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This is to certify that the

thesis entitled

PREDICTIVE TOOL WEAR OF COATED TOOL SYSTEMS

presented by

Raja Krishnan Kountanya

has been accepted towards fulfillment
of the requirements for

M. S . . Mechanics
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PREDICTIVE TOOL WEAR OF COATED TOOL SYSTEMS

By
Raja Krishnan Kountanya

A THESIS
Submitted to Michigan State University
in partial fulfillment of the requirements

for the degree of

MASTER OF SCIENCE

Department of Materials Sciences and Mechanics

1998

ABSTRACT
PREDICTIVE TOOL WEAR OF COATED TOOL SYSTEMS
By
Raja K.Kountanya

Tool wear has been a serious concern in the economics of modern machining. In
the light of the developments in materials technology, the demand for suitable cutting
tool materials has entailed a detailed study into the various mechanisms that bring about

tool wear.

Temperature has been known to play an important role in the mechanisms of wear
of cutting tools. The literature is rich in studies that focus purely either on the temperature
fields or the mechanisms that bring about tool wear. This study adopts a dual approach to
the problem by empirical quantification of cutting temperatures and semi-analytical
modeling of tool wear so as to eliminate the need for laborious testing for optimal tool
materials. Abrasion [Rabinowicz, 1961] and chemical dissolution [Kramer and Suh,
1980] are understood to dominate the wear process of the tool for the work materials
studied. Experimentally obtained wear data and that predicted theoretically, brought to

light many interesting aspects in the tool wear problem.

Cutting tests were conducted on plain carbon steels of A181 designation 1018,
1045, 1065, 1070 and 1095 with carbide cutting tools with a single coating of TiN, TiCN
and A1203. Temperature of the cutting tool was measured using an infrared pyrometer
with a fiber optic attachment. An inverse estimation was then carried out to estimate the
interface temperatures. Flank wear rate increased with cementite content. Crater wear rate
increased with temperature attesting to the common notion that thermally activated wear
mechanisms brings about crater wear. In general, good correlation between experiment

and theory was found.

ACKNOWLEDGEMENTS

I wish to place on record my sincere thanks to my advisor Dr.Patrick Youngseon
Kwon, for rendering all the technical and moral support during the course of this work. I
am deeply indebted to his enduring patience and understanding, which was invaluable

towards bringing this study to fruition.

My thanks to Thorsten Schiemann, from Aachen, Germany, whose insightful
suggestions and practical knowledge clarified many a doubt, which arose during the
experiments and the data processing. I consider my association with him as a delightful
experience, which taught me the essence of teamwork. I am indebted to Ben Simkin, a
fellow graduate student for his valuable time in obtaining the images on the Scanning

electron microscope.

My thanks to my thesis committee members, Dr.Thomas Bieler and

Dr.K.N.Subramaniam for their thoughtful suggestions in the writing of this thesis.

My thanks to Mr.Fritz Smydra of Lansing Community College for his expertise
on the CNC lathe. His availability on some crucial weekends is deeply appreciated. My
thanks also to Kurt Niemeyer and Mr.J.C.Brenton of the machine shop in the College of

Engineering, M.S.U. for 'lending me a hand ' on many an occasion.

Last but not the least, my thanks to all my friends in and around campus,
including Biman Ghosh, Sunil Chandran, Suhail Ansari, Monica Lokre and Sridhar
Seshagiri whose encouragement was the 'magic tonic' to keep me going. The moral I
gained from the whole experience was the simple truth in the sanskrit saying

‘Siddhirbhavati Karmaja’, meaning ‘Do Unto Perfection’.

iii

TABLE OF CONTENTS

LIST OF TABLES
LIST OF FIGURES
LIST OF SYMBOLS

OVERVIEW

CHAPTER 1

INTRODUCTION TO TOOL WEAR

1.1 Geometrical aspects of tool wear

1.2 Literature Survey in relation to tool wear

1.3 Importance of Temperature in relation to too] wear

CHAPTER 2

STATEMENT OF THE PROBLEM

2.1 Rabinowicz Three-body and Two-body wear models
2.2 Chemical Dissolution wear of Kramer and Sub

2.3 Comprehensive models for Crater and Flank Wear

2.4 Accounting the variation of hardness with temperature

2.5 Comparison with Taylor’s model for tool wear

iv

vii

viii

\O-bNN

1o
10
12
14
15
16

CHAPTER 3

CUTTING TOOL TEMPERATURES

3.1 Inverse Problem of measuring the interface temperatures
3.2 1D Ellipsoidal model of Yen and Wright

3.3 Oxley’s method of obtaining the flank temperature

CHAPTER 4

EXPERIMENTAL WORK

4.1 Turning Experiments

4.1.1 Cutting Conditions and related issues

4.2 Method of temperature measurement

4.2.1 Principles of Infrared Pyrometry

4.2.2 Implementation of the Infrared Pyrometer for the Experiments
4.3 Data Acquisition

4.4 Work Materials

4.4.] Dimensions of the bar-stock, cutting length and duration of cuts
4.4.2 Spherodize-annealing of the steels

4.4.3 Microstructures, Composition and Hardness of the steels

4.5 Inserts used in the experiments

4.5.1 Insert Geometry and ISO/AN SI designation

4.5.2 Grade of substrate used

4.5.3 Coating Materials and related details

4.5.4 Designation for the identification of the inserts

4.6 Area Measurement

4.7 Tool Wear Measurement : set-up for Flank Wear and Crater Wear

CHAPTER 5

RESULTS AND DISCUSSIONS

5.1 Results of the Computer Algorithm
5.1.1 Abrasive Wear

17
18
18
21

22
22
23
23
24
24
27
28
28
28
29
31
31
33
34
34
38
38

40
4o
42

5.1.2 Dissolution Wear

5.2 Temperature trends with cutting speed
5.3 Contact Area Trends

5.4 Flank Wear

5.5 Crater Wear

5.6 Calibration

CHAPTER 6
SUMMARY AND CONCLUSIONS

APPENDICES

BIBLIOGRAPHY

vi

42
44
45
47
52
54

57

59

91

Table 4.1:

Table 4.2:

Table 4.3:

Table 4.4:

Table 4.5:

Table 4.6:

Table 4.7:

Table 5.1:

Table 5.2: Curve fitted relations for the dissolution wear of the coating materials
(x is temperature in °K, w.r. stands for wear per unit time)

Table A]:
Table A2:
Table A3:
Table A4:
Table A5:

Table A6:

LIST OF TABLES

Composition of the hot-rolled steels used in the testing
(All in wt%)

Composition of the spherodized steels used in the testing
(All in wt%)

Cutting tool signature

Properties of the substrate

Therrnophysical property data of substrates

Process parameters for the spherodized steels testing
Process parameters for the unspherodized steels testing

List of wear rate equations for the calibration purpose
(T is in °C, w.r. stands for wear per sliding distance)

mechwear.out for the TiN coating
mechwear.out for the TiCN coating
mechwear.out for the A1203 coating
chemdiss.out for the TiN coating
chemdiss.out for the TiCN coating

chemdiss.out for the A1203 coating

vii

29

29

32

33

33

36

37

42

73

73

73

74

74

74

Figure 1.1:
Figure 2.1:

Figure 2.2:

Figure 2.3:

Figure 3.1:

Figure 3.2:

Figure 4.1:

Figure 4.2:
Figure 4.3:
Figure 4.4:
Figure 4.5:
Figure 4.6:

Figure 4.7:

Figure 4.8:

Figure 4.9:

LIST OF FIGURES

Zones of damage of the cutting tool due to machining
Illustration of two-body and three-body abrasion

Oversimplified model of a conical abrasive wearing a
bearing surface to illustrate abrasive action

Schematic illustrating chip flow and continuity conditions
[Kramer and Suh, 1980]

Sources of heat generation in a cutting tool

Temperature distribution in a cutting tool for the
inverse temperature estimation

Photograph of the lathe with the experimental set-up
(Inset: pyrometer end probe)

Illustration of the chip-breaker set-up over the insert

Final assembly of the pyrometer over the insert
Microstructures of the steels used in the testing

Hardness of the steels used in the machining tests

Illustration of the various geometric details of the cutting tool

Sample-naming of the TiN insert used on
A181 1045 (spherodized) for the low speed cut.

Illustration of the method for the area calibration

Set-up on the TMM for measuring flank and crater wear

Figure 4.10: Sample of the crater profiles (Cut-ID xoa1018sp-2)

viii

10

12

13

17

18

22

26

27

30

31

32

35

38

39

39

Figure 5.1: Charts illustrating the ranking of the coatings for the
two abrasive wear mechanisms

Figure 5.2: Figure illustrating the ranking of the coatings for the
dissolution wear mechanism

Figure 5.3: Temperature records of cuts on AISI 1095 steel with
TiN coated K420 carbide (Curve fitting shown alongside)

Figure 5.4: Photographs of K420 carbide coated with TiN, TiCN
and A1203 respectively afier machining spherodized

AISI 1045 steel at increasing cutting speeds at
f = 0.356 mm/rev and doc. = 1.905 mm.

Figure 5.5: SEM micrographs (800x) of the worn flank surfaces of
the coated tools after machining spherodized AISI 1045
steel for low and high speeds.

Figure 5.6: Plots showing the variation of the volume flank wear per
sliding distance for the three coating materials.
(Sp-spherodized steel, un-unspherodized steel)

Figure 5.7: Plots showing the variation of the crater wear rate for
the three coating materials.
(Sp-spherodized steel, un-unspherodized steel)

Figure 5.8: Plots of the predicted and the experimental wear data for
A1203 while machining spherodized steels.

Figure 5.9: Plots of the constants obtained from the calibration process.

(sp — spherodized steel, un — unspherodized steel)

ix

41

43

44

46

49

51

53

55

56

KT

VB

VBmax

V,Vm

P1

is

manna

LIST OF SYMBOLS

Crater width (um)
Crater depth (um)
Crater center distance (um)
Crater area (mmz)
Average wear land width (um)
Maximum wear land width (um)
Nose radius of insert (mm)
Volume of worn material
Roughness angle of the abrasive (deg)
Sliding distance (m)
Normal force of interaction between the surfaces (N)
Tool material hardness (Kg/mm2)
Abrasive Hardness (Kg/mmz)
Calibration constant for abrasive wear per sliding distance
Calibration constant for dissolution wear
Solubility (dimensionless)
Molar volume (cm3/mol)

Gibbs’ free energy (Kcal/mol)

>6

x,y, 2

B:

Excess free energy of solution (Kcal/mol)
Universal gas constant (kcal/mol.°K)
Stoichiometric coefficients

Volume worn away on the relief face (mm’)
Flank wear land (mm)

Constant for the softening (/°C)

Hardness (Kg/mm2)

Temperature (°C)

Dimensionless temperature

Distance in elliptical coordinates (mm’)
Parameters describing the base ellipse (mm)
Complete elliptical integral of the first kind
Steady-State cutting interface temperature (°C)
Temperature in the far-field of the tool (Taken to be 25°C)

Molar Volume (cm3/mol)

xi

OVERVIEW

Metal cutting has been a field of study of immense commercial importance. Since
even minor improvements in productivity can result in enormous cost savings and profits
in high volume production, there has been a never-ending demand for research in this
field even after ten decades of industrialization. In the light of the enormous amount of
work done in the past and the work going on presently, the scope of research has become
very focussed. Hence, study of any problem in metal removal processes involves careful
design of experiments, instrumentation and evaluation of results. A solid infrastructure is

therefore necessary for in this field.

Development of cutting tool materials has been foremost in all the research in
related to metal cutting. A significant stride in this regard has been the innovation of
adopting certain coating technologies for depositing thin hard coatings on common tool
substrates. This technique, developed in the early 19705, dramatically improved tool life
and productivity. Today more than 75% of turning operations and 40% of milling and
drilling operations are performed on coated carbides. While hot-hardness and chemical
inertness have been recognized to be the two important parameters for coating materials,

the exact dependencies and the rationalization are yet to take a concrete shape.

Chapter 1

INTRODUCTION TO TOOL WEAR

Cutting edge wear is still one of the unsolved problems in metal cutting. No matter
how superior the tool material is, it is common experience that tool life, defined as the
time for which tool wear is within acceptable limits, is always finite. Consequently, the
reasons for interest in tool wear are threefold: (1) Lower workpiece quality due to
deterioration in finish and dimensional accuracy, or damage to expensive workpieces if
an edge fails catastrophically in a cut. The second is (2) the cost of changing cutting
edges, the cost associated with the concomitant time delays and damage caused by
unexpected failure and (3) the increased power consumption due to the excessive rubbing
of the tool with the work material. Work continues for three basic purposes: (1) how to
accurately predict wear, (2) how to detect wear from measurements while machining, and
(3) how to minimize wear by the development of new materials. This study was focussed

on the first purpose.

1.1 Geometrical Aspects of Tool Wear

In order to characterize tool wear the proper geometric parameters have to be
defined. The ISO standard 3685-1977(E) [1] for tool life testing was formulated with this
express purpose so that there can be a basis for comparison between various cutting tool
materials. Figure 1.1 illustrates the geometric parameters associated with tool wear in a

typical right hand turning tool.

Section A—A Tool face

V

crater area KA

 

 

KB
Kl

 

 

 

 

 

 

 

flear notch

 

KB= crater width

Kl=crater centre distance
KT=crater depth

KA=crater area (self defined)
VB=average wear-land width
VBnax=maximum wear-land width
r=radius of cutting edge View on major flank

Figure 1.1: Zones of damage of the cutting tool due to machining

1.2 Literature Survey in relation to tool wear
In the light of the numerous attempts to investigate tool wear, the following section
will detail briefly the important information gathered from the recent literature

concerning this study.

Ramalingarn and Wright [2] performed studies on well-characterized work materials
such as F e-C-silica powder-metallurgy compacts where they found flank wear to increase
with silica content. They concluded that alloy chemistry did not describe machinability
and found the wear process on the tool flank to be dependent on interfacial temperatures.
The prowing action of the abrasive particles was clearly shown by means of quick-stop
sections. They also concluded that while the tool material could soften considerably at the
temperature prevalent at the interfaces, the abrasive particles did not soften by the same
amount since the duration for which they were present in the shear zone was extremely
small, of the order of milliseconds. Tool chip temperatures were higher for carbide tools
due to the higher speeds used on them. A modest assumption that the flank temperature

was 300-400°C lower than the chip-tool interface temperature was made.

Byrd and Ferguson [3] studied the influence of hard inclusions on flank wear of
carbide tools. A1203 particles were artificially impregnated in 1020 and 4620 steels using
the PM technique. They concluded that higher temperatures encountered in the
machining of steels do not in any way impede the abrasive action of hard inclusions in
the microstructure. However, they also had difficulties in establishing a threshold level in
the degradation of machinability due to hard particles in the microstructure. Vol.% of

hard particles seemed to provide a clue to the machinability.

Brun et a1. [4] performed experiments on 40%vol. SiC aluminum alloy with different
tool materials. Since they were cutting aluminum, cutting temperatures encountered were
expected to be low. They found that tool materials harder than SiC performed much
better than the others. This dependence on hardness represented a complex behavior

wherein abrasion was dependent on temperature at the interfaces and the altered

properties of the cutting tool at this temperature. Kramer [5] investigated the wear
resistance of binary carbide coatings. The WC substrates began deforming thermally
nearly at 1030K and showed a rapid decline in the compressive strength. At higher
cutting temperatures, he concluded that chemical dissolution was the most important

wear mechanism.

Ramalingam and Watson [6] considered the factors responsible for the scatter in tool
life. The role and significance of the oxygen rich non-metallics on tool wear and
machinability was examined. Tool chip interface temperatures of carbide tools were
800°C and above under normal industrial practices. Diffusion was excluded as a
contributing wear mechanism since it was thought that diffusion was a well behaved,
non-equilibrium mechanism that would lead to a deterministic solution for tool life. In
carbide tooling the ‘prowing’ process was believed to give rise to plastic flow in the
carbide at the tool-chip interface. Hence, a plausible reason for the stochasticity was the

variation in distribution of the abrasive particles in the work material.

Kramer and Kwon [7] concluded that tool wear was primarily due to two mechanisms
namely chemical dissolution and abrasion. The dissolution of the tool material in the
moving stream of chip material may be treated as a dilute solid solution formation and
can be modeled as a regular solution. They also concluded that even though wear models
such as the 3-body and 2-body abrasion do not describe the constraints on the abrasive
particle in machining, they can be adopted for modeling abrasive wear. The ferrite matrix
was believed to be quite soft at the temperatures present in the zone of deformation
during cutting, offering minimal restraint to the inclusions. The details of two-body and
three-body abrasion, an important highlight of this study, will be presented later on in this

thesis.

Kramer and Sub [8] developed the dissolution wear model, which will be the topic of
discussion later on. They concluded that solution wear is predominant in the carbide class

of tool materials, which is independent of the diffusion characteristics. In all cases the

tool material was assumed to be dissolving in a-iron. The solubility in y-iron was
excluded due to the commonly observed sluggishness of the 01—97 transition. Essentially,
this study proved that the free energy of formation of the ceramic coating determined the
effectiveness of the coating. The hypothesis was confirmed from the ranking obtained
from the experimental results of crater wear and that obtained from the thermodynamic
calculations. The difference of this model from the formerly believed diffusion theory is

that dissolution is an equilibrium process whereas diffusion is not.

Dearnley [9] cut'various plain carbon steels carbide tools with a single or composite
layers of TiC, TiN and A1203. Cutting temperature was measured using the tool-work
thermocouple method and was further verified using the metallurgical technique.
Uncoated tools showed a larger HAZ (Heat Affected Zone). The temperature difference
between the uncoated tools and the coated tools was found to be less than 150°C.
Although clear evidence of seizure was noted on the rake surface, there were indications
of reduced interfacial contact on the flank. The A1203 coatings and the ceramic inserts
showed the greatest propensity for ridges on the rake face via discrete plastic deformation
giving further evidence to abrasion on the flank since temperatures are lower on the
flank. This was the basis for the conclusion that wear of A1203 was primarily due to a
decohesion mechanism. Moreover, flank wear trends never followed the trend indicated
by the dissolution or diffusion mechanisms indicating that they may not be the rate
controlling mechanisms, i.e. the mechanisms contributing the most to tool material
removal, on the flank. However, they were useful in interpreting the preferential

dissolution of the WC substrate.

