge wear is pretty close for low speed drilling. In
relating Figure 38 (b) to (c) you can see how the edge wear has stayed relatively the same, while
the flank wear increase exponentially for high speed drilling after around 40 holes drilled due to
the cutting temperature. As the cutting forces go up, the cutting temperature increases [Weinert
and Kempmann, 2004]. The increase in cutting forces over the hole number as evident in Figure
34.
87

(a) Flank and edge wear measurement

(b) High speed drilling
Figure 38 Flank and edge wear progression of WC drills for various drilling
speeds

88

Figure 38 (cont’d)

(c) Low speed drilling
Figure 39 shows the effect of wear land size at flank and edge on the drilling forces.
Maximum thrust force increases with increasing flank wear length regardless of drilling speed.
The flank wear length is proportional to the flank surface area, which is the area normal to the
cutting direction. It is for these reasons that as the flank wear length increase, the parallel area of
contact between the tool and the work piece increases which translates into increased thrust
force. On the other hand, the maximum torque is well related to the edge wear length. As the tool
edge rounds and becomes blunter, the area of contact increases resulting in an increase in torque
need to cut the materials.

89

(a) Flank wear length vs maximum thrust

(b) Edge wear length vs maximum torque
Figure 39 Effect of wear progression on drilling forces

90

In summary, as the WC drill goes through the top CFRP plate abrasive wear occurs
predominately on cutting edge of the drill due to the hard carbon fibers. As, the chisel edge of the
drill begins to penetrate the top of the Ti plate and the speed substantially reduces, severe Ti
adhesion takes place. In Ti drilling, two different wear mechanisms are involved that mainly
wear on the flank face. First, abrasive wear occurs from the hard phase in the Ti alloy and,
second, carbide grains can be pulled out of the Co binder as Ti adhesion is removed. After
making the first hole, the Ti adhesion on WC cutting edge is brushed away by the carbon fibers
when CFRP drilling the next hole. Even after the Ti adhesion is removed, Ti adhesion will be
rebuilt on the WC drill cutting edges when drilling the bottom Ti plate.

4.3.3 PCD Drill Wear

The PCD drill shows a very small amount of adhesion compared to the WC drill. In
general the PCD drill has properties of high thermal conductivity and low friction at the cutting
surface [Stephenson and Agapiou, 2006]. Based on these characteristics, we expect that the
cutting temperature of the PCD drill be relatively lower than that of the WC drill and, thus result
in a reduced amount of Ti adhesion. In general, PCD tools provide lower propensity to adhesion
with titanium. However, the major problem with PCD tools is the brittleness. The brittleness of
PCD and harsh drilling conditions of the CFRP/Ti stacked material often lead to micro-fractures
and catastrophic failure as seen in Table 10. Table 10 shows the progression of the micro-fracture
(or chipping) at the cutting edge of the tool for low speed drilling. Moreover, the PCD tool at the
higher cutting speed failed catastrophically after making a few holes. The PCD drill failed
rd

catastrophically after making the 73 hole at low speed setting.

91

Figure 40 shows CLSM images displaying micro-fracture on the cutting lip. This microfracture might occur as the PCD drill penetrates the Ti plate. The transition between to two
materials allows for a slight impact and the beginning of the Ti metal chip. This small chip
thickness generates high stress concentration at cutting edge. When you combine that with high
cutting temperature due to low thermal conductivity of the Ti plate, edge chipping or fracture can
occur in the brittle PCD tool [Stephenson and Agapiou, 2006]. The chipping size on the cutting
edge was measured as shown in Figure 40 (d) and (e). The chipping size at hole 60 was
approximately 200μm in length and 60μm in depth.

Table 10 SEM images of the PCD tool at low cutting speed
(Tool experienced catastrophic failure at hole 73)

92

z( m)

(a) HE image before etching (50X)

(b) 3Dimensional profile before etching

(c) HE image after etching (50X)
Figure 40 Micro-fractures observed on flank side of PCD at low speed drilling and
at hole 60
93

Figure 40 (cont’d)
(B)

(A)

200
150
0

100

-40
0

50
50

100

150

200

250

300

0 y( m)

x( m)
(d) 3D profile after etching

Horizontal Profile (A)
-10
(A)
-30
-50

0

50

100

150

200

x( m)
Vertical Profile (B)

250

300

-10
-15
(B)
-20
0

40

80

120

160

200

y( m)

(e) Cross-sections of the 3D profiles
Figure 41 shows the CLSM images and tool wear profiles of the cutting edge where
micro-fracture was not observed. Less Ti adhesion was observed on the PCD drill compared to
the WC drill. As shown in Figure 41 (d), both flank wear and edge wear were observed on the

94

PCD tool, similar to the WC tools. When comparing the two tools materials at low speed, the
PCD tool exhibits better resistance to flank wear (Figure 41). Due to the superior hardness of
PCD compared with carbon fibers, tool wear should be much smaller during drilling of CFRP.
Figure 42 presents that the edge wear of PCD is only one third of that of the WC but flank wear
length for PCD is a half of that for WC. This supports an advantage that PCD drill shows
superior performance in cutting CFRP.

(a) HE Image before etching (50X)

(b) HE Image after etching (50X)

(c) 2D cross-sectional profiles
Figure 41 CLSM images and flank wear profiles of PCD drill before and after
etching at low cutting speed
95

Figure 42 Flank and edge wear progression of the PCD drill in low speed
In summary, the PCD drill may have minimal tool wear when cutting the top CFRP plate
due to its high hardness. As the drill penetrates the bottom Ti plate, micro-fracture occurs at the
outer edges due to brittleness of PCD. In addition, mild Ti adhesion will take place on the cutting
edges. When the PCD drill penetrates the CFRP plate again, carbon fibers abrade Ti adhesion
away. However, Ti adhered at the cutting lips again and micro-fracture may be enlarged when
penetrating the bottom Ti plate.

96

4.4 CONCLUSIONS

Tool wear analysis on uncoated WC and PCD drills was conducted in drilling of CFRP/Ti
stacks using various microscopic instruments. The conclusion can be drawn as follows:
As the number of holes drilled increase, the drilling forces of both WC and PCD drills
gradually increase in both CFRP and Ti due to tool wear. The drilling forces produced by
WC drill at high speed increased much faster than those at low speed. For the PCD drill,
the drilling forces did not increase as significantly over the hole number.
More adhesion of Ti was observed on WC drills than PCD drill. This adhesion was
removed by carbon fibers in drilling of CFRP. No significant micro-chipping occurred on
WC cutting edge tool but micro-chipping occurred on outer corner edge. For the WC drills,
the wear pattern is generally smooth and uniform along the cutting edges. The flank wear
was dominant at high cutting speed due to high cutting temperature in Ti drilling while the
edge wear and flank wear showed similar amount of wear length at low cutting speed.
Micro-chipping or micro-fracture was observed on PCD tool, in particular at the outer
corner edge. The PCD drill showed small Ti adhesion in Ti drilling and exhibited high
wear resistance in drilling of CFRP. Overall, the PCD was superior to WC tool in stack
drilling. However, major chipping was observed at the cutting edge when drilling Ti due to
the brittleness nature of PCD drills.
In CFRP/Ti stack drilling, different wear mechanisms for CFRP and Ti alloy took place.
Hard carbide fibers abraded more on cutting edge, which caused edge wear while hard
phase in Ti extended the flank wear land in addition to carbide grain pullout when Ti
adhesion was removed.

97

Chapter 5
A STUDY ON DROPLETS AND THEIR DISTRIBUTION FOR MINIMUM QUANTITY
LUBRICATION (MQL)
In modern metal cutting applications, the cutting temperatures and forces on a cutting
tool are typically high, which drastically reduces tool life. For this reason, the cutting zone must
be well lubricated and adequately cooled by applying coolants in most machining applications.
For optimal machining operations, the coolants must be maintained throughout their usage and
eventually disposed. Recently, minimum quantity lubrication (MQL) or near dry machining
(NDM) has been applied to reduce the cost associated with the maintenance and disposal of the
sheer volume of fluid being used and the conformation to environmental policies [Weinert et al.,
2004, Ueda et al., 2006 and Rahman et al., 2002]. In MQL, the high pressurized air sprays microscale lubricant droplets onto the tool-work interface. Even though it is very promising for many
machining applications [Rahman et al., 2002, Varadarajan et al., 2002 and Sharma et al., 2009],
several fundamental issues associated with MQL such as optimal conditions for the operations
and environmental concerns have not been fully addressed [Park et al., 2008].
Typical oil-based lubricants used in MQL do not have the heat capacity of water-based
coolants. Accordingly, the most important function provided by the lubricants in MQL is to
reduce the friction between tool and work material. Thus, the most important point in the MQL
process is that the lubricant must adequately cover the contact areas, i.e., tool-chip and tool-work
interfaces. Particularly, droplet size and distribution in terms of nozzle distance and air pressure
have to be considered because the smaller droplet, which can provide better penetration into the
cutting surface, are more preferable especially for micro machining applications where the
cutting surface is relatively small [Jun et al., 2008]. In addition, droplet distribution, which
associated with the wetting area by lubricant, can affect the cutting performance as well. Other
98

important variables are wettability on the tool surface [Park et al., 2008] and thermal-chemical
stability of lubricant. For example, vegetable oil shows poor the stability against oxidation and
thermal degradation while mineral oil has important health related issues as discuss later. In this
regards, several works have been reported not only for improvement of the thermal-chemical
stability by additives or chemical modification of the vegetable oil but for optimal lubricants in
MQL application [Sharma et al., 2009, Fox and Stachowiak, 2007, Suda et al, 2002 and Itoigawa
et al., 2006]. Thus far, the efforts for understanding of these MQL parameters have not been
considerably made. Therefore, a better understanding of the MQL application conditions such as
droplet size and distribution, air pressure, nozzle distance, and flow rate is needed prior to
machining process in industrial applications.
The operator’s health hazards have been a major concern whenever lubricants/coolants
are used in machining operations. For instance, in flood cooling, the workers are exposed with a
variety of respiratory disorders including occupational asthma, hypersensitive pneumonia and
loss of lung function and skin disease such as irritation, oil acne and skin cancer [Raynor et al.,
2005]. The major industrial concern for MQL is the use of very fine, airborne fluid droplets
which potentially pose a potential health risk for machine operators. This can be circumvented
by installing the vacuum system in tandem with the MQL sprayer. However, the extent of health
hazards depends on the concentration and the type of lubrication fluid used (mineral, vegetable
or animal, water soluble and semi- or synthetic) as well as the duration of exposure. The
characteristics such as mist droplets diameter and mass concentration generated are varied
depending on the type of oils [Thornburg and Lieth, 2000]. Mineral oil may cause more serious
damage to human body than vegetable oil. Mineral (petroleum- based) oil is retained in the lung
and caused lung fibroses while vegetable oil is eliminated from the lung within a few days

99

without any obvious damage to the lung [Pinkerton, 1928 and Shoskes et al.,1950]. With the
droplet (mist) size less than 4μm, even MQL can cause various occupational diseases. In fact it
was reported to deteriorate lung function even for short-term exposure [Thornburg and Lieth,
2000]. Therefore, the national institute for occupational safety and health (NIOSH) has
3

recommended that the exposure to metal working fluid mist is limited to 0.5mg/m for total
particulate mass as the recommended exposure limit (REL) [NIOSH]. Thus, a great interest
exists in the machining community in the development of a convenient technique to measure the
droplet sizes in terms of lubricant types and spray conditions.
To address various issues mentioned above, confocal laser scanning microscopy
(CLSM) is used for the first time for this purpose. CLSM is an established advanced microscopy
technique used in a variety of fields to obtain accurate surface topography [Hanlon et al., 2001].
In particular, CLSM can be successfully applied to characterize the crater wear of cutting tools
[Olortegui-Yume and Kwon, 2007]. Wavelet transform can be useful when the dynamic or
localized properties of surfaces are needed as opposed to the fixed wavelength harmonics
assumed by classical fourier transform (FT) or gabor transform [Maksumov et al., 2004 and
Rioul and Vetterli, 1991]. When wavelets are applied in cascade to the surface data, a multiresolution representation of the surface can be obtained, i.e. surface representations at multiple
scales [Fu et al., 2003]. The availability of the wavelet representations facilitates the researcher
to perform tasks such as data de-nosing and identification of roughness features.
In this chapter, we have developed a new technique to measure the droplet sizes in a
typical MQL process and their distribution after the droplets have been sprayed onto the polished
silicon wafer surface. CLSM has been utilized for three-dimensional (3D) surface
characterization of the droplet surface for the droplet volume measurement. The raw data from

100

CLSM contains noise and artifacts due to the inherent nature of data acquisition process. To
circumvent these problems, wavelets analysis has been applied by the use of de-noising schemes
and multi-resolution decomposition which helps to extract the droplet profile. In addition, edge
detection algorithm has been used to estimate the droplet volume for the calculation of the
diameter in an airborne droplet. These developments have been used to study the effects of air
pressure on the droplet size and distribution.

5.1 EXPERIMENTAL SETUP

Sharnoa SVC-60 CNC milling machine was used to provide the movement for the silicon
wafer mounted on the CNC table while the MQL dispensing unit (Unimax-Part No. 9570-7-512- UNIST) equipped with a flexible nozzle was used to spray vegetable oil-based lubricant
(UNIST-Coolube 2210). The MQL droplets were sprayed onto a silicon wafer to minimize the
surface roughness in order to accommodate the task of extracting the droplet shape in the
wavelets analysis. The set up for MQL spray experiment is shown in Figure 43. The droplets
discharged from the nozzle with the flow rate of 3.2ml/min, which was measured by detecting
the loss in the oil height though a graduated thin glass cylinder, are passed through the screening
plate with a hole in order to block most of the oil mist stream and thus avoid droplet overlap on
the wafer. After trying various hole diameters, it was set at 0.8mm because the droplet size did
not decrease any further below this value. The droplets were discharged as a function of air
pressure, P, and nozzle distance, D, in order to investigate their influence over the droplet size
and distribution. To evaluate the effect of air pressure, mass accumulation test was developed,
where an absorbent sheet placed on a wax paper was placed on the polished silicon wafer. With
the mass accumulation test, the oil droplets that are eventually attached to the silicon wafer can
101

be distinguished from the airborne droplets. This test was performed in terms of air pressure and
nozzle distance. Similarly, the sprayed droplets from the nozzle were absorbed into the mass
measuring paper at the pressure of 4 and 8psi at the flow rate of 3.2 ml/min.