Cho and Komvopoulos [10] studied wear mechanisms of multi-layered coated tools.
Severe abrasion was noted at the flank because of the lower temperature than the crater,
the more rigid work material and the constraint of the moving work and the tool. In the
case of A1203, they noted that dissolution may be neglected at all cutting speeds and
mechanisms such as plastic flow, thermo-mechanical fatigue and fracture were expected

to prevail at all temperatures.

Cook [11] observed that the average wear land temperature seemed to approach the
tool-chip interface temperature as tool wear progresses. At higher speeds, crater wear
rates were primarily a function of the temperature. Functionality of flank wear with
temperature was not deducible precisely. Stj emberg and Thelin [12] noted that increasing
the coating thickness increased the overall resistance to crater wear. The time needed to
expose the substrate underneath was the determining factor in this regard. It was
suggested that the temperatures in the flank are 300°C lower than that at the chip-tool
interface. All coatings were harder and more ductile at lower temperatures. Notch wear

was noted to be a chemical phenomenon.

Kramer [13] suggested that at moderate cutting temperatures, excessive rubbing
occurs between the flank and the work material. F lank wear determined tool life at low
speeds due to mechanically activated wear caused by microfracture, thermal and
mechanical fatigue and abrasion by hard inclusions. Milovic et a1. [14] noted that the
fluctuating stress conditions that can exist within the BUE in the machining of free
cutting steels can be the reason for the superior performance of HSS tools as against
carbide tools, owing to their superior toughness. The coating reduced the interface
temperature by as much as 125°C and hence could be used to turn the material at a speed
higher by 25 m/min. The thermal conductivity of carbide tools, which is higher than that
for HSS, was another reason for the higher heat abstraction from the interfaces and hence

higher speeds needed for machining without a BUE.

Chubb and Billingham [15] studied the wear mechanisms in high speed machining.
They found that once the coating was removed from the flank, the mechanism of wear
appeared to be a combination of abrasion and diffusion. WC appeared to wear by
diffusion as evidenced by the smooth boundary between the WC particles and the steel.
Hence tool life due to flank wear was closely related to the breakdown of the coating on
the flank. Kim [16] noted that abrasive wear dominated the wear on the flank surface and
diffusion on the crater surface. He also concluded that excessive coating thickness could

retard wear performance.

Kim and Durham [17] noted that cutting temperature as high as 1600°C could be
reached with alumina tools without tool failure. The temperature at the flank could be
100°C lower. Tools with a higher thermal conductivity and a higher hot hardness showed
a higher resistance to flank wear. Lee and Richman [18] noted that coated tools resist
cratering even after the coating had been removed in some places. Hardness of a material
as a coating was noted to be very different from that as a bulk material. Cooling of CVD
coatings after deposition at 1000°C was noted to develop tensile residual stresses in them.
Hardness measurements on coatings were particularly difficult because of the smooth

surfaces that are demanded for accurate measurements.

Suh [19], in his classic paper, outlined the essential ingredients for the making of high
performance coatings. Both mechanically and chemically activated wear processes were
pronounced to depend sensitively on temperature. In general, though tools should be at
least 4-4 1/2 times harder than the work material, this is not applicable to coatings since
crater wear rates and hardness of the coatings did not correlate very well. Residual
stresses existed in the coatings and were a function of the CTE differences between the
coating and the substrate. An important point made in connection with this study is that
contrary to A1203, SiC or Si02, Fe3C has a very high free energy of formation and is
likely to dissociate at high temperatures. There has however, been no experimental
evidence to date in this regard. He also mentions that A1203 coatings may have problems
adhering to the substrate since A1203 did not permit diffusion of carbon atoms across the
interface which relates to the common experience of depositing an intermediate layer of

an adherent material like TiN, as was the case in this study.

Subrarnaniam et a1. [20] performed high speed machining on AISI 1045 steel. They
noted that crater area increased with cutting speed. Evidence of twinned martensite in the
chips quenched in water were in accord with the average interface temperature
predictions for the secondary shear zone since steel undergoes a martensitic

transformation at these temperatures. The thermodynamic potential for dissolution is the

most important criterion for the design of a coating to minimize crater wear at high

speeds.

Trent [21,22,23] is one of the pioneers in recognizing the exact interfacial conditions
on the rake face. He was also the first to perceive the importance of recognizing the
mechanisms that control tool wear. In his 1963 series of papers, he was the first to depart
from the then accepted notion of the interfacial conditions. It was then believed that there
was relative motion between the chip and the tool on the rake face. It has now been
established beyond doubt that chip flow resembles fluid flow with an initial region where
there is relative movement and thereafter there is complete seizure thereby forming the
'secondary' shear zone. Among the other pioneers in the field are G.Boothroyd,
P.L.B.0xley, 0.0ptiz and M.C.Shaw. Their innumerable contributions to the literature

have enabled a very scientific understanding of metal cutting today.

1.3 Importance of Temperature in relation to tool wear

From the above section the importance of cutting tool temperatures in tool wear can
be appreciated. Not only does a higher temperature bring about softening of the tool but it
also makes the cutting tool vulnerable to thermally-activated mechanisms such as
dissolution and diffusion. It can also be seen from the literature survey that only a few
attempts have been made at a thorough quantitative study of tool wear involving
temperatures. The present study will create precedence for one in the future so as to

enable a very scientific and rational framework for predictive tool wear.

Chapter 2

STATEMENT OF THE PROBLEM

The principal aim of this study is, as mentioned before, is a quantitative study of tool
wear involving cutting tool temperatures. Among the various mechanisms, which have
been presented in the literature, two main mechanisms namely, abrasion and chemical
dissolution are selected and formulated for the purpose. A calibration scheme is then

proposed wherein tool wear can be predicted for a given cutting temperature.

2.1 Rabinowicz Three-Body and Two-body Abrasive Wear Models

Among the various abrasive wear models in tribology literature, the one appropriate
to modeling abrasive tool wear is the three-body abrasive wear model. Rabinowicz et al.
[24] performed experiments wherein two surfaces slid against each other with the
abrasives introduced in-between [Fig.2.1]. They drew conclusions related to wear rates,
sliding conditions and material hardness. A number of materials and abrasives were
chosen and a general empirical relation was found to fit the data. The final form of the
equations, as applicable to tool wear, is as shown in equations 2.1.

z-aouyAimon J—Mylhmsim

     
   

 

 

Q

Figure 2.1: Illustration of two-body and three-body abrasion.

 

ll

thanB

 

 

 

V” = , fl" < 0.8
31: I:
~25)
= xL “116(5) , 1.25 > 5- > 0.8 Equations. 2.1
"' 5.3Pt If, I;
= than19[fL] , 5‘ > 1.25
"’ 2.43Pt If, P

a

where tan 6? is the average tangent of roughness angle of the abrasive grains (3 measure
of the particle shape or sharpness), x is a sliding distance, L is the normal force of
interaction between the surfaces, 17 is the hardness of tool and If, is the hardness of the
abrasive. Eqs. 2.1 calculate the abrasive wear volume as a function of a sliding distance,

x.

In theory, the equations always governed the volume of material removed, unless the
process of abrasion was preceded by the formation of cracks. In the presence of cracks,
material is removed by brittle fracture as well as by abrasive wear. Since the tool wear
pattern usually seen on the flank, which consists of characteristic groove marks,
resembles a plastic ploughing process, the empirical relations can be expected to describe

three-body abrasion in tool wear also.

While three-body conditions exist when two bodies slide against each other with the
simultaneous rolling of a hard abrasive particle in-between, two-body conditions
represent a hard surface sliding over a relatively soft surface. The relations for two-body
wear are relatively simpler. The volume removed per sliding distance is expressed in

equation 2.2.

d_V_Ltan6
dl_ 72p

 

Equation 2.2

I

where L = Load between interacting surfaces

12

0 = angle for indentation of the conical abrasive particles

p, = hardness of the abraded surface, here the tool flank.
It should be noted that, in this model, the hardness of the abrasive does not appear in
the relation for the wear per sliding distance. Also the hardness of the abrasive is assumed

to be infinite. Figure 2.2 shows the schematic used to obtain the relation.

V p ‘ fig-‘P—it’: V \\\ ///
I / ~ oume \ \ \\\\\\\V //Boarin¢
4 d! >

 

 

Figure 2.2 Oversimplified model of a conical abrasive wearing a bearing surface to illustrate abrasive action

2.2 Chemical Dissolution Wear of Kramer and Sub

Kramer and Sub [8] treated crater wear as an equilibrium process of dissolution and
that a thermodynamic potential existed for the solution of the tool material in the chip
material. They obtained excellent correlation between the wear rates predicted by theory
and that from experiment. This is in opposition to the thermal process of diffusion, which
was then widely believed to be the cause of crater wear, since diffusion is a non-
equilibrium rate process. Diffusion kinetics [8] were proved to be relatively slow at
normal cutting temperatures. Figure 2.3 shows the schematic used to describe the

phenomenon.

The solution wear rate, in terms of the solubility and other parameters, is given by

equation 2.3.

Chemical Dissolution Wear Rate = [BMCVO'S] Equation 2.3

13

where M is the molar volume, C is chemical solubility of the tool material in the work
material, V is the cutting velocity and B is a calibration constant. The 0.5 power in the

velocity term comes from the Schmidt number in mass transfer [7].

   

 

3 S E V"
"f-CONTACT
LENGTH

 

 

      
 

 

Figure 2.3: Schematic illustrating chip flow and continuity conditions [Kramer and Suh, 1980]

For a compound AxByCz, the solubility is given by equation 2.4.

 

 

 

 

AG — AG" - AG" - AG" -RT 1 + l + 1n .
CA 8, =EXpH 4"” x A y 3 z C (x nx y try 2 2)]] Equation 2.4

 

(x + y + z)RT
where
A043}; = Free energy of formation of the tool material
AG ‘" A = Excess free energy of solution of component A in the chip

 

AG” ,3 = Excess free energy of solution of component B in the chip

14

 

AG” (r = Excess free energy of solution of component C in the chip
R = Universal gas constant
T = Absolute temperature in Kelvin

Notes
1. a-Iron is assumed to be the phase into which the tool material dissolves even though
the transformation to the y-phase occurs at 996K.

2. Regular solution behavior is assumed.

Because A and B cannot be determined directly without experiments, the previously
discussed models can predict only relative wear rates where relative wear rate is defined
as a ratio of wear rates between a candidate coating material and a reference coating

material.

2.3 Comprehensive Models for Crater and Flank Wear

Many workers [10,15] have enunciated flank wear to be a case of pure abrasion.
Hence flank wear can be modeled as a case of abrasion alone [equations 2.1, 2.2] and
hence wear needs to be calculated as a wear volume for this purpose. Furthermore, It is
widely believed [11,16] that flank wear rate is best expressed on a sliding distance basis

in view of the fact that unequal cutting times can give a distorted picture of flank wear.

It can be seen in Shaw [25] that

b 2 tanB
, = %— Equation 2.5
Where B, = Volume worn away on the relief face
w = Wear land on the tool
0 = Relief angle of the tool.

Hence the wear volume on the flank is proportional to the square of the wear land.

Since the Rabinowicz model calculates wear on a sliding distance basis, the appropriate

15

[25] formula for calculating the abrasive wear rate on the flank derived from equations

2.1 and 2.2 is shown in equation 2.6.

Flank volume wear per Sliding Distance = AK (Pam) )/(Pr") (3-body abrasion)
= A(l/Pt) (2-body abrasion)

Equations 2.6

where K is the factor from Equations 2.1. Experimentally obtained wear land values have
to be expressed as (VB)2/Sliding-Distancel [Figure 1.1] to correspond to the abrasive

wear calculation and A is a constant to be obtained from experiments.

For modeling crater wear, one has to account for both mechanisms [13]. Wear rate
has to be expressed on a time basis since the dissolution wear rate is not expressible on a

sliding distance basis. Therefore [7,8]

Crater Wear Rate = [AVK (P,‘"'” )/(P,“) + BMV0'5 chgycz ]. Equation 2.7

It is therein assumed that the abrasive and dissolution volume wear rates are directly
proportional to the crater depth and hence experimental wear rate is simply KT/cutting

time [Figure 1.1].

2.4 Accounting the variation of hardness with temperature

Most importantly, one has to consider the thermal softening of . the tool and the
abrasive while machining. A suitable relation has to be evolved to account for the
softening of the materials concerned. Kramer and Kwon [7] used an exponential function

of the form of equation 2.8,

H(T) = Hoe’” Equation 2.8

 

' Sliding-distance (m) in a turning operation is cutting velocity (m/min) x cutting time (mins)

16

where H(T) was the hardness of a material at a given temperature 7°C. H0 and a are
constants obtained form a curve fitting process on the empirical data. In the present
analysis, the same form is adopted along and developed along with the other
thermophysical properties in the computer algorithm to be described later on.

2.5 Comparison with Taylor’s model for tool wear

F.W.Taylor was perhaps the first person to recognize the importance of cutting tool
materials to modern civilization. Often proclaimed as the co-inventor of high-speed steel,
his tool life criterion has been in use ever since he came up with it. Stated simply as in
equation 2.9, it relates various process parameters and the tool life in an heavily empirical

form.

W"a"’ f ' = C Equation 2.9

where V is the cutting velocity, T is the tool life, a is the depth of cut, f is the feed and 1,
m, n and C are constants obtained from a series of machining tests performed as specified
in the ISO standard 3685-1977(E) [1].

However, simple as it may be, from the standpoint of adoption in the industry, it
cannot be sustained indefinitely since a large number of work materials such as metal
matrix composites and multi-layered coated tool materials, with a equally enormous
number of geometries, are now becoming available. Not only does tool life has to be
known apriori for given process parameters, the predictability of the machining process is
equally important because of the close tolerances and stipulations being made nowadays

on the surface finish of the component.

Chapter 3

CUTTING TOOL TEMPERATURES

Temperatures in metal removal processes have long been of interest to many
researchers. There are essentially three temperatures to be concerned with in the turning
process [Figure 3.1], (1) the shear plane temperature, which represents the bulk of heat
generation in the chip formation process, (2) the chip-tool interface temperature, which is
influenced by a number of factors including the secondary shear zone and (3) the flank
temperature or the work-tool interface temperature, which is considerably lower than the
other two. Flank temperature is important for its influence on the wear mechanisms
prevalent at the flank in addition to its minor contribution to the heat generated in the
chip-formation process. This chapter elaborates on the theory for the temperature

measurement in the experiments conducted.

   
 
   
 

 

chip-tool interface

workpiece work-tool interface

 

shear plane

Figure 3.1: Sources of heat generation in a cutting tool

17

18

3.1 Inverse Problem of measuring interface temperatures

Cutting tool temperatures are usually measured using a chip-tool thermocouple
technique [11]. This technique essentially gives the area weighted mean of the
temperatures at the chip-tool interface, the work-tool interface and the shear plane. Other
models developed recently such as the 1D ellipsoidal model of Yen and Wright [26]
avoid the problem of inaccessibility of the interfaces and determine the cutting tool
temperature as a whole field measurement and thus enable inverse estimation of interface
temperatures. These involve measuring the cutting tool temperature away from the

interface and adopting an inverse estimation of the chip-tool interface temperature.

3.2 1D ellipsoidal model of Yen and Wright[26]

The model essentially assumes that a one-eighth ellipsoid [Figure 3.2] in the domain
of the cutting tool represents an isothermal surface. The cutting tool is therein assumed to
be a semi-infinite body. Heat generated at the shear plane and the chip-tool interface
manifest themselves as heat input into the cutting tool in a defined elliptical area of
contact on the rake face. This heat input is represented as a constant temperature in this

elliptical area and used as the boundary condition for the heat diffusion equation.

      
 

rake face

chip-tool contact area (base ellipse)

“an... (>0

 

 

.
,o
.9"!
a
I“
......
I ,-
........
0" “a!
I. .........
''''''''''
..........
. ..
“v“.f ...........
.....
................
........

 

Figure 3.2: Temperature distribution in a cutting tool for the inverse temperature estimation

19

For the steady state temperature distribution, one has to solve the Laplace equation
in ellipsoidal coordinates. However, it can be simplified for the special case of the one-
dimensional (l-D) steady-state problem, where the temperature distribution in the tool
body is a function 0, only of the radial coordinate g as in equation 3.1[26].

 

 

d d0 .
E£Rg 61—5) = 0 Equation 3.1
. T; -T..
wrth R¢=\/(a2+§)-(b2+§)-§ and o=T T
R — co

where r, is the radial coordinate in the 1-D ellipsoidal model [mm2]
a,b are the parameters describing the base ellipse with a>b [mm]
0 is the relative steady state temperature
TR is the steady-state chip-tool interface temperature [°C]
Tg is the temperature at the location determined by 5, [°C], and

To. is the ambient temperature [°C].

The boundary conditions specified are: -

All other faces are insulated (heat convection = 0 or negligible).
Therrnophysical properties of the tool material are constant.
The tool is rigid and tool wear is negligible.

Uniform temperature TR at the chip-tool interface.

9993”?”

Temperature at infinity or the far field of the tool is T...

With some mathematical manipulation, a function for 0 can be found as in equation 3.2.

a .5 l
2K0") [Mkaz +x)-(b2 +x)-x

2
iFESin-{J— 72% '32—]
=1+ [([02 422]
02

 

 

dx

 

@(4‘)=1—

 

 

Equation 3.2

 

20

 

where F (1; | or) is the complete elliptical integral of order or.
K(m) is the complete elliptic integral of the first kind, and

x is a variable of integration.

The final form of the solution for the one-parameter case (a = b) appears as: -

Equation 3.3

 

where
T, is the steady-state cutting interface temperature
T is the steady-state remotely measured rake face temperature
T... is the temperature in the far field of the tool (Taken to be 25°C)
x is the distance of the point of measurement from the origin of the axes, and

a is the radius of the circular tool-chip contact area.

While the model can be developed more generally for three dimensions, closed form
solutions exist only for the 1D case with one, two or three parameters. The one-parameter
scheme is extremely efficient and facilitates accurate computation of the interface
temperature. Hence from knowledge of the temperature at a point located remote to the
interface, i.e. at a distance x from the interface, the mean interface temperature at the
interface can be estimated using equation 3.3. This is the form for the inverse estimation

which was used in this study.

The inverse estimation scheme is however sensitive to measurement error, Lin et al.
[27], while implementing it for the infrared method, concluded that the inverse

predictions for the heat conduction problem is an ill-posed problem and instabilities are

21

likely. However, they established that temperature measured in the IR system was in
agreement with that predicted by the metallurgical approach at low cutting speeds
without chip control. They had also used a square tool in their experiments. Hence small
deviations from theory are likely in using the same scheme for tools with a clearance
angle, as in this study. The cutting speeds for which the effectiveness of the scheme was

tested was also comparatively higher than that used in this study.

In a similar approach, Rall and Geidt [28] concluded that most of the energy in the
the metal cutting process appears as heat that is transferred to the chip, even though the

small proportion of heat transferred to the tool influenced tool life.

3.3 Oxley's method of obtaining the flank temperature

The temperature difference between the chip-tool interface and the flank has been a
much debated topic. In the literature survey presented before, one can see a number of
opinions raised regarding this. In this study, the relation proposed by Oxley [29] was
undertaken. Oxley’s conclusions were drawn from the isotherms of the infrared
photographs taken by Boothroyd in an earlier work [30] where stresses and temperatures
on the flank wear land were studied. Boothroyd’s work also concluded that large
variations with temperature do not occur (<16°C) over the flank wear land and that

temperature was higher away from the cutting edge.

Stated simply, the tool work interface temperature is 0.82-0.95 times the mean chip-
tool interface temperature (All temperatures being in Kelvin). An average value of 0.89
was used for this study. Slight deviations are possible due to the sharp HSS tools used in
the Boothroyd’s tests [30]. Moreover, the higher thermal conductivity of the carbide tools
[12] would tend to make this factor higher than that for HSS. In the absence of an exact
dependency, this nevertheless suffices for this context.

Chapter 4

EXPERIMENTAL WORK

4.1 Turning Experiments

The experimental work was carried out at the premises of the Lansing Community
College in Lansing, MI. The machine used was a Milltronics Manufacturing 20 HP
medium sized lathe with provisions for infinitely variable speed and programmable feed
and depth. It had a rigid tailstock, important for ensuring minimal chatter while turning
long bars. The lathe was facile since the RPM could be controlled to keep the cutting
speed constant. Figure 4.1 shows the photograph of the lathe with the experimental set-

up.

 

Figure 4.1: Photograph of the lathe with the experimental set-up (Inset: pyrometer end probe)

22

23

4.1.1 Cutting conditions and related issues

Dry cutting experiments were performed at a constant feed of 0.3 56 mm/rev
@.Ol4"/rev1 anl depth of 1.905 mm (0.075") while the cutting speeds selected were at
300, 500, 700 and 900 sfpm (between 90 to 275 rn/min). Cutting speeds were chosen
after referring to the insert manufacturer’s recommendations. The feed and depth of cut
were determined after a few trial runs with the inserts to give a consistent cut with
minimal chatter, sparking and good surface finish while maximizing on cutting time to
reach steady state cutting conditions. The machine was also programmed for the RPM

rather than the cutting speed to keep the cutting velocity constant.

A rigorous study of flank wear has to take into account this effect of the “cut-in”
wear, which develops on the flank immediately after the commencement of cutting.
Dearnley [9] had noted that flank wear rates estimated after longer cutting times were
more representative than that after shorter cutting tests. In all cases at least a minute of
cutting time was achieved to supercede the effect of this cut-in wear and ensuring that the
tool reaches a steady state of wear. Exactly one cut was performed for every satisfactory
data point. Data sheets consisting of the wear data and the process parameters were
prepared for every cut and were compiled separately both as a hard copy and on the

computer. The machining data (cutting speed) is shown in tables 4.6 and 4.7.

4.2 Method of Temperature Measurement

As explained before, there are several ways to measure cutting tool temperatures. In
this study, infrared pyrometry was adopted. Conventionally this method is suitable only
for high temperature measurements such as that in ovens and other places. However, new
detector materials have been developed and hence this efficient and quick way of non-
contact temperature measurement has become accepted all over the world. This method
was deemed more reliable than other methods of cutting tool temperature measurement
such as embedding thermocouples because: -
1. Bonding of thermocouples to the cutting tool is always a problem, given the dynamic

nature of the process.

24

2. Thermocouples will tend to thermally ‘load’ the cutting tool and may alter the
temperature patterns on the tool.

3. The pyrometer technique is reputed for the fast response it possesses.

4. Since the volume of cutting tests to be performed was very large, a non-contact

method saves a lot of time in interchanging and indexing the cutting tool.

Coupling of fiber optics to infrared detectors represents the latest progress in the
field of non-contact temperature measurement and control. Previously, fiber optics was
excluded from consideration since they were made of glass and plastics, which are
opaque to infrared radiation. This attachment can be beneficially used in cutting tool
temperature measurement since it allows the probe to travel along with the cutting tool

during the feed.

4.2.1 Principles of Infrared Pyrometry

The underlying principle of this technique is that all matter emits electromagnetic
radiation at a temperature above 0K proportional to the fourth power of the absolute
temperature. A temperature dependent emissivity is the parameter that adds a dimension
to the physics of the method. Though emissivity itself carries several meanings according
to the circumstances, it represents the behavior of a real surface and can be generally
defined as the ratio of the radiation emitted by the surface to the radiation emitted by a
black body at the same temperature.

4.2.2 Implementation of the Infrared Pyrometer for the experiments

The detector head of the pyrometer was designated 0S1513 General Purpose
Sensor used in conjunction with Model 3026 Single Channel Thermal Monitor]. It
possessed a response time of 10-msec and was calibrated for the temperature range of 84-
300°C. The end probe of the fiber optic cable was a glass tipped steel probe of 3" lengg
(Not as shown in the photograph in figure 4.1) and the spot size was found to be 0.785

mm2 prior to the commencement of every cut. Accurate positioning of the probe was

 

' Courtesy OMEGA Vanzetti, Inc.

25

done using a backlight source. There was also a provision for emissivity correction in the
system, though this was not used in this study. The calibration of the pyrometer was
performed using the BB-4 black body calibration source’. The non-linearity was found to

be less than 1%. The absolute accuracy of the pyrometer was found to be 3°C.

Since the detector head was capable of measuring radiant energy in only one
wavelength i.e. a single color pyrometer, it had to be corrected for the emissivity.
Emissivity is always an important consideration in such non-contact temperature
measurements especially while using these techniques for metallic surfaces, which are
usually shiny. A high emissivity is desirable since it eliminates the adulterating effect of
the background radiation. Among the coatings used for the testing, the TiN and TiCN
inserts had a very shiny surface. In a concurrent study that was carried out, the emissivity
of the TiN surface was found to be 0.101, which was dangerously low for temperature

measurement.

To circumvent these problems, a thin coating was black high temperature paint
(Flat Black) was sprayed on the inserts prior to the tests. Proper drying of the paint was
ensured since this could harm the measurement during the experiments. The emissivity of
this paint, which was known to be 0.92 from OMEGA, was the value to which the
monitor always set. It is recommended that in future, the dual color detector head be used

since this can measure temperature independent of emissivity.

 

1 Courtesy OMEGA Inc.

26

BRAIDED CABLE

END Prone.

/

----------- 1 CHIP Baum-ta

 

  

 

 

 

 

 
   

INSERT X

 

 

Clamp _- Chip-Breaker

  
 

<..- P...— __'-'-

Hole for
End Probe

 

Test Insert

 

 

/

Chip-Tool Interface

Figure 4.2: Illustration of the chip-breaker set-up over the insert

27

In order to prevent the interference of the chip in the field of view of the infrared
probe and to exercise some control on the chip, a chip breaker was designed using one of
the same inserts to hold the pyrometer in place. Figure 4.2 shows the schematic of the set-
up shown to illustrate this point. It should be noted that there might be some disturbance
in the temperature field due to this device. But in all cases it was found that it was

sufficiently far away from the interface as shown in figure 4.2. The final assembly is as

shown in figure 4.3.

FmrmOmc
. ’ Arnamarr

 

l . ’ i
l 5 f l Qiamlusra"1'"~ . , p _ / .' \.

 

 

 

 

 

 

 

 

 

Figure 4.3: Final assembly of the pyrometer over the insert

4.3 Data Acquisition
Temperature data was collected in an automated data collection system. This

consisted of a data acquisition board, signal conditioning and software which was written

for the purpose of collecting and storing the data in a spreadsheet. The timer of the

28

computer was used, enabling the on-line monitoring of the temperature. Though initially
the software was written for acquiring data in 5 channels-2 temperatures and 3 cutting
forces, only the channel for the pyrometer was ultimately used. The thermocouple

underneath the insert and the dynamometer were not implemented.

4.4 Work Materials

The work material consisted of steel rounds with the AISI designation 1018, 1045,
1065, 1070 and 1095. Being plain carbon steels, cementite constitutes the bulk of the
inclusions in their microstructures. All the work materials for the experiments were
procured from Alro Steel Corporation, Lansing, MI excepting the AISI 1065 bar, which
was acquired from Timken as a compliment. It was learnt that the AISI 1065 steel was

more commonly used as a plate rather than as a bar stock in the industry.

4.4.1 Dimensions of the bar-stock, cutting length and duration of cuts

The work material used for the experiments were steel rounds commercially
obtained fiom ALRO Steel, Inc., Lansing, MI. The bar stocks were nominally of
diameters between 3” and 6” and length was about 2-1/2' initially. The 1065 bar alone
was 2” in diameter to begin with. However, all the bars had been reduced to a diameter of
about 1.5" at the end of the tests. The cutting length was different for different bars for

stability reasons. As mentioned before, the shortest cutting time was 1 min and 3 sec.

4.4.2 Spherodizing-annealing of the steels

The spherodize-annealing process transforms the cementite in the steel in to
spheroids and brings the steel to a dead-soft condition, thereby removing any shape effect
in the abrasion of the tool while machining. Those that were required to have a
spherodized microstructure (1018, 1045, 1065 and 1095) were sent for heat treatment to
Atmospheric Annealing, Inc., Lansing, MI. The sheer size of the bars excluded any
possibility of annealing in a laboratory. This was also the reason due to which control
over the grain size of the cementite in the steels was not possible. The as-received steels

for the tests (1018, 1045 and 1070) were mainly in a hot-rolled, normalized condition.

29

The details of the annealing process for the AISI 1045 steel is shown below: -
1. 0-1000°F in 6 hrs, then hold for 4 hrs.

.O‘SJ'PP’N

Then cool to room temperature in air.

1000-12000°F in 5 hrs, then hold for 3 hrs.
1200-1310°F in 5 hrs, hold for 50 hrs.
1310-1200°F in 3 hrs.

1200-1150°F in 0.5 hrs.

The process was similar for the other steels.

4.4.3 Microstructures, Composition and Hardness of the steels

compositions of the bars are as shown in table 4.1 and 4.2.

Table 4.1: Composition of the hot-rolled steels used in the testing (All in wt%)

 

C

p

8

SI

1N1

Cr

Mo

Cu

Sn

AI

V

The hardness and the composition of the steels are examined in this section. The

T1

 

1018
1045
1070

 

 

0.208
0.476
0.684

 

0.702
0.744
0.780

 

0.015
0.011
0.013

0.026
0.037
0.024

 

0.212; 0.069

0.273

1
0.215111 0.042

1 0.051

0.133
0.077
0.165

 

0.018
0.015
0.016

 

0.258
0.109
0.046

 

0.011
0.004
0.006

 

0.020
0.037
0.020

 

0.003
0.004
0.000

 

0.001
0.000
0.000

0.004
0.005
0.004

 

 

 

 

Table 4.2: Composition of the spherodized steels used in the testing (All in wt%)

 

C

4-

p

S

 

1018
1045
1065
1095

 

 

0.160
0.476
0.640
0.887

 

 

0.828
0.744
0.800
1 .024

 

0.010
0.011
0.014
0.019

 

’ 0.028
0.037
0.010
0.025

 

31
"0"."1‘03’
0.273
0.280
0.309

0.051
0.070
0.145

 

""6361’4‘

Cr

Mo

Cu

Al

V

‘—

 

Tl

 

”01079
0.077
0.150
0.316

 

 

0.011
0.015
0.020
0.141

 

"0.053
0.109
0.130
0.156

 

0.004
0.004
0.009
0.007

 

0.020
0.037
0.024
0.004

 

0.001
0.004
0.002
0.170

 

0.000
0.000
0.000
0.000

 

0.001
0.005
0.002 .
0.004

 

 

Figure 4.4 shows the photomicrographs of the microstructures. As one can see,

the as-received steels showed a very pearlitic structure. For the spherodized steels, this

lamellar cementite were transformed into spheroids by the spherodizing process. It is also

evident that the cementite in the spherodized steels was not of a uniform size and this

could have a role to play in the tool wear process. In particular, the 1095 steel showed a

 

30

very small cementite size, which was believed to be because it was vanadium killed

[Table 4.2]. Figure 4.5 shows the hardness of the steels in various states.

   

 

1018(as-received) 1045 (as-received) 1070(as-received)
1018 (spherodized microstructure) 1045 (spherodized microstructure)

1065 (spherodized microstructure) 1095 (spherodized microstructure)

Figure 4.4: Microstructures of the steels used in the testing

31
m: moron steels Dem
240 “WWW mmsnnoruumtaoaeu

.BdoreSphnrodiziro 3
. . . 1
7 ‘ ; " uAiterSpharodizim ‘

   
   

Dachau [M
a‘.
o

     

' quuvmvm I

1018 1045 1070
AISI Dealngatlon of Steel

1018 1045 1065 1095
AISI Stool Damnation

Figure 4.5: Hardness of the steels used in the machining tests

4.5 Inserts used in the experiments
The inserts for the experiments were procured on request from Kennametal, Inc.,
Latrobe, PA. The turning tool holder was also procured fi'om them. Adequate number of

spare parts for the tool holder were also purchased for exigencies during the experiment.

4.5.1 Insert Geometry and ISO/ANSI designation

The geometry of the inserts had the ISO designation SPGN 19 04 12. The integral
chip breaker was intentionally avoided (l) to simplify modeling of tool temperatures, (2)
to make an accurate measurement using the pyrometer and (3) to apply the 1-D
ellipsoidal model. A large nose radius was chosen to avoid catastrophic failure of the
insert and a large IC (Inner Circle diameter) was chosen to suitably place the pyrometer
on top of the insert. It was felt that a proper choice of the nose radius is important since
this could bring about chatter problems and difficulties in flank wear measurement if
chosen improperly. It should also be mentioned that the depth of cut should be at least as
high as the nose radius to have a suitable chip curl. There were also many opinions that
the insert style was one of milling than turning. The inserts were used with a standard
CSRPR 856D tool holder. The assembled tool signature of the insert, measured with the
toolrnaker’s microscope is shown in table 4.3. Figure 4.6 illustrates the various geometric

parameters.

32

Table 4-3 Cutting Tool Signature

 

 

 

 

I Parameter Back Rake Side Rake End Relief Side Relief —End Cutting Side Cutting Nose
Angle Angle Angle Angle Edge Angle Edge Angle Radius

Symbol a. at, 0C 0, cc c, r
“We 0° 4°42' 4°42' 0° 15° 15° 3/64"

 

 

 

 

 

 

 

 

 

 

 

 

{\

 

 

 

 

 

 

ISO specification PGN 19 0412

I Lnose radius :12 mm
thickness 5 : 4.76 mm

length i, inner circle 10: 19.05 mm

insert type N; without hole, without chipbreaker
tolerance class G

relief angle P (beta): 11 degree

shape 3: square, nose angle 90 degree

 

 

Figure 4.6: Illustration of the various geometric details of the cutting tool

33

4.5.2 Grade of substrate used

The substrate was a K4202 grade, which consists mainly of WC and a C0 binder
has a slight alloying of TiC and TaC. The American standard for the same is C6-C7. The
details are shown in table 4.4 [31]. Thermophysical properties of the grade collected from
Santlranarn [32] are shown in table 4.5. It has a medium binder content and a large grain
size meant for general purpose steel machining. It is also renowned for the right balance

of wear resistance and toughness.

Table 4.4: Properties of the substrate

 

 

 

 

 

 

 

 

 

 

 

K420
ISO Specification Nominal Specifications
P25-P35 Porosity A04-B00-C00
M25-M30 WC (Gsum) l-8um
Nominal Composition — wt% HRA 91.2
Co 8.5 Hc(0e) 140
TaC 10.9 Dens. (g/cc) 12.65
TiC 7.4 TRS (Mpa) 2170
WC 73.2
Table 4.5: Thermophysical property data of substrates
Composition Grain Hardness CTE —Thermal Density g/mmz TRS
size HRA um/m°C Conductivity MPa
W/m.K

94WC-6Co Fine 92.5-93.1 5.9 108 15.0 1790

94WC-6Co Medium 91.7-92.2 5.4 100 15.0 2000

90WC-10Co Fine 90.7-91.3 6.0 80 14.6 3100

72WC-8TiC- Medium 907-91 .5 6.8 50 12.6 1 720

1 1.5TaC-

8.5Co

 

 

 

 

 

 

 

 

 

 

2 Courtesy Kennametal, Inc.

34

4.5.3 Coating Materials and related details

The TiN coating was commercially available from Kennametal as KC7102. It
consisted of the K420 substrate with a 411m coating of PVD TiN over it. The two other
coatings TiCN and A1203 were custom made for this study as follows. Uncoated inserts
were sent to Balzers, Inc, Lansing, MI, where a 3.5pm coating of TiCN was performed.
The PVD process used for this purpose was Reactive Ion Plating. Uncoated inserts were
also sent to Valenite, Inc., Troy, MI, where a 311m CVD coating of A1203 was performed.
Due to the well-known problems of the A1203 coating adhering to carbide substrate [32],
a 111m intermediary layer of TiN was deposited between the A1203 and the carbide
substrate. The nature of the CVD process used was not disclosed. The hardness data of
the coatings was also collected from the literature and a variety of sources to be
elaborated later on. The TiN coating was a golden colored coating whereas the TiCN had
a bluish gray color and the A1203 coating had a dark black color. Also the TiN and TiCN
coatings were shiny while the A1203 coating was rough and dull. This great difference in
colors and texture persuaded the thin black coating in the temperature measurement as

explained before.