P

D

Moving Table

Moving Table

Moving Table

Figure 43 MQL Experimental Set-up

102

5.2 CLSM DATA ACQUISITION AND IMAGE PROCESSING

The important characteristic of CLSM is the confocal aperture which eliminates the
reflected light from above and below the focal plane, thus obtaining surface information from a
very thin “optical slice” of the surface is feasible. Surface reconstruction is achieved by storing
and manipulating a stack of optical slices within a set of the vertical travel of the stage. The final
output of the built-in software in CLSM is a z-matrix (z(x,y)) containing the height information
of the surface. For a better handling of the image data, the z-matrix is encoded in a grayscale,
called height encoded (HE) image. Along with HE images, maximum brightness (MB) images,
which are extended depth of field views, can be also obtained. It is important to note that most
confocal systems can work in both confocal and optical (conventional) modes. The detail
information on CLSM can be found in [Hanlon et al., 2001 and Olortegui-Yume and Kwon,
2007]. In this work, A Zeiss LSM210 Confocal system was used to capture 2D images (confocal
slices) and HE images (topography) of individual droplets using Plan Neofluar 5X/0.15 and
50X/0.80 objectives. Figure 44 (a and b) shows the 2D CLSM droplet images at 4psi and 8psi,
respectively, obtained with the 5X/0.15 objective. These images are used for identifying the
droplet size distribution and the droplet coverage area fraction deposited on the wafer. The
boundaries of each droplet were detected using edge detection algorithm (EDA) called ‘Canny’
which looks for local maxima of the intensity gradient in an image. Then the areas enclosed by
detected boundaries were filled and considered as the sprayed droplets as shown on Figure 44 (c
and d). It was assumed that the areas smaller than 2 pixels (1 pixel=5.87265μm X
4.16792μm/5X objective) were rejected. On the other hand, for droplet size measurement, HE
images were obtained for various sizes of individual droplets, as shown in Figure 45.

103

The main advantage of the wavelet transform for image processing is to extract the
surface topography clearly from the raw image data without losing any surface details [Park and
Kwon, 2009 and Olortegui-Yume and Kwon, 2010]. In this work, the two-Dimensional Discrete
Wavelet Transform (2D-DWT) was used in a multi-resolution scheme with a two-channel filter
bank, which consists of a pair of filters, low-pass and high pass filters based on a chosen mother
wavelets, Daubechies7(db7) [Rioul and Vetterli, 1991]. For acquisition of the droplet topography
(3D), the wavelet de-noising was required to eliminate the electronic noise and artifacts from
individual droplet images (HE images). The de-noised surface topography data was then postprocessed using multi-resolution wavelets analysis. Finally, the droplet size can be obtained by
calculating the droplet volume (see Sec.5.3).

0.5mm

0.5mm

(a) Droplet distribution at 4psi

(b) Droplet distribution at 8psi

Figure 44 Droplets distribution optical images from CLSM and bitmap images
obtained by EDA at 4psi and 8psi

104

Figure 44 (cont’d)

(c) Bitmap image for (a)

(d) Bitmap image for (b)

50μm

(a) 2D optical image

(b) HE image of a droplet of (a)

Figure 45 Droplet images for droplet size measurement

105

5.3 DROPLET VOLUME CALCULATION

Once the droplets are deposited onto the flat silicon wafer surface, each droplet spreads
and changes its morphology while preserving the volume of each droplet. Using the HE image
information, the volume of droplet can be calculated as presented in Eq. (9).

N

M

V

∆V ,

∆x · ∆y · h ,

h
(9)

where ∆V , is the discrete volume at a point P ,
droplet at a point P , and h

, , , , h , is the surface height of

is considered as the height of the droplet edge. The total

volume of a droplet can be obtained by summing up the discrete volume, ∆Vi,j . The height z ,
is obtained as a result of HE image decoding. M and N correspond to the size of the HE image.
Once the droplet volume calculated, the diameter of a spherical airborne droplet,

D

defined as Eq. (10).

D

2

3·V
4π

/

(10)

106

, can be

5.4 RESULTS AND DISCUSSION
5.4.1 Droplet size and distribution

Figure 46 (a) shows air pressure effect on the oil mass accumulation test. After the oil
was sprayed in terms of the nozzle distance and air pressure, the oil mass accumulated on the
absorbent paper was measured. The average mass accumulated decreased as the nozzle distance
and the air pressure increased as shown in Figure 46 (a). Moreover, the oil mass could not be
measured at a higher air pressure because the spray was too vigorous so that the oil mist cloud
was generated in the air rather than accumulated on the absorbent paper. Thus, accurate data
could not be obtained at higher air pressure. Based on mass accumulation test, the short nozzle
distance is more preferable and too high air pressure might not wet the cutting surface effectively,
especially with the increase in the nozzle distance if the small cutting zone is considered.
The 2D CLSM images of sprayed droplets obtained under various operating conditions
(Figure 46 (b)) were converted into their bitmap versions (Figure 46 (c)), which were ultimately
used to calculate the average dimensions of droplets as well as the area fraction covered by
droplets. Table 11 summarizes the quantitative results of these calculations. It was observed that,
as the nozzle distance increased, the surface covered by the droplets decreased. The spray
conditions for the maximum surface coverage were observed at a nozzle distance of 30mm and
12psi nozzle pressure. This is because more droplets are released from the nozzle within a unit
time and thus more of them can cross the screening plate at a higher air pressure. However, the
area fraction decreased substantially at the nozzle pressure of 12psi as the nozzle distance
increased. For example, as seen in Table 11, at the air pressure of 12psi, the covered area fraction
dropped from 15.1% at the nozzle distance of 30mm to 1.7% at that of 70mm. The reduction of

107

the area fraction is more substantial at a higher air pressure. The droplet size at higher pressure is
smaller than that of lower pressure.
Therefore, it can be concluded that the nozzle distance should be less than 50mm and if
the high air pressure is required, the nozzle distance should be short enough to provide adequate
lubricant at the cutting zone.