4.5.4 Designation for the identification of the inserts

Given the huge volume of tests that were performed, it was imperative that a
proper designation system for the inserts be evolved. This was done and the same
designation used for the data sheets, the temperature records and the crater photographs
stored on the computer. Figure 4.7 illustrates a sample naming scheme used. Tables 4.6
and 4.7 Show the process parameters of the cuts with reference to this naming scheme. In

all 84 cuts were performed, of which, two were invalid.

 

2 Courtesy Kennametal Inc.

35

xng1045un-li1

I

> Cut Number of the edge

) Edge Number

 

 

p Unspherodized/Spherodized

 

 

 

> AISI steel designation

 

» Insert alphabet or name

 

 

> Coating TiN - n
TiCN - 0
Ale3 - 0

 

, Grade of carbide used x - K420

Figure 4.7: Sample-naming of the TiN insert used on AISI 1045 (spherodized) for the low speed cut.

36

 

 

 

 

 

2...: 1.38.22. .28 4.2333.

3.38 3.8.8.. 3.5 2.3339.

.2: 2.38.22. 83. 2.3332

8.8. 2.3889. 2.8. 33339.
6.2 6.2

3.8” 1.38:3 .28 1.3332.

a... . a «238...: 3.8a 2.3332.

2...: 2.8.3.. 8.3. 2.4332.

3.8. 138:3 8.8. .2332.
2o: 5:

.28 1.38:5. 3.8m 1.33.»:

we. . a 2.38:5. 3.38 2.33.?

8.3. 2.38. .5 8.3. 2.3325.

3.3. 338:5. 8.8. 133.9:

2: 2338. 2; 2332
2.5 1.38.80. «38 «-22262
2.2... 2332

.6 6.2 s... so 3.: 328 $2228..

3...: 3.328.. 3.3. 22228..

322 .2322; 8.8. .2222...
6.2 m2....

2 . .3 4.28.8. 3.48 25:28..

82.3 «.2383: .33 4.2.8.8..

«:2 22328.. 8.3. "2228..

8.8. 138.80. ~22 .2228x
5: 2o:

5.8 1.3825. 8...: 138:5.

as: 13825 Son 133...:

~32 3.32.5. 8.3. 02235.

2.8. 13825. 8.8 .2235.

2:. 2232 2: 2222
.5525 82.5.3 .5525 8.65.3
68%. 9.25 3.an 9.2.8 .3an 0.33 :3... Sam 225 arses. 9.230 3.an new use.

 

wage. £86 3333...... 05 .8 £808.53. 808k— 66 03a...

 

37

 

 

 

 

 

2.35 1.3332.
8.8 1.3332.
38. 8.3332.
man 2 8..
35.8.... .3 a...
6.2
8.38 1.3332.
8.88 2.3332.
8.8. $533.32.
8.... ....3332.
2o:
8.8 $833.55.
8.8 $833.55.
8.3. «5:33.85.
8.8 7533.5...
23 £525.33.
8.nh~ I 4 . ||-.. --.-.. .. . . ..a V'goho_=°x om.th ......yi....................... .. .. . .. . ......C... ........I-I|.Ii...!..........tb..y - uvgw~°_—°x
833 $888.52. 8...: $822.2.
8.8. $558.52. 8.3. $822.2.
8.8 .-558...2. 8.8 75.22....
.03... 5.2
8.8 755832. 838 5822.2.
8...: $558.52. 833 $822.2.
8.8. $558.52. 8.3. $822.2.
8.8 7558.5... 8.8 .8222.
20... 20:
2.58 7888.5... 8.38 5558......
8...: $558.5... 853 $55.25..
8.8. $5.825... 8.3. $822.5.
8.8 7558...... 8.8 7558......
2: 32588. 2: 3.55.8.2
A58)... ASE—5
82.3 3.2.6 .5553 3.58 522.3 .83 8.5.5.3 8...... 82.3 3.5.0 .5853 3.58 3.2.5.. .83 8.5.5.80 82..

 

 

wage. ...—08m Egunams. o... .5. flaogm 8.89... ....v 035...

 

38

4.6 Area Measurement

Photographs of the crater were taken with a LECO SZH stereo microscope with a
LECO 2001 image analyzer. Samples of the photographs taken are shown later in figure
5.4. Measurement of chip-tool contact area, which is a vital input in the inverse
temperature estimation, posed a great challenge. After attempting a number of methods, it
was found that it was most accurately measured using an image processing software.
However, the spatial calibration was a critical issue. Figure 4.7 shows the methodology
used in this regard. The diagonal formed by the ends of the nose radius was the easiest
and the most accurate dimension, available right on the insert specimen. It also avoids
any parallax error, if any. The image processing software also enabled adjustment of the
contrast of the image. This was important since precise demarcation of the chip-tool

contact area was sometimes absent.

 

Nose Radius ——\

__,..--""i .6838 mm

 

 

Insert

Figure 4.8: Illustration of the method for the area calibration

4.7 Tool Wear Measurement : set-up for Flank and Crater Wear

Flank wear was measured using the Mitutoyo TM-505 toolmaker’s Microscope at a
magnification of 200. The microscope was equipped with digimatic heads enabling
measurements accurate up to 1 um. The set-up is shown in figure 4.8. Crater wear >5 um
was measured on the stage and the optics of the microscope at a magnification of 30, with
the Fowler D1040 Digital Test Indicator. The accuracy on this was lum. The calibration

was checked using one-tenth thousandths gage blocks. For very shallow craters, the Sloan

39

Dektak IIA surface profilometer in the premises of the Physics Department of M.S.U.

was used. A sample of the profiles obtained is shown in figure 4.9.

 

Figure 4.1: Set-up on the TMM for measuring flank and crater wear

10:2 SCAN: 3,000 mm VERT: 0 A
17:07 02-26-98 SPEED: HIGH HORIZ: 1,930 um

80,000

75,000

70,000

 

65,000
60,000
0 500 1,500 2,500
A CUR: 01,301 A Q 1,052 um
M CUR: 01,301 A Q 3,988 pm SLOAN DBKTAK 11

Figure 4.2: Sample of crater profiles (Cut-ID xoa1018sp-2)

Chapter 5

RESULTS AND DISCUSSIONS

5.1 Results of the Computer Algorithm

The results of the computer algorithm [Appendix A] yield a general ranking of the
three candidate coatings fi'om the hardness and chemical properties. A coating with a
greater hardness has a better abrasion resistance and a coating with a lower free energy of
formation has a greater resistance to chemical dissolution. Though cementite is softer
than the coating materials at all temperatures, the variation of the abrasive wear-rate with

temperatures, as given by the theoretical formulae in Chapter 2 was more keenly seen.

Interestingly, 3-body abrasion showed a decreasing rate with temperature. This is
justified since even in the work of Rabinowicz [33,24], a rapid fall in wear-rate was
observed with decreasing hardness of the abrasive. Two-body abrasion, however, shows
an increasing rate with temperature, simply because the hardness of the abrasive is
assumed to be infinite at all temperatures and only softening of the tool material is taken
into account. From a practical standpoint, the validity of this assumption is indeed

questionable.

Dissolution, being inherently a thermally activated mechanism, increases steeply
with temperature. The exact methodology of the numerical computation is elaborated in
Appendix A. Figures 5.1 and 5.2 show the output of the computer program for the three

coatings in absolute terms.

40

41

Theoretical 3-Body Abrasive Wear Rate

 

1.E-04 l"
I
i _.°_..TiN
. . .g. . . TICN
1E'06 +Alum'na

  

Volume Wear Rate
(Arbitrary Units)

 

 

 

1 E-OB . El .....
. , a ....... B-
""""" 13
1.E-10 4 . 1
700 900 1 1 00 1 300
Temperature (’C)
Theoretical 2-Body Abrasive Wear Rate
6.E-03 l

 

Volume Wear Rate
(Arbitrary Units)

 

 

700 900 1 100 1300
Temperature (°C)

Figure 5.1: Charts illustrating the ranking of the coatings for the two abrasive wear mechanisms.

42

5.1.1 Abrasive Wear

It can be seen in figures 5.1 and 5.2 that wear-rate decreases with temperature for
the 3-body abrasive wear model whereas it increases with temperature for the 2-body
wear model. A curve fitting of the form Ae‘“T was performed for figure 5.1 for prediction
of wear with the experimental results. The final equations evolved for the 6 cases are

shown in table 5.1.

Table 5.1: List of Wear-rate Equations for the calibration purpose
(T in °C, w.r. stands for wear per sliding distance)

 

 

 

 

Coating Compound Three-body Wear Equation Two-Body Wear Equation
TiN w.r. = 7.5506E-04e'9'3077E’03T :w.r. = 3.91735-04e‘596254’“
TiCN w.r. = 3.629515061-2”0°89“3T Ew.r. = 1.8166E-04e2'°°”E'°3T
i
1
N20: w.r. = 1.374230465751550” §w.r. = 30735150413.“74°“3T
i

 

 

 

5.1.2 Dissolution Wear

As can be seen from figure 5.2, dissolution wear-rate increases with temperature.
Also, A1203 is considerably more inert that the other two coatings. This corresponds to
the common experience of using A1203 coatings for finish machining since higher
temperatures are encountered due to the high cutting speeds employed for finishing. The
curves were then fitted in the form Ax“, where x is the temperature in °K. The reason for
using this form was the steep drop in dissolution wear with decreasing temperature. The

final equations are shown in table 5.2.

43

Table 5.2 Curve fitted relations for the dissolution wear of the coating materials

(x is tern rature in °K, w.r. stands for wear r unit time)
P9 pe

 

Coating Dissolution Wear Equation

 

TiN w.r. = 1.1155815-54x1-6234N1

ncn w.r. = 185755431252"E+O1

N20: w.r. = 8.1920E-96x2'82°15*°‘

 

 

 

 

 

 

 

 

 

 

1.5E-02 F—

1 "C W :3
g l -9— TICN
5 g 1.0E-02 ll A— Alumina <>
2 5 m ‘
a ,
2 5
° 2
8 s.
2
O
>

HS
1300

 

Temperature (°C)

Figure 5.2 Figure illustrating the ranking of the coatings for the dissolution wear mechanism.

44

5.2 Temperature trends with cutting speed

A sample of the temperature profiles for the cuts on spherodized steels, TiN coated
K420 while machining spherodized AISI 1095 steel, is shown in figure 5.3. One can note
an increasing trend with cutting speed. The curve fitting, to obtain the steady state cutting

temperature, is explained in Appendix D is shown in the same graph.

180 _
160. 2m"
-' i
140 '1 « 1 "5" mm
*1

l I nIfl-Qg-W 1W . r :v 1“ I. I”

-
/

. i -c' 1' r _, . '
120 ' " ‘ '
I
.l 1 fl . IJ
: ; ' v
1 v 1 V

 

Temperature (°C)

 

0 20000 40000 60000 80000 100000
Tina (ms)

Figure 5.3 Temperature records of cuts on AISI 1095 steel with TiN coated K420 carbide
(Curve fitting process shown alongside)

Although the pyrometer technique was easy to conceive and implement, the
accuracy in measurement had to be validated with the data obtained in the past. The chip-
tool interface temperature measurements compared favorably with the results obtained by
Subramanian et al. [20] who employed the same work material, similar cutting conditions
and grade of carbide. They had estimated the interface temperatures based on the cutting
forces of Boothroyd’s model [30]. The flank temperatures observed in the present
experiments showed an increasing trend with the cutting speed, except in a few cases
[Appendix C]. Complications due to abnormally large contact areas may have contributed

to the high temperatures observed in these low speed cuts. Another possible reason could

45

be that the peak temperature was far removed from the cutting edge thereby decreasing

the apparent distance of the pyrometer to the interface.

The inaccuracies in the inverse temperature estimation schemes such as the 1-D
ellipsoidal model is fortunately low since the variation in the temperature field in the
chip-tool interface is mitigated at the far field by the dampening of higher spatial
temperature frequencies by the tool body [26]. Though the curve fitted peak temperature
based on pyrometer readings [See Appendix D] generally increased with cutting speed as
shown in figure 5.3, there were many instances where the pyrometer did record lower
temperatures at higher speeds. The reason suggested is that the interface temperature in
the inverse estimation was a function of both the pyrometer temperature and the contact
area (which showed a decreasing trend as explained in the following section). This is
direct evidence of the presence of steep temperature gradients in the cutting tool during
cutting since this implies that the temperature of the area of heat input into the tool is far
higher than the temperature outside this region. From the argument of the ellipsoidal
model, this stands justified since, the volume for heat removal is higher farther away

from the tool [figure 3.2], and hence represents a greater capacitance for storage of heat.

The trends of the rake face temperature with cutting speed for the as-received
(normalized) steels were not as uniform as that obtained with the spherodized steels
[Appendix C]. The pearlitic structure of the steel, which represents a lower mean-ferrite
path [25], could have made the flow characteristics of the steel substantially different
from that of the spherodized counterpart. The undetermined pre-work in the steel could

have also influenced the chip-tool interface temperature.

5.3 Contact Area Trends

A significant finding in the results [Appendix C] was that the chip tool contact area
showed a decreasing trend with cutting speed in both the spherodized and unspherodized
cases. Subramaniam et al. [20] observed an increase in the crater area with cutting

speed. The difference between the chip tool contact area and the crater area is that while

46

the chip-tool contact area corresponds to the region of heat input into the tool and the area
of partial or complete seizure [40], the crater area corresponds to the amount of coating
and substrate material carried away by the chip. The contact length was similar among
the various coatings, which was also observed by Dearnley [9]. As is evident from figure
5.4, crater wear commences sooner at a higher cutting speed, possibly due to the
chemical inertness and the resistance to surface traction offered by the coating material at
lower speeds. It can also be noted that the alumina coating did not wear as much as the
other coatings at the same cutting speed Considering the remarkable chemical inertness

of A1203 [7], this elucidates the thermo-chemical component of crater wear.

TiN TiCN A1203

 

Figure 5.4: Photographs of K420 carbide coated with TiN, TiCN and A1203 respectively alter machining
spherodized A181 1045 steel at increasing cutting speeds at f = 0.356 mm/rev and doc = 1.905 mm.

47

Among the various work materials investigated, the chip-tool contact area showed a
slightly decreasing trend with carbon content of the steel. The contact area is conjectured
to be a strong function of the fracture toughness [Appendix C] of the work material, since
this property governs the deformability at the shear zone.

5.4 Flank Wear

Flank wear has widely been recognized as the most appropriate criterion for tool life
[11]. Chubb and Billingham [15] found that the removal of the coating in the region of
the flank could accelerate the rate of wear. Hence, the resistance of a coating in the
progression of flank wear deserves greater attention. Dearnley [9] and Cho and
Komvopoulos [10] concluded that the WC phase in the substrate is prone to dissolution
into or or y phases of steel and dissolution wear supercedes any other mechanism of wear

for the case of uncoated carbide grades without any TiC or TaC alloying [13].

In the present work, as expected, the flank wear-rates increased with the cementite
content of the steels. As noted by Kim and Durham [17], the region of the flank wear
could be divided into several zones of damage which includes the zone reminiscent of
superficial plastic deformation and plowing by small carbide grains [10]. In this study,
flank wear was measured at the region of wear land exhibiting uniformity in wear pattern.
Ramalingam and Wright [2] had noted that alloy chemistry does not satisfactorily explain
the machinability of steel since nominally identical alloys yield different machinabilities
in different heats. Thus, it has been corroborated herein that the constitution of the steel
in terms of its hard and soft phases is responsible for flank wear and hence its
machinability. Abrasion by other means (such as the wear debris generated by the tool)
cannot give rise to a uniform wear pattern since this form of abrasion is stochastic in

nature.

The effect of competition between the various mechanisms in coated tools is the
problem at large and requires careful introspection. It was seen in the present work that

the A1203 coated inserts showed scouring marks uniform in length from the cutting edge

 

48

[Figure 5.5] while the other coated inserts showed scouring marks but with a shiny band
where the carbide substrate was exposed. It can also be seen in Appendix C that TiCN
coating showed the lowest wear-rate. In the context of coated tools, it is common
experience that coated tools exhibit a lower crater wear-rate due to the impedance to solid
solution formation offered by the coating. Given these facts, the superior chemical
inertness of A1203 [7] and the higher hot hardness of TiCN [31], it can be concluded that
temperatures prevalent at the flank do allow preferential dissolution of WC to take place
but the progression of flank wear is resisted by the mechanical superiority of the coating.
This resistance is a property greatly determined by the hardness of the coating at the flank
temperature.

It should be noted that the preceding arguments would be true only if abrasion is the
predominant mode of mechanical wear. This may not hold if more complex phenomena
such as thermal cracking, mechanical fatigue cracking, chipping or fi'acture start adding

to the damage of the cutting tool.

49

100 m/min 270 m/min

TiN

 

   

 

00
.E
‘3

Z o
‘11-:

Q o

{-1 c:
.2
H
8
.5:
D

M NW“ wa~u,.s.r~' 's..,..~....,__. .
C
A.“

 

Figure 5.5: SEM micrographs (800x) of the wom flank surfaces of the coated cutting tools after machining
spherodized A181 1045 steel for low and high cutting speeds.

50

In the overall analysis, the TiCN coating showed the highest resistance to flank
wear. A decreasing trend in the wear-rate with flank temperature was noted in the case of
spherodized steels. This can be attributed mainly to a higher softening rate of the
abrasive particles. Suh [19] mentions that Fe3C is capable of dissociation at high
temperatures, vis-a—vis oxide and nitride inclusions. This is also a possible reason for the
decrease in wear-rate with temperature, though not quantitatively ascertained. In
particular, the alumina coating exhibited the greatest sensitivity to flank wear due to a
higher softening rate [7] reflecting the contrasting properties of the A1203 and the TiCN
coatings. While Brun et al. [4, Section 1.2 (paragraphs 2,3,4)] had observed excessive
wear-rates while machining silica reinforced aluminum and the same was the case with
Ramalingam and Wright [2], wear was tolerable in the machining of plain carbon steels.
In view of the fact that SiC, SiOz and A1203 are more refractory materials than cementite,
as can be seen from their hot hardness [7], they can inflict a greater damage on the tool
than cementite. One can therefore generally conclude that inclusions and reinforcements
in a material should be chosen so that they strengthen the material but do not degrade

machinability at higher temperatures encountered in high speed machining.

The final part of the analysis is an attempt to calibrate the comprehensive wear
equations for the three coatings. Figure 5.6 shows the relevant graphs for the flank wear
trends for three coatings. As can be seen from the figure 5.6, there is a convincingly
decreasing trend in the spherodized steels and a moderately increasing trend with the
unspherodized steels. The increasing trend in the case of the latter can be mainly
attributed to the fact that the cementite particles are firmly rooted in the ferrite matrix.
This supports modeling of the former with the three-body model and the latter with the
two-body wear model. The decreasing trend should not give the notion that flank wear
decreases with cutting temperature, rather, the rate of damage on the flank decreases with
cutting temperature. The trends are not clearly evident with TiCN since the hard coating
[31] was impervious to flank wear, due to the inadequate duration of the cuts. This

testifies to the fact that abrasion is the rate-controlling mechanism on the flank.

 

 

 

TiN o . 1018(sp)
30 , :1 1045(sp)
25 x Ag--- 1065(sp)
A 20.
E _._.___10958
a 15 . , x ( p)
5 i X C] A“A 7.x
oi 33"%"9<>'°
700 900 1100 1300
Flank Temperature ('C)
TiCN ..- 0 __1018(sp)
30
_ --- 1045 s
25 1:1 ( P)
E 20 l A 1m5(59)
: 15 1 . x» 1095<sp1
3 1o;
51'
o 4612;:wa
700 900 1100 1300
Flank Temperature (‘6)
N20: . --e——1018(SP)
30 1 x . g.__1045(sp)
:2 . A A 1065(sp)
3 ' L A , x , 1095<sp1
"‘ 15 — _ _ _x
i x8. ~
- 1° 8....
5 . Cl '0
0 . f _. _..
700 900 1100 1300

Flank Tem pe rature (’C)

51

 

 

TiN —e—— 1018 M)
30 - —--1:—J~—- 1045(un)
25 ' A 1070 (un)
.. 20 -
E
E 15 .
S 10 .9
5 xx 13;“ i
. Are... “P“
700 900 1100 1300
Flank Temperature ('C)
TiCN ——e—— 1018 (un)
30 l
25 . __ 43--- 1045 (un)
3 ZOJ 1070(un)
: 15 . A '
Sr 10.
51
o 4 0——0 Mona

 

 

 

700 900 1100 1300
Flank Temperature (‘0)
A110, ___+__ 1018 (un)
30 . a--- 1045(un)
25 .
A 20 .l A 1070 (UN)
E
E 15 .
10 . A
5 4 $“AB-B“"” _..g
0 . a ,
700 900 1100 1300
Flank Temperature (‘6)

Figure 5.6: Plots showing the variation of the volume flank wear per sliding distance for the three coating

materials. (sp—spherodized steel, un-unspherodized steel)

52

5.5 Crater Wear

Crater wear on coated inserts presents an interesting topic for investigation. An
increasing trend [Figure 5.7] was noted in most cases. The general pattern of wear on the
rake face consisted of a central region where the carbide substrate is exposed and a
peripheral area where the coating seemed to resist removal by the flowing chips [Figure
5.4]. Crater wear in coated inserts is pronouncedly lesser than in uncoated carbides. It is
therefore doubtful as to whether the same criteria for tool life [1] can be employed, and
hence, suitable amendments are needed for a tool life criteria for coated cutting tools. It is
speculated that there is a connection between the erratic nature of the rake face
topography in coated tools, as observed in this study, unlike uncoated tools, to the wide
variation in tool life of coated tools noted by Chubb and Billingham [15]. It was also
found that the point of greatest depth on the crater was 500-700um away from the cutting
edge implying that the peak temperature occurred in the vicinity of this point. Moreover,
the variation in the texture from the smoothly worn carbide substrate to the parental

roughness of the coating suggested a wide range for the temperature on the rake face.

An explanation for the dependence of crater wear with cementite content can be
elicited from the preceding point. Dissolution wear has no particular relation to the
cementite content in the steel. That being the case, what was the reason for the increasing
trend in the crater wear-rate with carbon content in the steel [Figure 5.7], if dissolution
was the rate-controlling mechanism on the rake? A closer examination of the rake face
revealed that the depth of the crater was somewhat proportional to the width of the crater.
It can therefore be concluded that, at the periphery of the crater the coating was being
abraded, whereas at the center of the crater, the exposed substrate was being dissolved
continuously. Hence abrasion brought about the widening of the crater exposing a greater
part of the substrate and dissolution brought about the deepening of the crater. This
justifies the need for both abrasion and dissolution in the quantitative modeling of crater
wear. It is suggested that in the design of multi-layered coatings, the outer layer be a
coating resistant to abrasive wear and the intermediary layer be one resistant to

dissolution.

 

 

 

 

 

 

3"7'1‘0186571
: a 1045(89)
TiN i ._-A--_1065(sp)
61 I -—e——1095(sp)
l ,_
5 4 ' °
:1. 2 4: 0 If 'A“;- A o
. 0'
0 ; AD“?! ‘0‘ o I
800 1150 1500
Temperature (°C)
0 ‘ " .7 018 (spf
_ g 1045(sp)
TICN _ A--- 1065(sp)
6 -- 0 1095(sp)
‘ 9
.5 4 ' .'
5
:r. 2 ; 6””(39
01 5 3;,
. o_
0800 1150 1500
Temperature (°C)
‘ 3“,”10T8Tsi)"
A1203 (3 1045(sp)
am 1065(813)
5 1 .91.” 1095(sp)
9 _
.5 4
E !
: o
E 2 l o. D
1 AD
0 1 ° . ___.
800 1150 1500

Temperature (°C)

53

pmlmin

umlmin

urn/min

 

 

 

 

 

 

 

 

Figure 5.7 Plots showing the variation of the crater wear-rate for the three coating materials.

(Sp-spherodized steel, un-unspherodized steel)

0 1018(un) .
TiN
6 ---{3_--1045(un)
, _A_1070(un) l
4, |______ _ _.
! ° I3
I A°<>~ 36:11.1
0 + I I 8 SE! 1
800 1150 1500
Temperature (°C)
1018(un)
TiCN "1045(un) l
6 ‘ IE “A_1070(un)
4
A“. 1:
2
EA]
0 __ .' .
0 9 ° 334 o 5
800 1150 1500
Temperature (°C)
0 1018(un) ;
A|203 -_-{}-_1045(un) |
5 +1070(un) E
4 _
2 .' A?” """ A El
1 A: 13.-----
3 El
0 l 99 . .
800 1150 1500
Temperature (°C)

54

5.6 Calibration

With use of the equations in tables 5.1 and 5.2 one can obtain a calibration for the
wear-rates based on equations 2.6 and 2.7 and hence values for the constants A and B. As
a recapitulation, the value for 'A' has to be obtained for flank wear, since it is dominated
only by abrasion and both 'A' and 'B' for crater wear since both mechanisms operate.
Further, for flank wear, it should be remembered that the three-body wear mechanism is
assumed to operate in the machining of spherodized steels whereas the two body
mechanism is assumed to operate in the machining of unspherodized steels. Furthermore,
in the case of abrasion on the crater only the three-body mechanism is assumed to
operate, since temperatures are excessively high for the abrasive to retain any hardness.

The details of the governing relations of the wear mechanisms were explained in Chapter
2.

As an example, calibrations for A1203 for flank wear and crater wear while
machining spherodized steels is illustrated in figure 5.8. This was performed on the
software SigmaPlot®. The values are listed in Appendix C. Figure 5.9 shows the plot of
the constants 'A' and 'B' obtained from the calibration process. There is a general
agreement in the trends of the models and that of the experimental results, however, it
was not impressive for the as-received steels probably because of the pre-work in the

steels and presence of distributed hard spots, which injure the tool in an arbitrary fashion.

Since ‘A’ and ‘B’ are quite different in their order of magnitude, they are plotted on a
logarithmic scale in figure 5.9 for visualization purposes. It can also be observed in figure
5.9 that the values for individual 'A's and 'B's are within the same order of magnitude.
This is promising since this substantiates the correctness of the modeling, despite the fact

that only one cut was taken for each machining condition.

Volume Flank Wear per
Sliding Diatance(‘m1Im)

Wear Rate ( ’mlmin)

 

55

Flank Wear Calibration for AIzO,

 

 

 

 

 

30 .. o 1018(sp)
25 _, t predicted 1018(sp)
\\ x. A 1045(sp)
2° ‘\ _ _ _ _ predicted 1045(sp)
15 ° A35." + :1: 1065(sp)
10 7 + s :; ... ‘ ....... predicted 1065(sp)
. O ... 1095(sp)
5 A _ . __ . _ predicted 1095(sp)
0 a . . 5
700 850 1000 1150 1300
Flank Tem perature (°C)
C rater Wear Calibration for A1203
6 1 0 1018(sp)
5 + predicted 1018(sp)
‘ A 1045(sp)
4 1 ......... predicted 1045(sp)
3 x 1065(sp)
_A- '+ predicted 1065(sp)
2 +1, i + 1095(sp)
. X A .. . predicted 1095(sp)
o 7,_ _ ..-“ , . .
800 1000 1200 1400

Crater Tem perature ('C)

Figure 5.8: Plot of the predicted and the experimental wear data for A1203

while machining spherodized steels

Flank Wear Calibration for
spherodized eteele
30
- 25 atom-2)
% rout-p)
3 20 roost-p)
E 13100500)
15
g 10
'6
" 5
2 l
O ,
10000000 [ , #
... 1000000
3 ‘3 100000 ‘
% a 10000 1‘
> g 1000.
In a:
<" _e 100
s 10
1 .
10000000 . 7 Hi ,A,,, 7,_
1000000
8 % 100000 1‘
3 D 10000
Te :1 '
i; g 1000
<- § 100 .j
5 \

 

 
     

Crater Wear Calibration for epherodized eteele

56

A (Arbitrary Units

 

   
   
   

Flank Wear Calibration for
Unepherodized Steele
4000 1
3000 amount)
.10451011)

I 1 070(un)

O

TIN TiCN Al203

 

D 1018(sp) A value I 1018(sp) B value I 1045(sp) A value D1045(sp) B value
I 1065(sp) A value I 1065(sp) B value I 1095(sp) A value 1095(sp) B value

 

D1018(un) A value
D1045(un) 8 value

Crater Wear Calibration for unepherodized eteele

 

TiCN Al203
I 1018(un) B value I 1045(un) A value
I 1070(un) A value I 1070(un) B value

Figure 5.9: Plots of the constants obtained from the calibration process.
(sp - spherodized steel, un — unspherodized steel)

Chapter 6

SUMMARY AND CONCLUSIONS

The following summarizes the work performed: -

Wear of the flank and the crater of coated carbide tooling while machining plain
carbon steels was investigated. Wear models in tribology and the literature were adopted
in modeling and studying the phenomena. An attempt was made to relate tool wear and

cutting temperatures in a quantitative fashion.

The infrared pyrometry technique was utilized in the measurement of rake face
temperature. The 1D ellipsoidal scheme of Yen and Wright [44] was used in the inverse
estimation of the chip-tool interface temperature. Oxley's recipe for the flank temperature
was adopted. The dependence of interface temperatures was on the process parameters
was studied. The infrared pyrometry technique proved to be a versatile and easy method
enabling easy measurement of the rake face temperatures of the tool, but relied heavily
on the accuracy of the inverse estimation schemes for its accuracy in prediction. The

related problem of the interfering chips was efi‘ectively dealt with.

The three-body and two-body wear models were used for describing abrasion of
the flank surface of spherodized and unspherodized steels respectively whereas the
dissolution and the three-body wear models were used in obtaining a comprehensive

description of crater wear in both steels. Recommendations for further developments in

57

58

tool materials such as multi-layered coatings and work materials include having an
intermediary layer resistant to dissolution and an outer layer resistant to abrasion. The
criteria for the predictability of tool wear have been outlined clearly. Reinforcements in a
work material should be chosen so that they don’t degrade the machinability of a

material.

Among the major findings in this study are that steep gradients in temperature are
present on the rake face of the cutting tool. Chip-tool contact area decreases with cutting
speed. Contact area is proportional to the fracture toughness [Appendix C]. Flank and
crater wear were found to increase with the cementite content in the steel. Spherodized
steels showed a decreasing flank wear per sliding distance with temperature whereas hot-
rolled, unspherodized steels showed a moderately increasing trend. Both types of steels
showed an increasing rate of crater wear with temperature. Modeling of tool
temperatures, which was undertaken, showed that there were steep temperature gradients

in the cutting tool and temperatures depended on the area of heat input into the tool.

The appropriate conclusions drawn include affirming that abrasion was the rate-
controlling wear mechanism in the wear of coated carbide tools while machining plain
carbon steels, though preferential dissolution of the carbide substrate is also possible.
Crater wear involved both abrasion and dissolution. The wear models used in tribology

proved to be reasonably accurate in predicting wear rates on both the crater and the flank.

APPENDICES

Appendix A

COMPUTER ALGORITHM

A.l The main part of mainprog.f

character coating*8,elea(6)*2,compound*8

OOOOOOOOOO

(10

30

31

52
53

68
69

70

‘00000000

integer xa(6)

'Coating' IS THE INPUT VARIABLE FOR THE COATING, 'elea' IS THE GLOBAL

ARRAY CONTAINING THE ELEMENTS OF THE COATING, READ FROM THE FILE COMP1.DAT
THERE ARE FIVE FILES IN ALL. MAINPROG.F WHICH IS THE MAIN PROGRAM FILE,
MECH1.DAT, WHICH CONTAINS THE HARDNESS DATA, CHEM1.DAT CONTAINS THE CHEMICAL
DISSOLUTION DATA AND EXFreeEn.dat WHICH CONTAINS THE EXCESS FREE ENERGY OF
SOLUTION OF INDIVIDUAL ELEMENETS IN ALPHA IRON. THE PROGRAM CAN WORK

FOR AT THE MOST TERTIARY COATINGS. IN ORDER TO UPGRADE FOR MORE COMPLEX
COATINGS ON HAS TO JUST ADD STATEMENTS RELATED TO THE ELEMENTS IN THE 'elea‘
ARRAY. ALSO, FOR NON STOICHIOMETRIC COATINGS ONE HAS TO ENTER THE NUMBERS WITH
AN INTEGRAL NUMBER OF ATOMS IN THE MOLECULE.

The following unit reads the file 'COMP1.DAT' for the compound
name and its elements.

INPUT COATING AND ITS DETAILS
write(*,30)
format('What is the coating?')
read(*,3l) coating
format(A8)
open(unit=2,file='COMP1.DAT',blank='ZERO',status='OLD')
read(2,53,end=68) elea(l),elea(2),elea(3),xa(l),xa(2),xa(3),compound
format(A2,A2,A2,12,12,12,A8)
if(compound.eq.coating) then
goto 70
else
endif
goto 52
write(*,69)
format('SORRY COATING IS NOT TO BE FOUND IN THE LIST')
stop
close(2)
call mechwear(coating)
call chemwear(coating,elea,xa)
end

Listing of 'subroutine mechwear'

THE FOLLOWING SUBROUTINE GENERATES THE MECHANICAL WEAR DATA FOR

THE COATING BASED ON THE HARDNESS OF THE ABRASIVE AND THE COATING.

BOTH TWO BODY AND THREE BODY WEAR DATA ARE GENERATED.

THE DECLARATIONS ARE AS FOLLOWS: -Ctng STANDS FOR THE COATING.