(a) Mass accumulation in terms of air pressure and nozzle distance
Figure 46 Average mass accumulation, sprayed CLSM images, and
corresponding bitmap images at nozzle distances and air pressures

108

Figure 46 (cont’d)

4 Psi

8 Psi

30mm

50mm

70mm

90mm

110mm

(b) The Tiff images from CLSM

109

12 Psi

Figure 46 (cont’d)

4 Psi

8 Psi

30mm

50mm

70mm

90mm

110mm

(c) Bitmap images by EDA
110

12 Psi

Table 11 Average sizes, droplet count, and area fractions
NOZZLE AIR PRESSURE
NOZZLE
DISTANCE

4psi

8psi

12psi

Dav
(μm)

Nd

Af
(%)

Dav
(μm)

Nd

Af
(%)

Dav
(μm)

Nd

Af
(%)

30mm

12.93

888

9.3

11.28

1335

9.6

11.32

2066

15.1

50mm

15.18

595

8.9

12.66

471

4.5

10.54

797

5.0

70mm

13.94

362

4.5

12.29

217

2.0

11.85

217

1.7

90mm

15.47

151

2.4

11.05

485

3.4

11.70

160

1.2

110mm

13.32

144

1.6

11.04

71

0.5

10.15

58

0.3

Af: Area fraction covered by droplets (%)
Nd: Droplet count
Dav: Average diameter (μm)

5.4.2 Droplet volume calculation by wavelet transform

The general procedure used to calculate the droplet size includes the CLSM data
acquisition, the wavelets de-noising step, the wavelets multi-resolution analysis, and the droplet
volume calculation as presented in Figure 47. In addition, an empirical model is developed to
estimate the droplet diameters sprayed in the air based on the relationship between 2D CSLM
images and 3D HE images (see Sec. 5.4.3).

111

Figure 47 Flow chart for the calculation of the droplet size and distribution
Figure 48 (a) shows the raw HE image of an individual droplet obtained at the compress
air pressure of 6 psi and the nozzle distance of 55mm using the 50X/0.80 objective in the CLSM.
Each pixel size for the 50X objective is 0.14745μm×0.764706μm. In the HE image (Figure 48
(a)), the surface height at a position (x,y) is encoded such that the bright pixels indicate high
positions relative to dark pixels. For more accurate calculation, only droplet surface area was
cropped using image processing software as shown in Figure 48 (a). Figure 48 (b) shows the 3D

112

surface reconstruction of the droplet surface using the raw HE image in Figure 48 (a). The
droplet HE image was post-processed using a wavelet-denoising algorithm (Daubechies) to filter
out electronic noise and unusual pits/peaks (artifacts) present. This denoised HE image was used
as an input for the multi-resolution analysis up to level 7 using Daubechies7 (db7) mother
wavelet. After inspecting the decoupled approximation and detail surfaces, a threshold level was
defined beyond which distortion impairs the useful extraction of the droplet shape. Raw HE
images contain the topography of the surface mixed with noise and artifacts. However, even if all
the noise and artifacts are eliminated, the remaining information in the HE images could contain
the tilting of the silicon wafer stages and of the substrates (wafers in this case). Extracting the
droplet shape from a tilted surface is considerably difficult; thus, tilting was compensated using
mean slope obtained by the average of the slopes in both x and y directions both tilted shapes are
shown in Figure 48 (b) and (c). The slope compensation is carried out after wavelet multiresolution analysis and prior to the droplet volume calculation.

Cropped
area

(a) HE image of a droplet
Figure 48 HE image of the droplet (a), 3D surface profile (b) and 2D droplet
profiles in x (c) and y section (d) at level 3

113

Figure 48 (cont’d)

(b) Droplet surface topography from (a)

(c) Droplet surface profiles in x - section
114

Figure 48 (cont’d)

(d) Droplet surface profiles in y - section
Figure 49 presents the approximations and details surfaces (horizontal, vertical and
th

diagonal) obtained up to 7 level after applying the wavelet transformation including the
reconstruction (inverse transformation) process. It can be clearly observed that between levels 1
through 4 (Figure 49 (b)) the decoupling of the droplet (large features) from the details (small
features) is consistent with our purpose. In other words, the approximation surface contains
mainly the dominant shape (the droplet) while the detail surfaces contain much smaller features.
In this study, only the approximations were used for the droplet profiles and volume calculations
th

since the small features such as roughness and waviness were of no interest. After the 4 level of

115

th

th

decomposition, this behavior is no longer observed and between levels 5 and 7 larger features
are being extracted implying a major distortion in the droplet shape.
th

Figure 50 shows the approximation droplet surfaces for all levels up to the 6 level of
wavelet decomposition. It is clearly observed that the fourth level is the threshold limiting the
st

th

feasible extraction of the droplet shape. Thus, the images between 1 and 4 levels with the
droplet pattern clearly were used for the volume calculation.

st

(a) 1 level
Figure 49 The Droplet surfaces after wavelet filtering

116

Figure 49 (cont’d)

th

level

(b) 4

th

(c) 5

level

117

Figure 49 (cont’d)

th

(d) 7

level

st

nd

(a) 1 level

(b) 2

level
st

Figure 50 Droplet surface after wavelet transform in 1

118

th

to 6

level

Figure 50 (cont’d)

rd

(c) 3

th

level

th

(e) 5

(d) 4 level

th

level

(f) 6

level

Using the droplet surface of Figure 50 (c), the estimation of droplet volume was
3

conducted. The total calculated droplet volume using Eq. (1) at level 3 is 2.1084E-004mm . The
droplet volume calculated in each level can be shown as in Figure 51 (b). The calculated droplet
st

th

volumes from 1 to 7 level of wavelet decomposition are similar, but it shows a large deviation
th

above 7 level. With this droplet volume calculation technique, further analysis with other
droplet details such as wetting angle and average droplet size, can be conducted for various
commercial lubricants. In this sense, the equivalent droplet radius is estimated at level 3 as 36.92

119

μm using Eq. (2) and the equivalent droplet radius in several levels can be shown as Figure 51
(c). It shows the similar trend with droplet volume at each level. As shown in Figs. 51 (b) and (c),
th

the droplet surface profiles were completely lost at 8 level.

(a) Final droplet surface

(b) Estimated volume
Figure 51 Droplet surface at level 3 (a) and its corresponding estimated droplet
volume (b) and equivalent radius (c)

120

Figure 51 (cont’d)

(c) Equivalent radius

5.4.3 Small droplet size estimation
The capturing of HE images (3D data) for very small droplets with the average diameter,
Dav<100μm, can be challenging and time consuming. Thus, an estimation method was pursued.
A diameter estimation equation for a spherical airborne droplet (3D) can be achieved by using
the CLSM optical image (2D version) and the calculated volume of a droplet (3D). For example,
both the optical and HE images for the droplets are taken by CLSM. Using volume calculation
method in Sec. 5.3 and Eq. (2), 3D diameters of airborne droplets can be obtained. Figure 52 (b)
and (d) represent the droplet volumes calculated by wavelet-filtering for different droplet sizes
and Figure 52 (a) and (c) are their corresponding 2D droplet images, respectively. In addition, for

121

the small droplet estimation, relationship between 2D and 3D droplet diameters was determined
by using interpolation method as in Eq. (11).

D

D

0.0012D

D

0.1997D

D

0.0987

(11)

where D2D is the 2D diameter of a droplet on the silicon wafer and D3D is the diameter of an
airborne droplet. This empirical correlation between the diameters in the 2D images and the
diameters in the 3D data (HE images) should be limited to the combination of the lubricant and
the silicon wafer surface used in this study. To justify the use of this equation, other combination
of lubricant and surface need to be studied. Finally, 2D and 3D droplet diameters for various
sizes of the droplets for the vegetable oil can be obtained as summarized in Table 12. Finally,
without measuring the extremely small droplets, the size of oil droplet sphere sprayed from the
nozzle can be estimated. Using Eq. (3), the size distribution of droplets in Figure 45 (a), 2D
images, was analyzed and the average 3D droplet size was around 10μm. Besides, the most of
the droplet sizes in this test is less than 20μm, which might be small enough to be effective in
practical MQL application, as shown in Table 13. The proposed method of measuring the droplet
size is particularly useful to evaluate the very small droplet sizes if the inhaling of MQL droplets
by operators is harmful.

122

(a) 2D diameter (149.7μm)

3

(b) Droplet volume (9.2337E-005mm )

(c) 2D diameter (173.4μm)
Figure 52 Droplet volumes corresponding 2D droplet sizes
123

Figure 52 (cont’d)

3

(d) Droplet volume (1.8273E-4mm )

Table 12 Droplet size calculation
2D droplet diameter (μm)
0
0
1
96.1
2
149.7
3
173.4
4

3D droplet diameter (μm)
0
29.2
56
70.4

183.2

73.8

Table 13 Droplet size distribution
Number of Droplets (2D)
Diameter
Ave size (43.78μm)
9
D < 20μm
30
20μm <= D < 40μm
14
40μm <= D < 60μm
12
60μm <= D < 80μm
21
80μm <= D < 100μm
17
100μm <= D

124

Number of Droplets (3D)
Ave size (10.02μm)
123
17
1
0
0
0

5.5 CONCLUSION

The following conclusions can be drawn from the present chapter:
•

The important MQL parameters such as droplet size and distribution were studied in
terms of the nozzle distance and air pressure.

•

The experimental technique to measure the droplet sizes and distribution for MQL
combining CSLM and wavelets analysis proposed and proved to be successful.

•

Using the edge detection algorithm, the boundary of each droplet on the substrate was
identified and the area fraction covered by the droplets based on the area and average size
(diameter) of droplets can be determined for various nozzle-workpiece distance and
nozzle air pressure.

•

Based on the observation, it can be inferred that the higher nozzle pressure provided the
more droplets but the smaller droplets were obtained while a smaller amount of droplets
was deposited on the surface when the nozzle distance was increased. The maximum area
fraction covered by droplets was obtained at the nozzle distance of 30mm and the air
pressure of 12psi. However, high air pressure does not provide the condition to wet the
cutting zone especially with the nozzle distance more than 50mm.

•

The volume of a droplet was calculated using wavelet transform (Daubechies7 mother
th

wavelet). The wavelet decomposition of the droplet surfaces was performed up to 7
th

level and the 4 level is the threshold limit for the feasible extraction of the droplet shape.
st

th

Therefore, the images between 1 and 4 levels can be used for the volume calculation.

125

•

The size of small droplets (less than 100μm) can be estimated by using the equation that
represents an empirical correlation between diameter of 2D optical droplet images and
3D droplet images. The average diameter of 3D droplets was 10.02μm. This droplet size
information in terms of lubricant type and spray condition can be useful not only to
evaluate the penetration ability of lubricant into cutting zone but to determine the
operator’s health hazards, especially for very small droplets.

•

The proposed approach of the MQL parameters in this work can be used to determine
optimal MQL conditions prior to each machining application, which ultimately increases
the efficiency of the machining process. The optimal MQL parameters will be depending
on the particular MQL oil used and an operating condition of a particular MQL system.
Thus, this paper introduces a new method to for such parametric study

126

Chapter 6
MQL WITH NANOGRAPHENE-ENHANCED LUBRICATES: BALL MILLING
EXPERIMENT

In a view of green manufacturing, dry machining has drawn much attention from
industries recently. Remarkable improvement on tool coating technology has facilitated the
complete elimination of cutting fluid in some machining processes. In most applications,
however, the cutting fluids are still required due to the high friction and adhesion tendency
between work and tool materials and the difficulty of chip and heat removal [Sreejith and Ngoi,
2000, Klocke and Eisenblatter, 1997 and Kelly and Cotterell, 2002]. Therefore, in some
applications, an optimal solution for lubrication condition can be found between dry machining
and flood cooling. In this regards, minimum quantity lubrication (MQL) is a viable solution to
overcome the disadvantages of both dry and flood cooling in machining processes.
The effectiveness of MQL process has been identified in turning [Dhar et al., 2006 and
Kamata and Obikawa, 2007], milling [Rahman et al., 2002 and Lacalle et al., 2006] and drilling
[Tasdelen et al., 2008, Heinemann et al., 2006, Davim et al., 2006 and Filipovic and Sephenson,
2006]. In addition, MQL process parameters, such as nozzle distance and direction, discharge
pressure, flow rates, etc., have been studied to optimize the machining conditions [Park et al.,
2010]. By means of confocal laser scanning microscopy (CLSM) and edge detection algorism
techniques, optimal nozzle distance and air pressure of vegetable based oil was determined [Park
et al., 2010]. It was found that both nozzle position and oil flow rate had the important roles on
machining performance [Lacalle et al., 2006 and Ueda et al., 2006].
Several works have been published to evaluate the characteristics of various lubricants
for finding optimum lubricants in MQL machining. The mixture of water and oil-free synthetic

127

lubricant provided longer tool life compared to both synthetic ester with and without alcohol in
deep-hole drilling due to the high cooling ability and low viscosity of water [Heinemann et al.,
2006]. A new lubricant type, oil film on water droplets, was proposed to provide a better cooling
ability on top of improved lubricity and showed less cutting forces in end milling [Itoigawa et al.,
2006 and Yoshimura et al., 2005]. Cutting performance in the tapping test as well as secondary
factors (biodegradability, oxidation, and storage stability) for synthetic polyor esters was superior
to vegetable oil [Wakabayashi et al., 2006 and Suda et al., 2002].
Solid materials (nano-sized) have been developed as a lubricant itself, which can
eliminate cutting fluids, or as an additive for lubricant. For the solid lubricant materials,
molybdenum disulfide (MoS2), graphite, boron nitride and polytetrafluoroethylene (PTFE) have
been used as dry powders or coating materials [Jianxin et al., 2009, Krishna and Rao, 2008 and
Xu et al., 2004]. In the machining test, these solid lubricants reduced the cutting forces and
surface roughness [Krishna and Rao, 2008, Reddy and Rao, 2006]. On the other hand, the solid
lubricants can be also mixed with the oil based lubricants. In the grinding processes, MQL
lubricant with MoS2 nano-sized particles was introduced in [Shen et al., 2008] and MoS2
grinding lubricant showed excellent performance on cutting forces in spite of its low dissociation
temperature at 350°C in oxidizing environments. The performance of diamond nano-particles in
oil was evaluated in terms of friction and a wear test using ball on disc type equipment, which
showed high loading capacity and less friction and wear [Tao et al., 1996]. The nano-graphene
platelet enhanced lubricants were studied in various conditions [Park et al., 2010, Huang et al.,
2006, Lee et al., 2009 and Alberts et al., 2009]. The nano-particle enhanced lubricants are
expected to provide better performance because nano-graphene has not only a higher dissociation
temperature but also the high aspect ratio, which naturally orients the graphene phase for better
128