ALL VARIABLES WITH 'COMP' ENDING DENOTE COMPOUNDS, ANY VARIABLE WITH

h IN IT DENOTES HARDNESS.THE HARDNESS OF

ANY MATERIAL CONCERNED IS EXPRESSED IN THE FORM H(T)=HO*EXP(-ALP*T). THIS

WOULD THEN REPRESENT THE SOFTENING IN THE MATERIAL AT A HIGHER TEMPERATURE. ALL

59

60

C REPRESENT TEMPERATURE. THE SUFFIX 'ref'STANDS FOR THE REFERENCE COMPOUND AND THE
SUFFIX
C 'ab' STANDS FOR THE ABRASIVE. TO CHANGE THE ABRASIVE FOR TEH MODEL ONE HAS TO
C CHANGE THE FIRST ELEMENT IN THE MECH1.DAT FILE. SINCE THE HARDNESS OF A MATERIAL
C MAY NOT ALWAYS BE EXPRESSIBLE IN THAT FORM, AN INTERMEDIATE TEMPERATURE
C IS CHOSEN SO THAT FROM THAT TEMPERATURE ONWARDS, THE HARDNESS TAKES A
C DIFFERENT FORM. HENCE THE TWO TEMPERATURES t1 AND t2.
subroutine mechwear(ctng)
character comp*8,refcomp*8,ctng*8,abcomp*8
real h1,alp1,h2,alp2,refhl,refalp1,refh2,refalp2
real ahl,aa1pl,ah2,aalp2
integer t1,t2,reftl,reft2,t,i,at1,at2,abrat
real th,ah,refh,relwear,abrwearcoeff,twobodyrelwear,abrwearcoefftwobody
C IN THE MECHl.DAT FILE, THE FIRST RECORD CORRESPONDS TO THE REFERENCE TOOL MATERIAL
C FOR THE RELATIVE WEAR CALCULATION. THE SECOND CORRESPONDS TO THE ABRASIVE
PARTICLE.
C HENCE THE ORDER OF DATA COLLECTION FROM THE FILE.
c GETTING THE DATA FOR THE REFERENCE TOOL MATERIAL
open(unit=2,file='MECH1.DAT',blank='NULL',status='OLD')
read(2,500) refcomp
500 format(A8)
read(2,510) refh1,refalp1,reft1
510 format(F7.1,FlO.8,I4)
read(2,520) refh2,refalp2,reft2
520 format(F7.1,F10.8,I4)
C GETTING THE DATA FOR THE ABRASIVE MATERIAL
read(2,521) abcomp
521 format(A8)
read(2,522) ah1,aalp1,at1
522 format(F7.l,F10.8,I4)
read(2,523) ah2,aalp2,at2
523 format(F7.1,F10.8,I4)
close(2)
C THE NEXT UNIT RELATES TO SERACHING THE DATA FOR THE CANDIDATE TOOL MATERIAL. AS
MENTIONED
C BEFORE, HARDNESS IS EXPRESSED IN THE SAME FORM AS BEFORE. AGAIN FOR MORE
CORRESPONDENCE
C WITH THE EXPERIMENTAL VALUES, THE HARDNESS FORMULA IS SPLIT INTO TWO REGIONS.
c LOCATING THE HARDNESS DATA FOR THE T/M IN THE FILE
open(unit=2,file='MECH1.DAT',blank='NULL’,status='OLD')
525 read(2,530,end=560) comp
530 format(A8)
readi2,540) hl,alpl,tl
540 format(F7.1,FlO.8,I4)
read(2,550) h2,alp2,t2
550 formatiF7.1,F10.8,I4)
if(comp.eq.ctng) then
goto 580
else
endif
goto 525
S60 write(*,570)
570 format('SORRY ABRASION DATA NOT FOUND FOR THE TOOL MATERIAL')
return
580 close(2)
C THIS IS THE MAIN UNIT WHICH GENERATES THE WEAR DATA. AN OUTPUT FILE IS OPENED FOR
WRITING
C AND HEADINGS ARE WRITTEN. A SEPARATE SUBROUTINE 'hardness' FOR THE HARDNESS
CALUCULATION USING THE
C FORMULA IS WRITTEN. THAT WILL BE SEPARATELY EXPLAINED. APPROPRIATE ERROR MESSAGES
ARE

C ALSO WRITTEN SIDE BY SIDE.

61

openlunit=3,file='mechwear.out',blank='NULL',status='NEW')

writel3,593)
593 format('Temperature(C) Relative Abrasive wear Two body relative wear
T/Hardness A/Hardness wear coeff(3-body) Wear coeff.(2-body).')

do 800 i=l,l4

t=(i-l)'lOO

if(t.le.reftl) then
refh=hardness(refhl,refalpl,t)

elseif(t.le.reft2) then
refh=hardness(refh2,refalp2,t)

else
write(*,600)
600 format('SORRY TEMPERATURE EXCEEDED FOR THE REF. MATL.:Termination')
return
endif

if(t.le.t1) then
th=hardness(hl,alp1,t)

elseif(t.le.t2) then
th=hardness(h2,alp2,t)

else
write(*,610)
610 format('SORRY TEMPERATURE EXCEEDED FOR THE T/M: Termination')
return
endif
C THIS UNIT MEASURES THE ABRSIVE TEMPERATURE. AS SHOWN IN THE NEXT LINE A MODEST
ASSUMPTION
C THAT THE ABRASIVE TEMPERATURE IS 90% OF THE INTERFACE TEMPERATURE IS MADE. THIS IS
AN
C IMPORTANT POINT AND IS QUESTIONABLE.
abrat=90*(i-l)
if(abrat.le.atl) then
ah=hardness(ahl,aalp1,abrat)
elseif(abrat.le.at2) then
ah=hardness(ah2,aalp2,abrat)
else
write(*,630)
630 format('SORRY,TEMPERATURE EXCEEDED FOR ABRASIVE MATERIALzTermination')
return
endif
C A SEPARATE SUBROUTINE TO CALCULATE THE RABINOWICZ WEAR RATE CALCULATION IS WRITTEN.
ITS NAME IS
C 'volume'. THE WORKING OF THE SUBROUTINE IS EXPLAINED IN DETAIL LATER ON.
'abrwearcoeff' IS THE
C VARIABLE WHICH CORRESPONDS TO THE 3 BODY WEAR. THE TWO BODY WEAR HAS A VERY SIMPLE
FORMULA AND
C HENCE NO SUBROUTINE IS NEEDED FOR THAT.
abrwearcoeff=volume((th/ah),th)
relwear = volume((th/ah),th)/volume((refh/ah),refh)
twobodyrelwear = refh/th
abrwearcoefftwobody=l/th
write(3,640) t,relwear,twobodyrelwear,th,ah,abrwearcoeff,abrwearcoefftwobody
640 format(I4,' ',E9.4,' ',E9.4,' ',F7.1,'
',F7.1,' ',E9.4,' ',E9.4)
800 continue
close(3)
return
end

A.l.2 Listing of 'subroutine hardness'

C THIS SIMPLE FUNCTION DETERMINES THE HARDNESS OF THE MATERIAL WITH THE ASSUMED FORM

A.l.3

C

ARE

C
BETTER.

A.l.4

C
INPUTS
C

THE

62

H(T)=HO*EXP(-ALPHA*T). THE INPUTS FOR THE SUBROUTINE ARE H0, ALPHA AND T

real function hardness(h,alp,temp)
real h,alp,rt
integer temp

rt=real(temp)
hardness=h*exp(-alp*rt)
return

end

Listing of 'subroutine volume'

THIS SIMPLE FUNCTION DETERMINES THE RABINOWICZ WEAR RATE CALCULATION. ITS INPUTS

THE HARDNESS RATIO AND THE TOOL HARDNESS. THIS IS ELABORATED IN THE LITERATURE

real function volume(ratio,hard)
real ratio,hard

ifiratio.lt.0.8) then

volume = 1/(3*hard)

return
elseif(ratio.lt.l.25) then

volume = (exp(-2.S*alog(ratio)))/(5.3*hard)
return
elseif(ratio.gt.1.25) then

volume = (exp(—6.0*alog(ratio)))/(2.43*hard)
return
else
endif

return
end

Listing of 'subroutine chemwear'

THIS IS THE MAIN SUBROUTINE WHICH COMPUTES THE CHEMICAL DISSOLUTION WEAR. ITS
ARE THE COATING, ELEMENTS ARRAY AND THE STOICHIOMETRIC NUMBER.

subroutine chemwear(ctng,ele,x)

character refcomp*8,ctng*8,comp*8,ele(6)*2,elem*2
real refmlv,mlv,refll,refm1,refnl,ref12,refm2,refn2
real ll,ml,n1,12,m2,n2

integer reftl,reft2,tl,t2,freen

integer x(6),refx(6)

integer free(6),refree(6)

integer i,t

real sol,refsol,relsol

THE REFERENCE COMPOUND FOR THE RELATIVE WEAR CALCULATION IS THE FIRST ELEMENT IN
CHEMl.DAT FILE. HENCE THIS CAN BE ALTERED SUITABLY.

READING THE DATA FOR THE REFERENCE COMPOUND
open(unit=2,file='CHEMl.DAT',BLANK='ZERO',status='OLD')
read(2,100) refcomp,refmlv
format(A8,F7.2)
read(2,110) refll,refml,refnl,reft1
format(FlO.1,F7.2,F7.2,I5)
read(2,120) ref12,refm2,refn2,reft2
format(FlO.1,F7.2,F7.2,IS)
close(2)

THIS UNIT COLLECTS THE DATA FOR THE CANDIDATE COATING FROM THE FILE.
LOCATING DATA FOR THE INPUT COATING FROM THE DATA FILE

63

open(unit=2,file='CHEM1.DAT',blank='ZERO',status='OLD')

160 read(2,170,end=198) comp,mlv
170 format(A8,F7.2)
read(2,180) ll,ml,n1,t1
180 format(FlO.1,F7.2,F7.2,IS)
read(2,190,iostat=ios) 12,m2,n2,t2
190 format(FlO.l,F7.2,F7.2,IS)
if(comp.eq.ctng)then
goto 200
else
endif
goto 160
198 write(*,199)
199 format('SORRY DATA NOT AVAILABLE. COMPUTATION HALTED')
return
200 close(2)
C IF THE MOLAR VOLUME IS ZERO, THEN THE COMPUTATION IS HALTED FOR OBVIOUS REASONS.
C TO SEE IF THE MOLAR VOLUME VALUE IS ZERO OR NOT

if (m1v.eq.0) then
write(*,201)

201 format('SORRY, THE MOLAR VOLUME IS ZERO. COMPUTATION HALTED')
return
else
endif
C THE NEXT SET OF THREE UNITS LOCATE THE EXCESS FREE ENERGY OF SOLUTION OF THE
ELEMENTS
C IN THE COATING. AT PRESENT THREE ELEMENTS CAN BE PRESENT. IF ONLY TWO ELEMENTS ARE
PRESENT
C THEN THE THIRD IS DUMMIED 'X' AND THE DATA CORRESPONDING TO IT IN THE FILE IS A
ZERO.
C TO LOOK FOR THE EXCESS FREE ENERGY OF SOLUTION IN ALPHA IRON FOR THE ELEMENTS
open(unit=2,file='ExFreeEn.DAT',blank='ZERO',status='OLD')
205 read(2,210) elem,freen
210 format(A2,I7)

if(elem.eq.ele(l)) then
free(l)=freen
goto 215
else
endif
goto 205
215 close(2)

open(unit=2,file='ExFreeEn.DAT',blank='ZERO',status='OLD')
206 read(2,211) elem,freen
211 format<A2,I7)
if(elem.eq.ele(2)) then
free(2)=freen
goto 216
else
endif
goto 206
216 close(2)

open(unit=2,file='ExFreeEn.DAT',blank='ZERO',status='OLD')
207 read(2,212) elem,freen
212 format(A2,I7)
if(elem.eq.ele(3)) then
free(3)=freen
goto 217
else
endif
goto 207
217 close(2)

C SINCE TiN IS THE REFERENCE COATING HERE, THE APPROPRIATE CONSTANTS ARE FIXED

64

C AND THE DATA COLLECTED. ONCE HAS TO REMEMBER TO CHANGE THIS IF ANOTHER COATING IS
TO BE
C CHOSEN AS THE REFERENCE COATING. THIS WAS DELIBERATELY DONE TO KEEP THE CODE
SIMPLE.
C FIXING THE CONSTANTS FOR THE REFERENCE MATERIAL

refx(1)=1

refx(2)=1

refx(3)=0

open(unit=2,file='ExFreeEn.DAT',blank='ZERO',status='OLD')
read(2,218) elem,refree(1)

218 format(A2,I7)

read(2,219) elem,refree(2)
219 format(A2,I7)

read(2,220) elem,refree(3)
220 format(A2,I7)

close(2)

C THIS UNIT CREATES AN OUTPUT OF THE SOLUBILITIES. TEMPERATURE IS REPRESENTED IN
KELVIN HERE.
C THE SUBROUTINE 'solubility' IS EXPLAINED SEPARATELY IN THE END. THE OUTPUT FILE IS
CHEMDISS.OUT
C WHICH HAS THE APPROPRIATE HEADINGS.
c NOW FOR CREATING THE OUTPUT OF THE SOLUBILITIES

open(unit=2,file='chemdiss.out’,blank='ZERO',status='NEW')
write(2,400)

400 format('temp(K) solubility ref solubility relative solubility')
do 490 i=1,l4
t=i*100 + 273

C NOTE THAT TEMPERATURE HAS TO BE REPRESENTED IN KELVIN FOR CALCULATIUNG
C THERMODYNAMIC DATA.
C THIS UNIT COMPUTES THE SOLUBILIY OF THE CANDIDATE COATING MATERIAL. IT INVOKES
C THE solubility FUNCTION FOR THIS. THE FUNCTION IS ELABORATED LATER ON.
if(t.lt.tl) then
sol=m1v*solubility(ll,m1,n1,t,x(1),x(2),x(3),free(1),free(2),free(3))
elseif(t.lt.t2) then
sol=mlv*solubility(12,m2,n2,t,x(l),x(2),x(3),free(l),free(2),free(3))
else
write(*,410)
410 format(' SORRY TEMPERATURE EXCEEDED FOR THE COATING MATERIAL’)
return
endif
C THIS UNIT CALCULATES THE SOLUBILITY OF THE REFERENCE TOOL MATERIAL. DATA HAS
C ALREADY BEEN COLLECTED FOR THIS.

if(t.lt.reft1) then

refsol=refmlv*solubility(refll,refm1,refn1,t,refx(l),refxiZ),refx(3),refree(1),refree(2),
refree(3))
elseif(t.lt.reft2) then

refsol=refmlv*solubility(ref12,refm2,refn2,t,refx(1),refx(2),refx(3),refree(1),refree(2),
refree(3))

else
write(*,420)
420 format(' SORRY TEMPERATURE EXCEEDED FOR THE REFERENCE MATERIAL')
return
endif

relsol= (sol)/(refsol)
write(2,430) t,sol,refsol,relsol

430 format(IS,' ',E10.5,' ',E10.5,' ',E10.5)
490 continue

close(2)

return

end

65

A.l.S Listing of 'subroutine solubility'

C THIS IS THE SUBROUTINE WHICH COMPUTES THE SOLUTBILITY IN ALPHA IRON BASED ON THE
FORMULA DEVELOPED

C BY KRAMER AND SUH [1980] IN THEIR PAPER. IT TAKES THE THERMODYNAMIC DATA OF THE
COATING MATERIAL

C AND RETURNS THE SOLUBILITY WHICH IS JUST A NUMBER. IT HAS TO BE CALIBRATED BASED
ON EXPERIMENTS

C IT CAN BE SEEN THAT IT DEPENDS SENSISTIVELY WITH TEMPERATURE. FREE ENERGY OF
C FORMATION OF THE COATING COMPOUND IS EXPRESSED IN THE FORM G=A +B LOG(T) + C.
C AGAIN TO TAKE INTO ACCOUNT, THE PHASE CHANGE IN THE MATERIAL, TEMPERATURE IS
C DIVIDED INTO TWO REGIMES. THE MOLAR VOLUME OF THE MATERIAL IS ALSO TO BE KNOWN
C TO ESTIMATE THE SOLUBILITY IN VOL%. THIS ENTRY IS NEXT TO THE COMPOUND NAME IN
C THE FILE CHEM1.DAT.
real function solubility(sl,sm,sn,st,sx1,sx2,5x3,sfree1,sfree2,sfree3)
real sl,sm,sn
integer st,sxl,sx2,sx3,sfreel,sfree2,sfree3
real rxl,rx2,rx3,rfreel,rfree2,rfree3,lnt,denom,numer,g,rt
R=1.987
If (sxl.lt.10) then
rxl =real(sx1)
rx2 =real(sx2)
rx3 =real(sx3)
else
rxl =l.0
rx2 =real(sx2)/ real(sxl)
rx3 =real(sx3)/ real(sxl)
endif
rfreel=real(sfreel)
rfree2=real(sfree2)
rfree3=real(sfree3)
rt =real(st)
g =sl + (sm*rt*alog10(rt)) + (sn*rt)
if(rx3.eq.0)then
lnt =rx1*alog(rxl) + rx2*alog(rx2)
else
lnt =rx1‘alog(rx1) + rx2’alog(rx2)+ rx3’alog(rx3)
endif
denont =R*rt*(rx1 + rx2 + rx3)
numer =g - (rxl‘rfreel) - (rx2*rfree2) — (rx3*rfree3)- R*rt*lnt
solubility=exp(numer/denom)
return
end
A.2 Final Remarks
C FINAL REMARKS: -
C ADDING A NEW COMPOUND TO THE EXISTING LIST IS TO BE DONE AS FOLLOWS. THE
C COMPOUND NAME WITH ITS RESPECTIVE ELEMENTS IS TO BE ADDED TO THE LIST COMP1.DAT
C THEN THE HARDNESS OF THE COATING IS TO BE ADDED AS INDICATED IN THE MECHWEAR
C SUBROUTINE. IF THERE ARE NEW ELEMENTS IN THE COMPOUND, THEY HAVE TO BE ENTERED
C IN THE FILE EXFreen.DAT. THEN THE FREE ENERGY OF FORMATION OF THE COMPOUND IS
C TO BE ENTERED IN THE FILE CHEM1.DAT AS INDICATED IN THE SUBROUTINE CHEMWEAR.
C IF NON-STOICHIOMETRIC COMPOUNDS ARE TO BE ENTERED, IT IS SUGGESTED THAT
C THE FRACTIONS BE NORMALIZED UPTO TWO DECIMALS AND THEN ADDED TO THE LIST AS
C EXPLAINED BEFORE.
A.3 Listing of 'COMP1.DAT'

Since TiN is the reference coating in the relative wear calculation, it is the first

record in the file. If some other coating is to be chosen, suitable changes are to be made

66

in this and other files and also in the program. In particular, the values in the variable
refx(1..3) has to be changed.