lubricity when applied through MQL process. In addition, the type and concentration of solid
lubricant in oil can be also important factors for an optimal use of the cutting fluids in machining
applications [Krishna and Rao, 2008 and Alberts et al., 2009]. The machining performance is
also affected by the additive in lubricant. Boric acid with 20% by weight mixed with SAE 40 oil
was found to be optimal lubricant for tool wear and surface roughness in turning of EN8 steel
[Krishna and Rao, 2008].
This paper studies the effect of the exfoliated nano-graphene enhanced lubricant by
conducting wetting angle measurement, friction tests using a ball-on-disk set-up and MQL ball
milling test. First, wetting angles for a variety of commercially available lubricants including
nano-graphene enhanced oil were tested in order to evaluate the wettability of lubricants.
Second, the friction coefficients of each lubricant and a dry nano-graphene coated insert were
measured as a function of load and speed to determine if the introduction of nano-graphene,
either in a discrete particulate, or multiple nano-graphene layer format, will reduce the friction.
Third, ball milling tests were performed to evaluate the proposed nano-graphene enhanced
lubricant in a practical machining application.

129

6.1 NANO-GRAPHENE ENHANCED LUBRICANT

Exfoliated graphite nanoplatelets (xGnP), produced by XG Sciences, Inc. (East Lansing,
Michigan), were obtained from acid intercalated expandable graphite and a microwave
exfoliation method [Park et al., 2010 and XG Science, Inc.]. In this chapter, these xGnPs were
mixed with vegetable oil to yield a nano-enhanced MQL lubricant, which was patented by Kwon
and Drzal [Kwon and Drzal, 2009]. Figure 53 shows the xGnPs used in this study, where the
diameters are between 1 to 15μm (Figure 53 (a)) and the thickness are 10nm (Figure 53 (b)).
The MQL vegetable-based oil (provided by Unist, Inc. Grand Rapid, MI) was mixed with
0.1 and 1.0 wt% of xGnPs in a high shear mixer (SpeedMixer DAC 150FVZ-K from FlackTek,
Inc.) for a stable suspension. When the xGnP-enhanced lubricant was used in MQL machining, it
was expected that the larger size of xGnPs can be easily adhered on the cutting surface. When
two interacting surfaces are sliding with the xGnP-enhanced lubricant, the friction is expected to
be reduced as one graphene sheet is sliding over another. As shown in Figure 54, however, the
large diameter (15μm) xGnP particles with both 0.1 and 1.0wt% were segregated only within a
day (Figure 54 (B and C)). The 0.1wt% xGnP with 1μm diameter oil mixture showed the
stability even over a year (Figure 54 (A)) while that of 1.0 wt% was separated after a few days
(Figure 54 (B)). Therefore, the oil mixed xGnPs with 1μm diameter was only used; 0.1wt%
xGnP oil was directly applied using MQL equipment while 1.0wt% oil was applied by means of
a magnetic stirrer to keep a uniform concentration spray during the machining.

130

10nm
(a) SEM image (Diameter)

10nm

(b) TEM image (Thickness)
Figure 53 (a) SEM and (b) TEM images of xGnP

A

B

C

D

Figure 54 Suspension stability of xGnP oils (A: 1μm / 0.1wt%, B: 1μm / 1.0wt%, C:
15μm / 0.1wt%, D: 15μm / 1.0wt%)
131

6.2 EXPERIMENTAL SETUP
6.2.1 Wetting angle measurement

The wetting angles of lubricants can be used to evaluate the wetting ability of lubricants
on a cutting tool surface. A droplet of lubricant with a high surface tension resting on a low
energy solid forms a spherical shape giving a high wetting angle. On the contrary, when the solid
surface energy exceeds the liquid surface tension, the flat droplet profile with a low wetting
angle is considered to have a good wettablility. As is shown in Figure 55, the wetting angle θ is
defined by measuring the tangent line at the interface between the droplet and the cutting tool
surface as shown in Figure 55.
Figure 56 shows the wetting angle measurement setup and various lubricants including
tap water were selected for wetting angle measurements on the TiAlN coated carbide. The
motorized syringe assembly manufactured by AST Product, Inc. dispenses a 0.5μl droplet onto
the coating surface. The CCD camera then captures the droplet image produced by the LED
black light. Thus, wetting angles were measured by detecting both boundaries of the droplet and
the solid surface in contact angle measurement software.

Figure 55 Wetting angle of a droplet
132

Figure 56 Wetting angle measurement setup

6.2.2 Tribometer test

As shown in Figure 57, the friction tests were performed with a liner ball-on-disc type
tribometer (CSM Instruments) where a steel ball (440C) with 6.36mm in diameter oscillated on a
TiAlN-coated carbide flat surface in order to evaluate frictional characteristics in various
lubrication conditions. The test conditions include dry, vegetable oil (Unist-Coolube 2210), eight
layers of graphene coatings and vegetable oil mixed with graphenes of 0.1 and 1.0wt%.
Graphene coating technique can be found in [Biswas and Drzal, 2009], Loadings were 1, 5 and
10N and sliding speeds were 0.25, 1.0 and 2.5cm/s.

133

Figure 57 Tribometer setup

6.2.3 Ball milling Test

The MQL dispensing device (uni-MAX) provided by Unist, Inc. (Grand Rapid, Michigan)
was used to provide the lubricant in a mist form to the cutting area. The lubricant, Unist Coolube
2210, was sprayed through an external co-axial nozzle. The nozzle outlet pressure and flow rate
can be adjusted with the air metering screw and pulse duration/frequency in the control panel,
respectively. However, we have used an optimized output pressure of 8psi and a flow rate of
1.5ml/min determined based on our previous work [Park et al., 2010 and Park et al., 2010(2)].
Ball-milling experiments were performed on a 3-axis vertical milling center (Sharnoa
CNC mill) in dry and MQL conditions as shown in Figure 58. AISI 1045 steel
(203.2mm×127mm×203.2mm) and 25mm diameter ball-nose TiAlN coated carbide inserts
(ZPFG250-PCA12M) produced by Hitachi Tool Engineering were used for work and tool
materials, respectively. The cutting started at one corner of the work material in the direction of
134

203.2mm and continued line by line until a layer of material was completely removed. The tool
wear was also measured after cutting each layer to record the progress of tool wear. The
machining and lubrication conditions are summarized in Table 14.

Figure 58 MQL and ball milling setup

135

Table 14 Machining conditions
Cutting Speed (RPM)

3500 & 4500

Feed Rate (mm/min)

2500

Axial DOC (mm)

1

Radial DOC (mm)

0.6
Dry
NRG mineral oil 1: 15 (Water)

Lubrication

Unist Coolube 2210 (Vegetable oil)
Unist oil with xGnP (0.1 & 1.0wt%)

Confocal laser scanning microscopy (CLSM) was used for the measurements of tool wear
at the center of the ball-nose tool (central wear) and flank surface (flank wear), which were
illustrated in Figure 59. Central wear can be attributed to 3-body abrasive wear due to the debris
from the work material sliding between tool and work materials while flank wear comes from 2body abrasion when cutting through complex pearlitic cementite phase constrained within the
work material.

(a) Central wear

(b) Flank wear

Figure 59 Tool wear measurements of a ball nose insert
136

6.3 RESULTS AND DISCUSSIONS

Figure 60 shows the result of wetting angle measurements on various lubricant candidates
for MQL machining. Based on the assumption that a smaller wetting angle indicates a better
wettablility to the tool surface, the wetting angle can be an important parameter that elucidates
the lubrication performance. The result shows that water itself or water-based lubricant showed a
higher wetting angle compared to an oil-based lubricant, as expected. Moreover, Unist oil with
xGnP showed the best wettablility among the lubricants. It is expected that the xGnP-enhanced
lubricant provides effective lubricity at the cutting interface in the machining process. This
argument will be verified in ball milling test in this section. The oil based lubricant included
100% mineral oil produced by NRG Resources (NRG oil), which is water soluble oil. However,
the pure NRG oil has a high viscosity (hard to spray) and is intended for use as a water soluble
lubricant. Therefore, in this study, 100% NRG oil was not tested further.

Figure 60 Wetting angle test results
Various lubrication conditions including dry, vegetable oil, eight layers of graphene
coatings and a combination of the latter two were tested under a variety of loads and speeds for
the tribological characteristics. In a dry condition, the friction coefficient was inversely
proportional to the relative sliding speed as shown in Figure 61 (a) while as the load applied
between ball and TiAlN-coated carbide increased, the friction decreased (Figure 61 (b)).
137

Similarly, Unist oil showed the same result as in dry. When the xGnPs are applied on the track of
sliding, it was not effective in reducing the friction because the loosely applied xGnPs cannot be
sustained in their positions at the interface between ball and surface. Alternatively, the eight
layers of graphene (5nm thickness) have been coated onto the TiAlN coated carbides using the
coating technique described in [Biswas and Drzal, 2009], which consequently reduced the
friction in the test. Figure 62 (a) and (b) compares the friction coefficients between dry and
xGnP-coated samples from beginning to 300s. The graphene layers showed excellent
performance in reducing the friction. Even if the most of graphene coatings were delaminated
after about 100 seconds, the friction did not increase compared to the dry case because the
xGnPs detached from the coating are suspected to be remained at the interface still working as a
lubricant. Especially, the friction was fairly stable at the beginning of the run, which represents
the outstanding lubricity of the graphene phase. In this regard, it can be stated that the friction
can be effectively reduced if the graphene phase can be applied appropriately at interface.
However, when Unist oil was applied onto the graphene-layered, TiAlN coated carbides, the
friction did not reduce further because the xGnP-debris from the coatings taken from the sliding
surface would neither be necessary to relocate into the interface nor be well mixed with the oil
being applied. This is why the nano-graphene particles must be uniformly suspended in the oil.
Similar to the wetting angle test result, the xGnP-enhanced Unist oil showed the lowest friction
coefficient as seen in Figure 63 (a). In addition, the effect of the concentration of xGnP in Unist
oil was investigated; and 0.1wt% xGnP of 1μm diameter showed a slightly lower friction
coefficient than 1wt% of that and pure Unist oil, as in Figure 63 (b). And note that the xGnP of
15μm diameter lubricant was not tested due to its poor suspension stability in the oil as
mentioned in Sec 6.1.

138

(a) Load effect (Dry)

(c) Speed effect (Dry)
Figure 61 Tribometer test results for various lubricants

139

Figure 61 (cont’d)

(c) Speed effect (Unist oil)

(a) Dry
Figure 62 Comparison of friction coefficients of dry and xGnP coating
140

Figure 62 (cont’d)

(b) Graphene coating (8 layers)

(c) Microscope image after 100s

141

(a) Various lubricants

10N at 2.5cm/s

Friction Coefficient

0.1

0% xGnP (Unist)
1.0% xGnP
0.1% xGnP

0.05

0

-0.05

-0.1
0

100

200

300

400

500

600

700

800

900

Time (s)
(b) Various xGnP concentrations
Figure 63 Comparison of friction coefficients of lubricants
Figure 64 represents the central wear measured at 3500rpm for dry, NRG water-soluble
oil, Unist oil, and Unist oil mixed with xGnP. It was reported that central wear occurred for
coated carbide end mills with relatively low feed rate (less than 1000mm/min) in high-speed

142

milling [Sokovic et al., 2004]. However, in this test, the central wear was observed, as seen in
Figure 64, even with the feed rate of 2500mm/min. Especially in dry machining the central wear
became significant because the center of ball nose insert is always contacting the work material
without the cutting fluid medium reaching the area. Thus, the wear particles cannot be evacuated
and abraded the cutting tool. Unlike Unist oil, the mineral NRG oil did not provide any
beneficial effect on the central wear; therefore, the NRG oil was eliminated from any additional
experiment.

(a) Dry

(b) NRG water soluble oil

(a) Unist oil

(b) Unist oil with xGnP (0.1wt%)

Figure 64 Central wear at 3500rpm at 8th layer
143

Flank wear was also measured at various conditions. Figure 65 compares the effect of
MQL lubricants on the flank wear at the spindle speed of 4500rpm. Unist oil mixed with xGnPs
was superior compared to dry and Unist oil only, regardless of the concentration of xGnP.
Moreover, considering edge chipping, the beneficial effect of xGnP became more obvious. As
seen in Figure 65 (a) and (b), the machining experiments in dry and Unist oil only show chipping
at the cutting edge while, as shown in Figure 65 (c) and (d), the cutting edge maintained its
sharpness. Therefore, the presence of xGnPs has deterred the chipping at the cutting edge. It can
be concluded that the xGnP stay at the cutting interface by providing the lubricity and relieving
the mechanical impact generated during the milling process.