TiN X l 1 OTiN
TiC X l 1 OTiC
AlO X 2 3 0A1203
TiC N 10 5 5TiCN
TiAlN l 1 lTiAlN
HfC X 1 1 OHfC
ZrC X 1 1 OZrC
TaC X 1 1 OTaC
V C X l 1 OVC
NbC X 1 1 ONbC
SiC X 1 1 OSiC
CrC X 7 6 OCr7C6
CrC X 3 2 OCr3C2
MoC X 2 l OM02C
AlC X 4 3 OAl4C3
B C X 4 1 OB4C

W C X 1 1 OWC

W C X 2 1 OW2C
TiO X 1 1 OTiO
TiO X 1 2 OTiOZ
TiO X 2 3 OTi203
TiO X 3 5 OTi3OS
HfO X l 2 OHfOZ
ZrO X l 2 OZrOZ
TaO X 2 5 OTa205
V O X 1 1 OVO

V O X l 2 OVOZ

V O X 2 3 OV203
V O X 2 5 OV205
CrO X 1 2 0Cr02
CrO X 1 3 OCrOB
CrO X 2 3 OCr203
SiO X l 2 OSiOZ
NbO X 1 1 ONbO
NbO X 1 2 ONbOZ
NbO X 2 5 ONb205
TiB X 1 2 OTiBZ
HfB X l 2 OHfBZ
ZrB X l 2 OZrBZ
W B X 2 5 OWZBS
A N X 1 l OAlN
CrN X 2 1 OCr2N
CrN X l 1 OCrN

B N X 1 l OBN
TaN X 1 1 OTaN
HfN X 1 l OHfN
ZrN X 1 l OZrN
NbN X 2 l ONbZN
SiN X 3 4 OSi3N4
MON X 2 1 OM02N

67

A.4 MECH1.DAT

The following is the listing of the data file containing the hardness data. As
mentioned before, the hardness of the material is expressed in the form H(T) (Kg/mm2) =
Hoe"1LT where T is in °C. H0 (in Kg/mmz) and a are constants known from [13]. Since this
may not be form of the empirical data for the whole range of temperatures from 0-
1300°C, this temperature range may be divided into two ranges and separate constants
may be evaluated separately for these ranges. The division in temperature ranges is taken
from the third field in the records. If no division in the temperature range is needed, then
the two temperatures are consecutive °C, as for example in the case of TiN. The first field
stores the Ho value and the second stores the a value while the third field stores the
temperature of validity of the approximation. If the temperature in the third field of the

second record is exceeded, then the computation is halted.

TiN

02563.6 0.001600 1200
02563.6 0.001600 1201
Fe3C

01200.0 0.001347 0400
03320.0 0.003891 1400
TiAlN

02198.4 0.000400 0600
04128.6 0.001400 1200
TiC

03300.0 0.001010 0600
10190.0 0.002890 1600
A1203

02468.5 0.001616 0500
03271.5 0.002180 1200
NbC

02400.0 0.001530 1600
02400.0 0.001530 1600
HfC

03000.0 0.001420 1600
03000.0 0.001420 1601
ZrC

03000.0 0.001660 1600
03000.0 0.001660 1601
TaC

01900.0 0.000775 0540
03620.0 0.001778 1600
WC

01723.9 0.001400 0400
02733.9 0.002300 1200
M02C

01600.0 0.001400 0350
01600.0 0.001400 0351

TiBZ

68

03500.0 0.001890 1000
01072.0 0.000705 1600
HfBZ

03000.0 0.001790 1000
01000.0 0.000686 1600
ZrBZ

02300.0 0.001780 1000
00640.0 0.000495 1600
W285

03100.0 0.001550 1200
30720.0 0.003470 1600
T102

02000.0 0.000116 0600
04000.0 0.000236 0800
TiO

01250.0 0.000599 0800
01250.0 0.000599 0801
HfN

02000.0 0.000857 1000
02000.0 0.000857 1001
SiC

02800.0 0.000090 0400
04050.0 0.001010 0800
Si3N4

01950.0 0.000437 0800
01950.0 0.000437 0801
TiCN

02787.3 0.000400 0400
05496.0 0.002000 1200
A.5 CHEM1.DAT

This file stores the free energy of formation data. As mentioned in the remarks of the
program, the Gibb's free energy of formation G (in cal/mol) of a compound is expressed

in the form of equation A.l[13]: -

G = A + BLog,0(T°K) + C Equation A.l

In addition to this value, the molar volume of the compound is also needed. This
is stored next to the compound name in cm3/mol. The values of A, B and C are stored in
that order as the first three fields in each record. To take phase change into account, it
may be required to split the temperature range of 373K to 1673K into two regions. The
temperature of phase change is stored as the third field in the first record. The
temperature beyond which this formula is not valid is stored as the third field in the
second record. The computation is halted after this temperature.

TiN

-080250.
-O80850.

TiC

-043750.
-044600.

HfC

-052165.
-052165.

ZrC

-047560.
-047995.

TaC

-035655.
-035655.

VC

-020540.
-050540.

NbC

-033910.
-032940.

SiC

-014000.
-027100.

Cr23C6

-098280.
-098280.

Cr7C6

-041710.
-041710.

Cr3C2

-025240.
-025240.

M02C

-Oll330.
-010110.

A14C3

-051600.
-003700.

B4C

-011413.
-014313.

WC

-009000.
-009000.

W2C

-014335.
-014335.

TiO

-123800.
-123800.

TiOZ

-225000.
-225000.

Ti203

-363000.
-363000.

31
0
0

00.
00.
.75
04.
04.
.40
00.
00.
.0

-54
-54

.96
00.
00.
.76
00.
00.
.26
00.
00.

.00
.OO

.00
.00

.00
.00

.00
.00

.48
.48

.00
.00

.35
.00

.30
.73

.00
.00

.00
.00

.94
.94

.53
.00

00
00

65
65

00
00

.26
.26

00
00

00
00

00
00

22

N

(A)

U"!

.20
22.

77

.41
.61

.89
.89

.20
.30

.41
.41

.81
.81

.45
.43

.68

18.1

10.
23.

-13.
-13.

141.
141.

22.
22.

45.
45.

65.
65.

.24
.24

.16
.16

.41
.41

.66
.15

00
00

65
65

.40
.40

11
11

30
30

00
00

00
00

1155
1900

1155
2000

2000
2001

1135
2000

2000
2001

2000
2001

0950
2000

1686
2000

1600
1601

1600
1601

1600
1601

1463
1600

0932
2000

2000
2001

2000
2001

2000
2001

1950
1951

1950
1951

1950
1951

69

T1305

-592500.
-592500.

Hf02

-266000.
-266000.

Zr02

-262000.
-262000.

Ta205

-490000.
-490000.

VO

-099000.
-099000.

V02

-l68000.
-168000.

V203

-294000.
-294000.

V205

-374210.
-140670.

Cr02

-142000.
-142000.

Cr03

-138000.
-123000.

Cr203

-267450.
-267450.

Si02

~215600.
~227700.

A1203

-400810.
-405760.

NbO

-096000.
-096000.

Nb02

-188000.
-188000.

Nb205

-452500.
-452500.

TiBZ

-067230.
-068580.

HfBZ

-078405.
-078405.

ZrBZ

-077400.
-076560.

18.

0
0

.0

00
00

.45
00.
00.

00.
00.

00.
00.

00.
00.

00.
00.

00.
00.

00.
00.

00.
00.

00.
00.

00

00.

00.
00.
.72
-3.

00.

00.

00.
00.

00

00.
.42
00.
00.
.06
00.
00.

50

3.

.00 108.
.00 108.
00 45.
00 45.
00 45.
00 45.
00 104.
00 104
00 18.
00 18.
00 37.
00 37.
00 59.
00 59.
00 105.
00 59.
00 44

00

00 60.
00 30.
.00 60.
00 60.
00 41.
00 48.
98 87.
.75 92.
00 18.
00 18.
00 35.
00 35.
.00 97.
00 97.
00 4
00 5
OO 2
00 2
12 -6.
.94 -18.

33
33

71
71

00
00

88

.88

57
57

42
42

22
22

26
22

O

97
97

64
22

93
93

71
71

62
62

.08
.25

.43
.43

09
35

1950
1951

2500
2501

0135
2136

2200
2201

2200
2201

1800
1801

2200
2201

0950
2201

1650
1651

0450
1000

2150
2151

1700
2000

0923
1800

3050
3051

3200
3201

1750
1751

1155
2000

2000
2001

1135
2000

70

71

AlN 12.58
-072170.0 00.00 23.30 0932
~073750.0 00.00 25.00 2000
Cr2N 0.0
-022000.0 00.00 12.00 2000
-022000.0 00.00 12.00 2001
CrN 0.0
-027000.0 -2.08 24.44 1800
-027000.0 -2.08 24.44 1801

EN 7.13
-063000.0 -4.40 37.55 2000
-063000.0 -4.40 37.55 2001
TaN 13.74
-058380.0 -1.76 25.51 2000
—058380.0 —1.76 25.51 2001
HfN 13.95

~088716.0 -2.44 30.39 1600
—088716.0 -2.44 30.39 1800
ZrN 14.42

-087000.0 00.00 22.31 1135
-087925.0 00.00 23.11 2000
Nb2N 0.0

-056750.0 00.00 22.50 0600
-056750.0 00.00 22.50 0601
Si3N4 44.11

-177000.0 -5.76 96.30 1686
-209000.0 00.00 96.80 1973
M02N 0.0

~017200.0 -4.60 28.95 1300
-017200.0 -4.60 28.95 1301
VN 10.82

-041360.0 -1.32 15.65 2000
—041360.0 -1.32 15.65 2001
TiCN 11.85

-062000.0 00.00 10.93 1155
-062725.0 00.00 11.81 1900

A.6 ExFreeEn.DAT
This stores the excess free energy of solution of the tool constituent atoms in a-iron.
It is therefore valid only for iron based work materials. Since 'X' is a dummy element, the

corresponding value is zero.

Ti -6900
N 5700
X 0
C 7600
Hf —2100
Si -16700
Zr -5000
V —9100
Nb —100
Ta -200

Cr 2200

72

MO -5000
W -7110
Al -10700
B -2100
0 -12600

A.7 Results of the computer program

Since in the experimental work carried out TiN, TiCN and A1203 coatings were used,
the results pertaining to only these three materials is presented here. Kramer and
Kwon[13] had used the bulk hardness of the materials in their wear prediction. However,
the presence of residual stresses in the thin film coating can markedly change these
values [31]. Hence the arduous task of procuring the hardness data of the coatings was
undertaken. While the hardness data for the TiN and the TiCN coatings was obtained
from Santhanam [32], the hardness data for the A1203 coating was obtained from Quinto
and Mehrotra [41]. Hardness of cementite used by Kramer and Kwon [13] was obtained
from Gove and Charles [43]. The hardness data were then curve fitted and used for the
theoretical wear prediction. The following sections will detail the output generated for the
various coatings. The thermodynamic data for TiN and A1203 was taken from Kramer
and Kwon[l3]. The calculation of the thermodynamic data for the TiCN coating is
detailed in Appendix B.

Table A. l: mechwear.out for the TiN coating

73

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Temperature Relative Abrasive Two body relative TlHardness NHardnees wear coefi. Wear eoeff.
(°C) wear wear (3-body) (2-body).
0 1 006100 1 .00E+00 2563.6 1200 1 .69E-08 3.90E-04
100 1.00E+00 1.00E+00 2184.6 1063 2.50E-06 4.58E-04
200 1.00E+00 1.005400 1861.8 941.6 3.70E-06 5378-04
300 1 .OOE+00 1 005100 1586.3 834.1 5.48E-06 8306-04
400 1 .00E+00 1.00E+00 1351.8 738.9 8.12E-06 7.4OE-04
500 1 .OOE+00 1 .OOE+00 1 151.9 576.4 5.61 E-06 8688-04
800 1 .OOE+00 1 .00E+00 981 .6 406.1 2105-06 1 .026-03
700 1 .00E+00 1 .OOE+00 836.5 286.1 7.88E-07 1 20an
800 1.00E+00 1.00E+00 712.8 201.6 2.95E—07 1.4OE-03
900 1 .OOE+00 1 .OOE+OO 607.4 142 1.1 1E-07 1 .655—03
1000 1.00E+00 1.00E+00 517.6 100.1 4.15E-08 1.93E-03
1100 1.00E+00 10064-00 441.1 70.5 1.56E-08 2276-03
1200 1 .00E+00 1.00E+00 375.8 49.7 5845-09 2.66E-03

Table A-2: mechwear.out for the TiCN coating

Temperature Relative Abrasive Two body relative TIHardnese NHardneee wear coefi. Wear ooeli.
('C) wear wear (it-body) (2-body).
0 5.57E-01 9.20E-01 2787.3 1200 9.4OE-07 3595-04
100 2.40E-01 8165-01 2678 1063 6.01 E-O‘I 3.73E—04
200 1 .04E-01 7.24E-01 2573 941.6 3.84E-07 3.89E-04
300 4.485-02 6.42E-01 2472.1 834.1 2.465-07 4.05E-04
400 1 .935-02 5.69E-01 2375.2 738.9 1 .57E-07 4.21 E-04
500 1 .95E-02 5.70E-01 2021 .9 576.4 1 .09E-07 4.95E-04
600 2.58E-02 5.93E-01 1655.4 406.1 5.428-08 6.04E-04
700 3.41E-02 6.17E-01 1355.3 286.1 2695-08 7388-04
800 4.51 E-02 6.42E-01 1109.6 201.6 1.33E-08 9.01 E-04
900 5.97E-02 6.69E-01 908.5 142 6616-09 1.1OE-03
1000 7.90E-02 6968-01 743.8 100.1 3.28E-09 1.34E-03
1 100 LOSE-01 7.24E-01 609 70.5 1.63E-09 1.64E-03
1200 1386-01 7.54E-01 498.6 49.7 8.07E-10 2.01 E03

Table A-3: mechwear.out for the A1203 coating

Temperature Relative Abrasive Two body relative TIHardness All-lardness wear coefi. Wear coefi.
('C) wear wear (Fr-body) (2-body).
O 1 .30E+00 1 .04E+00 2468.5 1200 2.20E—06 4.05E-04
100 1.32900 1.04E+00 2100.2 1063 3.30E-06 4.76E-04
200 1 .33E+00 1 0464-00 1786.8 941 .6 4938-06 5608-04
300 1 .35E+00 1 046400 1520.2 834.1 7396-06 6585-04
400 1 36800 1.05800 1293.3 738.9 1 .1 1E-05 7.73E-04
500 1 .38E+00 1 .05E+00 1 100.3 576.4 7735-06 9095-04
600 2.07E+00 1.11E+00 884.5 406.1 4.386-06 1.13E-03
700 3.11E+00 1.18E+00 711.2 286.1 2.45E—06 1.41 E-03
800 4.67E+00 1 .25E+00 571 .9 201 .6 1.38E-06 1 .75E-03
900 7.01 E+00 1.32E+00 459.9 142 7.76E-07 2.17E-03
1000 1 .05E+01 1 .4OE+00 369.8 100.1 4375-07 2705-03
1 100 1 .58£+01 1 .4BE+00 297.4 70.5 2.46E-07 3366-03
1200 2.37E+01 1.57E+00 239.1 49.7 1386-07 4.18E-03

 

 

 

 

 

 

 

 

 

 

74

Table A-4: chemdiss.out for the TiN coating

 

 

 

temp( K) vol.solubility vol.ref. solubility relative solubility J
373 2.12E-20 2.12E-20 1.00E+00
473 1 .67E-15 1.67E—1 5 1.00E+00
573 2.57E-12 2.57E—1 2 1.00E+00
673 4.47E-10 4.47E-10 1.00E+00
773 2.04E-08 2.04E-08 1.00E+00
873 3.90E—07 3.90E-07 1 .00E+00
973 4.05E-06 4.05E-06 1.00E+00

1073 2.72E-05 2.725-05 1.00E+00
1173 1.34E-04 1.34E-04 1.00E+00
1273 5.14E-04 5.14E-04 1.00E+00
1373 1.62E-03 1 .62E-03 1.00E+00
1473 4.36E-03 4.36E-03 1.00E+00
1573 1.04E-02 1 .04E-02 1.00E+00
1673 2.22E-02 2.22E-02 1.00E+00

 

 

 

 

Table A-5: chemdiss.out for the TiCN coating

 

 

 

 

 

 

 

 

 

 

 

 

temp( K) vol.solubility vol.ref.solubility relative solubility
373 2.12E-16 2.12E—20 1.00E+04
473 1 .42E-12 1 .67E-15 8.50E+02
573 4.38E-10 2.57E-12 1.71 E+02
673 2.46E-08 4.47E—1 0 5.51 E+01
773 4.88E-07 2.04E-08 2.39E+01
873 4.88E-06 3.90E-07 1.25E+01
973 3.04E-05 4.05E-06 7.51 E+00
1073 1.35E-04 2.72E-05 4.95E+00
1173 4.25E-04 1.34E-04 3.17E+00
1273 1.22E-03 5.14E-04 2.37E+00
1373 3.00E-03 1.62E-03 1.85E+00
1473 6.52E—03 4.36E-03 1.50E+00
1573 1 .28E-02 1 .04E-02 1.24E+00
1673 2.34E-02 2.22E-02 1.05E+00
Table A-6: chemdiss.out for the Al203 coating
temp( K) vol.solubility vol.ref.solubility relative solubility
373 2.25E-36 2.1215720 1.06E-16
473 6.29E-28 1 .67E-15 3775-1 3
573 1.97E-22 2.57E-12 7.65E-11
673 1 .43E-18 4.47E-10 3.19E-09
773 1 .03E-15 2.04E-08 5.05E-08
873 1 .65E-13 3.90E-07 4.23E-07
973 9.45E-12 4.05E-06 2.33E-06
1073 2.63E-10 2.72E-05 9.64E-06
1173 4.14E-09 1.34E-04 3088-05
1273 4.22E-08 5.14E-04 8.21 E—05
1373 3.07E-07 1 .62E—03 1.90 EM
1473 1 JOE-06 4.36E-03 3.90E-04
1573 7.58E-06 1.04E-02 7.33E-04
1673 2.83E-05 2.22E-02 1 .27E—03

 

 

 

 

Appendix B

COMPUTATION OF THE FREE ENERGY OF FORMATION OF TiCN

B.1. Balinit B coating of Balzers

The TiCN coating was a commercial coating designated Balinit 8'. Though it is
common to represent the compound as TiCN, in reality, it exists as TiC1.,,Nx depending
on the process parameters such as the N2 partial pressure in the PVD coating process.
Hence, the free energy of formation is a separate calculation for the purpose of

computing the dissolution wear.