(a) Dry

(b) Unist oil only
th

Figure 65 Flank wear at 4500rpm at 8

144

layer

Figure 65 (cont’d)

(c) Unist oil with xGnP (0.1wt%)

(d) Unist oil with xGnP (1.0wt%)

To investigate the effect of the xGnP concentration, both central wear and flank wear
th

measured after machining each layer up to 8 layer were plotted against spindle speeds as shown
in Figure 66. In terms of reducing the central wear, the xGnP-enhanced lubricant is substantially
better than traditional MQL vegetable oil at both cutting speeds. The overall performance of the
0.1wt% xGnP oil was superior to that of the 1.0wt% one. This result corresponds with the
friction test result as the higher friction for 1.0wt% xGnP could cause slightly increase in flank
wear. As mentioned in Sec. 6.1, due to the notable amount of the xGnP precipitation for the
concentration of 1.0wt%, it had some difficulty in spraying a uniformly distributed lubricant in a
machining application, which might lead to diminish the effectiveness of xGnP lubricant.
Remarkable improvement on the central wear was observed when the xGnP enhanced lubricant
was used, as shown in Figure 66. In fact, the central wear is almost non-existent with the xGnPenhanced vegetable oil with the concentration of 0.1wt%, providing the best results. Therefore, it
can be stated that the optimal concentration of xGnP is around 0.1wt%, which means that the

145

higher xGnP concentration in the lubricant does not necessary provide any additional benefit on
friction reduction and machining performance. In summary, the flank wear (VB) at high cutting
speed was larger than that at low cutting speed as seen in Figure 66 (c) and (d)). Only Unist oil
th

th

showed the larger flank wear at low speed at 6 and 7 layers. This is due to the catastrophic
increase in wear rate during the interrupted cutting process where mechanical and thermal shock
can be established. Again, flank wear rates of xGnP 0.1w% at low and high cutting speeds
increased steadily while Unist oil showed unstable wear rate, resulting in the outstanding
performance of xGnP 0.1wt% oil as MQL lubricant in milling process.

(a) Central wear at 3500rpm
Figure 66 Central wear and flank wear in terms of concentration of xGnP

146

Figure 66 (cont’d)

(b) Central wear at 4500rpm

(c) Flank wear at 3500rpm
147

Figure 66 (cont’d)

(d) Flank wear at 4500rpm

148

6.4 CONCLUSION

In this chapter, the effect of the nano-graphene additive (xGnP) in vegetable oil was
investigated by comparing lubricants (dry, water-soluble oil and vegetable oil only) in the
wetting angle measurement, the tribological behavior, and the ball milling. The results from the
wetting angle test and the friction coefficient measurement show that xGnP-enhanced vegetable
oil improved the wettablility on the cutting surface and reduced the surface friction. These results
were verified by the ball milling experiment, which indicates that MQL machining provides
better performance, especially for the central wear and chipping at cutting edge. In addition, the
larger diameter xGnP did not show a stable suspension in vegetable oil, which caused
segregation of graphene and oil. The concentration of xGnP 0.1wt% with 1μm diameter showed
outstanding cutting performance without any segregation problem.

149

Chapter 7
CONCLUSIONS

In this work, a tool wear analysis in various machining processes, such as turning, milling
and drilling , and a study on MQL parameters including droplet size and distribution, nozzle
distance, outlet air pressure, lubricant types, etc., which might affect the machining efficiency,
were performed. In addition, the combination of CLSM and image processing techniques
including wavelet transform was used for obtaining the 3D surface topography and developing
droplet size measurement techniques.
In turning of AISI 1045 steel, the dominant wear phenomenon was abrasion by hard
inclusion in work material and then adhesion, by high affinity of steel and carbide, took over
after carbide was exposed. In addition, 2-body abrasion model in conjunction with FEM
simulation was used for flank wear prediction and the predicted wear profiles showed very good
results compared with the experimental data. Based on the combination of model and
experimental approaches, no significant beneficial effect on multilayer coating was observed for
flank wear as a single hard coating is adequate to resist abrasion. In contrast, for the crater wear,
multilayer coating was much effective due to not only Al2O3 middle layer coating, which has
high wear resistance against dissolution, but also the changes in interface conditions of
multilayer coating as each layer was being exposed, which resulted in the maximum crater
location changes instead of concentrating on one location.
In a face milling study, multilayer cutting tools, double (TiN/TiAlN) and triple (TiN/
Al2O3/TiCN) layered coated carbide, processed by PVD and CVD respectively, were evaluated
in terms of various cutting conditions. As expected, depth of cut (DOC) did not play a significant

150

role in the evolution of flank wear while the cutting speed had an important role in flank wear.
Similar to the turning case, abrasion was found to be the most dominant tool wear mechanism in
milling. Edge chipping and micro-fracture were the tool failure modes. After comparing the
performance of the two inserts, it was concluded that the double layer coating was superior to the
triple layer coating under various cutting conditions due to the benefit coming from the coating
deposition processes themselves. Additionally, it is commonly claimed that the superior
performance of the multilayer coating comes from preventing the gross fracture or crack into tool
materials instead of propagating the crack along the coating interfaces. However, no such
observations were found in both turning and milling experiments. Mainly, the flank wear rate
was dependent upon the hardness of the coating materials.
A similar study on drilling has been conducted. In particular, drilling of carbon fiber
reinforced polymer (CFRP)/titanium (Ti) stacks was performed using carbide and PCD drills.
Due to the dissimilar mechanical and thermal properties of the CFRP/Ti stacks, wear behaviors
in drilling CFRP and Ti drastically was changed. The dominant wear mechanism for carbide
tools was the abrasion by fibers in CFRP drilling and Ti adhesion, covering entire cutting edge
while the main wear behavior of PCD drill was edge chipping. The adhesion of titanium seemed
to be the most important factor in tool wear and micro-chipping because the adhered titanium
was brushed away by the fibers in CFRP when drilling the composites, which accelerated tool
wear in drilling of stacks. PCD drills showed more wear resistance than carbide drills only in
terms of flank wear. However, due to the inherent brittleness of PCD drills, chipping was
observed at the tool edge when drilling titanium. As expected, higher cutting speeds caused
higher torque and thrust force, which resulted in a significant increase in tool wear, especially in
drilling of titanium.

151

Finally, for a better understanding of Minimum Quantity Lubrication (MQL) and its
effective use in practical industrial applications, MQL parameters such as droplet sizes, the
droplet distributions and wetting angles of various lubricants including the nano-graphene (xGnP)
enhanced vegetable oil were investigated. The measurement method of droplet size and
distribution was proposed using CLSM and wavelet transform. In addition, the empirical droplet
size estimation equation was also introduced for extremely small droplets, which is hard to be
measured. The distribution of the droplets was also studied to determine the MQL optimal
nozzle-workpiece distance and the nozzle discharge pressure using the edge detection algorithm.
To extent the applicability of MQL to more aggressive machining conditions, a potential additive
to MQL lubricant, which is a mixture of exfoliated nano-graphene particles and vegetable oil,
was developed. After the wetting angle and friction tests for the various lubricants, it was found
that oil based lubricant, especially nano-enhanced vegetable oil, showed a better wettablility and
tribological behavior. MQL-ball milling tests with nano-graphene enhanced lubricant were
performed to show a remarkable performance improvement in reducing both central wear and
flank wear as well as edge chipping. In addition, it was noted that higher concentration of nanographene did not provided any additional benefit on both central and flank wear.

152

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