It can be seen in [34] that a mixture of TiC and TiN in the required proportions can
form TiCN. It is also known that they exhibit very good miscibility in one another and
the resulting mixture shows a color according to the final stoichiometry. At a composition
of TiCo.5No_5 the compound is gray with a bluish tinge, as was the case for the coating in

this work. One can therefore assume reaction in equation B] for the formation of TiC]-

xNx. As a result, the free energy of formation of TiCMNx is given by equation B.2.

_)
(1 — x)(TiC) + xTiN (_ (TiCH N x ) Equation 8.1

AGJO (TiCMNx ) = (1 - x) AGf°(TiCN) + xAGf°(TiN) + RT(anx + (1 — x)Ln(1— x))
Equation 8.2
Essentially, this would mean a neglect of the excess free energy of solution of the

respective components in the formation of the final compound. Since, the exact

composition itself was speculative, as known from Balzers, Inc., this simple assumption

 

' Coutesy Balzers, Inc.

75

76

would suffice for the computation in this context. For a more detailed study one is

referred to [34], [3S] and [36].

B.2. Final Form for the Free energy of formation for TiCo,5N05
For a value of x = 05 , the formula for the free energy of formation (in Kcal/mol)

appears as: -

AG,°(Tic,_,N,_,) = -1.37728 T+O.5 (-43750+2.41 T) +0.5 (-80250+ 22.2 T) , T < 1155°c

AG,°(Tic,_,NO,) = -1.37728 T + 0.5(—44600 + 3.61 T) +
O.5(—80850 + 22.77 I), 1155° K < T < 1900° K

B.3. Molar Volume of TiC,-xN,
The molar volume can be obtained as a linear interpolation between the two
values. This is justified because both TiN and TiC exhibit the same crystal structure and

also because they have comparable molar volume values. Hence
MVTK.‘,_,N, = (1 " x)MVTi(.' + XMVTIN Equation 33

A linear interpolation for TiCo.5No,5 would therefore give rise to a molar volume

of 11.85 cm3/mol.

Appendix C

EXPERIMENTAL DATA

 

 

 

 

 

 

 

 

Insert Work Material Coating Material Wemperature ('C) Experimental A value Predicted
Designation Flank Volume Flank Volume
Wear Rate Wear Rate
(pm’Imin) (pmzlmin)
1018(sp) TiN
xnf10186p-2 893.26 2.86 2.84£+07 3.36
xnf101Bsp-3 1 104.70 3.27 0.42
xnf10183p-1 1176.60 1.09 0.21
xnf10186p-4 1230.10 2.46 0.12
TiCN
xca10189p-1 863.99 1.65 2.06E+08 1.75
xca10183p-2 1041.40 0.40 0.51
xca10189p-3 1 103.80 0.86 0.33
xca10186p-4 1198.20 0.69 0.17
AI203
xoa10189p-2 914.87 12.14 1.64E+07 11.70
xoa1O189p-1 923.92 9.29 11.11
x061 0189p-4 996.56 8.01 7.31
xoa1018sp-3 1054.90 7.14 5.23
1045(sp) TiN
xng1045sp-1 860.53 4.80 3.01 E+07 4.92
xng1045sp-2 911.54 1.93 2.98
xng104539-4 991.37 3.84 1.36
xng10459p-3 1000.60 1.52 1 .25
, TiCN
xcb1045cp-2 816.93 1.44 1.63E+08 1.93
xcb10459p-1 851.68 1.02 1.51
xcb10459p-3 874.53 1 .74 1 .29
xcb10459p-4 f 997.86 2.61 0.54
} AI203
xob10456p-1 841.48 13.81 1.24E+07 13.45
xob10453p-4 913.39 6.26 6.90
xob10453p—2 914.74 10.10 8.83
xob10456p-3 1 933.69 6.65 7.91

 

 

 

 

 

 

77

 

78

 

 

 

 

 

 

 

 

 

 

RISE-Designation Work Material 1 Coating Material WEperimental A value W
Temperature (°C) Flank Volume Flank Volume
Wear Rate Wear Rate
(pmzlmin) (pmzlmin)
1065(sp) TiN
xnh10659p-2 697.61 12.69 1.23E+08 13.99
xnh10653p-3 1002.40 7.03 5.01
xnh10653p-1 1065.70 4.70 2.69
xnh10653p-4 1125.30 3.27 1.50
TiCN
xcc1oessp-1 612.96 2.71 2.47E+06 3.01
xoc10659p—2 911.71 1.07 1.51
xcc10658p-4 1037.60 2.10 0.62
xccioessp—s 1057.10 1.81 0.54
1111203
xoc10656p-1 603.55 23.04 1.77e+07 23.94
xoc1065ep-2 671.91 17.34 16.16
xoc1065ep-4 Third cut with 1 990.50 6.47 6.17
A1203 on 1065(sp)
invalid. .
1095(sp) TiN
xni10953p-1 736.32 6.97 94.26
xni109ssp-2 667.06 24.69 1.71E+08 26.15
xni10953p-3 1079.20 11.72 3.26
xni1095sp-4' 1230.60 7.90 0.74
TiCN
xod10953p-1 603.92 1.40 1 .63E+08 2.1 1
xod1oessp-2 699.67 1.37 1.06
xcd1095sp-3 1042.60 2.40 0.40
xcd1095sp-4 1096.60 1.64 0.27
A120.
xod10953p-1 " 736.36 29.11 1 .38E+07 27.41
xod10953p-2 601.79 13.06 16.61
xod109ssp-3 937.46 14.69 6.62
xod10959p-4 966.57 6.44 7.21

 

 

79

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

fibuignation Work Material Coating Material Fla—nk——-Experimental A value med—
Temperature (°C) Flank Volume Flank Volume
Wear Rate Wear Rate
(umzlmin) (pm’lmin)
1018(unsph) TiN
xnl1018un—2 857.95 5.65 1549.90 2.39
xnl1018un-1 858.39 1.41 2.39
xnl1018un-4 876.11 3.14 2.46
xnl1018un-3 994.23 0.57 2.97
TiCN
xcl1018un-1 782.49 1.42 663.83 0.58
xcl1018un-2 878.32 1.36 0.70
xci1018un-3 920.70 1.03 0.76
xcl1018un-4 1165.70 0.32 1.24
Al203 1
xol10180n-1 805.01 4.10 3291.60 5.82
xol10180n-2 839.64 5.31 6.28
xol1018un-3 887.68 8.50 6.97
xol1018un-4 904.75 7.99 7.23
1045(unsph) TiN
xnm1045un-2 969.39 4.66 2430.74 4.47
xnm1045un-4 1056.00 2.84 5.14
xnm1045un-3 1057.60 10.82 5.15
xnm1045un-1 1114.50 2.41 5.64
TiCN
xom1045un-1 945.02 0.65 325.98 0.39
xom1045un-2 1009.50 0.77 0.45
xom10450n-3 1073.40 0.51 0.51
xom10450n-4 1223.60 0.33 0.69
AI203
First cut
interrupted due to 1
BUE ‘ 1
xom1045un-4 978.28 6.43 3 2221.19 . 5.73
xom1045un-3 1023.60 6.69 1 6.32
xom1045un-2 1205.50 6.71 ' 9.36

 

 

80

 

 

 

 

Insert Work Material Coating Material Wamperature ('C) Experimental A value PTediEteT
Designation Flank Volume Flank
Wear Rate Volume
(pm’lmin) Wear Rate
_ (um’lmin)
1070(unsph) TiN

xnn1070un-4 834.19 2.78 1235.24 1.83

xnn1070un-3 862.14 1.20 1.92

xnn10700n-1 a 962.26 1.50 2.25
xnn1070un-2 1 1134.40 3.41 2.96

TiCN
xcn10700n-2 ; 1020.70 1.25 624.51 0.67
xcn1070un-4 1 1021.40 1.24 0.66
xcn1070un-1 1025.30 0.59 0.88
xcn1070un-3 1 1101.40 0.64 1.03
Al203 1

xon10700n-3 660.04 6.44 3521.31 7.02
xon1070un-2 919.23 11.27 7.96
xon1070un-4 937.11 6.78 6.30
xon1070un-1 1 937.36 7.16 6.31

 

 

 

 

 

 

 

 

 

 

 

81

 

 

 

 

 

 

Insert Work Coating Cutting Speed Crater Experimental A «65"m’
Designation Material Material (m/min) Temperature Crater Wear Values Crater
('0) Rate Wear
(um/min) Rate
(pm/min)
1018(sp) TiN
xnf10185p-2 164.30 1037.40 2.81 587451.93 2.79
xnf1018sp-3 239.57 1274.96 0.29 1.09 0.57
xnf10185p-1 99.29 1355.76 1.84 0.33
xnf10185p-4 274.09 1415.66 0.43 0.88
TiCN
xca1018sp-1 101.32 1004.52 0.42 169605767 0.56
xoa1018sp-2 . 164.05 1203.65 0.64 1.31 0.33
xoa1018sp—3 3 274.44 1273.97 0.59 0.49
xca10185p-4 215.81 1360.03 0.52 0.63
110,03
xoa10185p-2 1 164.54 1061.69 0.96 19553.11 1.01
X08101889-1 1 102.06 1071.85 0.36 16425.43 0.60
x0810188p-4 207.55 1153.47 1.20 0.91
x0810188p-3 1 292.06 1219.02 1.36 1.44
1045(sp) 1 TiN
th10458p-1 100.06 1000.63 0.11 31276.26 0.20
xn910458p-2 . 162.36 1057.94 0.35 16.13 0.30
xng10458p-4 269.04 1147.64 0.99 0.78
xng1045sp-3 225.66 1 156.01 0.53 0.75
TiCN
xcb10458p-2 1 164.33 951.64 0.16 94713.57 0.25
xcb1045sp-1 1 100.67 990.69 0.16 21.56 0.25
xcb10458p-3 227.44 1016.36 0.63 0.47
xcb10456p-4 291.11 1154.93 1.67 1.69
1 Al203
xob10455p-1 1 100.14 979.22 0.73 16200.42 0.93
xob10458p-4 1 1 292.11 1060.02 2.20 169999.46 1.96
X0b10458p-2 , 1 164.92 1061.54 1.03 1.17
xob10459p-3 1 227.66 1063.06 1.50 1.56

 

 

 

 

 

 

 

 

 

82

 

 

 

 

 

 

 

 

Insert Work Material Coating Material Cutting Speed Crater Experimental A and PW
Designation (mlmin) Temperature Crater Wear Values d
('C) Rate Crater
(pm/min) Wear
Rate
(pm/min)
1065(sp) TiN
xnh10656p-2 170.22 1042.29 1.00 288673.03 1.43
xnh10656p-3 219.21 1160.03 2.08 8.96 0.93
xnh106590-1 103.18 1231.16 0.58 0.71
xnh1065sp-4 261.37 1298.12 1.88 2.05
TiCN
xoo10656p-1 103.88 947.18 0.19 720387.67 0.44
xoc10656p-2 157.73 1058.13 0.86 12.78 0.55
xoc1065sp-4 241.14 1199.81 1.82 1.46
xoc1065sp-3 218.50 1221.49 1.25 1.62
Ales
xoc10656p-1 105.56 936.61 0.61 14281.05 0.95
xoc1065sp-2 157.66 1013.42 1.31 75382.09 0.95
xoc10656p-4 Third out with 257.82 1146.66 1.45 1.47
A1203 on
1065(sp) invalid. .
1095(sp) TiN
xni10955p-1 104.24 861.07 1.07 99912.14 1.69
xnl10956p-2 168.79 1007.97 2.08 6.48 0.69
xni10959p-3 231.48 1246.33 2.73 0.83
xni10953p-4 277.21 1416.66 4.29 4.60
TiCN
xod1095sp-1 105.04 937.02 1.04 2047747.12 1.23
xod10956p-2 170.14 1044.83 1.73 21.99 1.31
xodiO956p-3 231.48 1205.20 2.05 2.70
xod109560-4 279.45 1268.12 4.91 4.61
A1203
xod10959p-1 107.73 861.11 1.05 12931.57 1.36
xod10953p-2 171.11 934.63 1.79 667761.79 1.47
xod109550-3 234.40 1087.07 2.69 2.76
xod10955p-4 274.74 1122.02 5.14 5.11

 

 

 

 

 

 

 

 

 

83

 

 

 

 

 

Insert Work Material Coating Cutting Speed Crater Experimental A and PM
Designation Material (tn/min) Temperature Crater Wear Values Crater Wear
(°C) Rate Rate
(pm/min) (pm/min)
1018(unsph) TiN

xnl1018un-2 154.40 997.73 2.73 173944.51 1.15
xnl1018un—1 90.90 998.22 0.89 1.10 0.67
xnl1018un-4 275.60 1018.13 0.50 1.68
xnl1018un-3 214.30 1150.85 0.43 0.39

TiCN
xci1018un-1 91.00 912.94 0.07 769522.06 0.43
xcl1018un-2 153.00 1020.62 0.38 1.63 0.36
xcl1018un-3 214.00 1068.24 0.83 0.38
xci10180n-4 275.70 1343.52 0.61 0.64

Al203
xol1018un-1 92.00 938.25 0.12 947.45 0.09
xol1018un-2 153.30 977.16 0.10 536148.45 0.19
xol1018un-3 214.80 1031.13 0.70 0.54
xol1018un-4 1 275.30 1050.31 0.79 0.87

1045101151311)" - TIN .. .-., .. .. ... _ _

xnm1045un-2 1 164.30 1122.94 0.74 575105.30 1.35
xnm1045un-4 ‘ 274.50 1220.26 2.56 " 7.32 i 1.44
xnm1045un-3 221.90 1222.06 1.19 1.23
xnm1045un-1 91.60 1285.99 0.14 0.93

TiCN
xcm1045un-1 91.60 1095.56 0.05 940325.38 0.26
xcrn1045un-2 152.60 1168.01 0.37 4.63 0.43
xom1045un-3 220.60 1239.81 1 .12 0.77
xom10450n-4 276.70 1408.57 2.68 2.74

A80:

First cut
. interrupted
due to BUE

xom10450n-4 3 284.10 1132.93 2.33 41731.31 2.42
xom1045un-3 221.00 1163.65 1.58 1202.08 1 1.42
xom1045un-2 152.20 1388.24 1.09 1 1.10

 

 

 

 

 

 

 

 

 

 

 

84

 

 

 

 

 

 

nsert Work Coating Cutting Speed Crater Experimental A and B Pmdicted
Designation Material Material (mlmin) Temperature Crater Wear Values Crater
('0) Rate (pm/min) Wear
. Rate
1 (pm/min)
1070(unsph) 1 TiN .
xnn1070un-4 l 272.40 971.03 1.17 101401.29 1.55
xnn1070un-3 1 214.50 1 1002.44 1.34 5.32 0.92
xnn1070un-1 i 90.30 1114.96 2.00 0.21
xnn1070un-2 152.30 1306.35 0.61 0.96
' TiCN 1
xcn1070un-2 152.60 E 1160.60 0.61 643672204 1.42
xcn1070un-4 271.90 1 1161.36 3.02 3.33 2.42
xcn1070un-1 92.00 1165.76 0.15 0.66
xcn1070un-3 214.00 ‘ 1271.27 1.71 1.47
A1203

xon1070un-3 213.50 1000.06 2.37 24944.36 2.37
xon1070un-2 152.50 1066.56 0.65 99651.56 1.29
xon1070un-4 ‘ 275.00 1066.67 2.65 2.13
xon1070un-1 1 92.90 1086.96 0.46 0.80

 

 

 

 

 

 

 

 

85

 

 

 

 

 

 

 

 

 

 

 

 

Wnafion Work Material Coating Material Emmett Crater Area W
(m/miIIl (mm’) (J) 1371
1018(sp) TiN

xnf10189p-1 99.29 3.45 120.0

xnf10183p-2 164.30 4.13

xnf1018ep-3 239.57 2.95 Average Area =

xnf1018sp-4 274.09 2.51 3.26
TiCN

xca10189p—1 101.32 3.81

xoa1018ep-2 164.05 3.24

xca10186p-4 215.81 3.28 Average Area 8

xca10186p-3 274.44 2.85 3.30
A1203

xoa10186p-1 102.06 3.60

xoa10189p-2 164.54 3.49

xoa10186p-4 207.55 3.36 Average Area =

xoa10189p-3 292.08 2.86 3.33

1045(sp) TiN

xng10456p-1 100.08 2.99 30.6

xng10459p-2 162.38 2.87

xng10459p-3 225.68 2.23 Average Area =

xng10453p-4 289.04 2.17 2.57
TiCN

xcb10455p-1 100.87 3.31

xcb10453p—2 184.33 3.06

xob10456p-3 227.44 2.49 Average Area =

xcb10456p-4 291.11 2.17 2.76
Al203

xob1045ep-1 100.14 3.45

xob10456p-2 164.92 2.98

xob1045sp-3 227.88 2.76 Average Area =

xob10455p-4 292.1 1 2.74 2.98

 

 

86

 

 

 

 

 

 

 

 

 

 

 

 

 

Wmignation Work Material Coating Material 1 Cutting Speed Crater Area WI
1 (m/min) (mm’) (J)
1065(sp) TiN 1
xnh10656p-1 103.16 2.55 10.0
xnh10656p-2 J 170.22 2.66
xnh10656p-3 219.21 2.45 Average Area =
xnh10655p-4 i 261.37 2.14 2.50
TiCN 1
xoc1065sp-1 103.88 3.16
xoc10653p-2 157.73 2.07
xoc10655p-3 218.50 2.16 Average Area =
xoc10653p-4 241.14 2.74 2.53
Al203
xoc10656p-1 105.56 3.21
xoc10655p-2 1 157.66 3.12 Average Area =
xoc10656p-4 1"T'ii'i‘rmi1—ATZCJ06'16651W3Fd'.""1 257.82 2.30 2.88
' 1095(sp) nN
xni10956p-1 104.24 3.47 2.7
xni10959p-2 168.79 2.52
xni1095sp~3 1 231.48 1.92 Average Area =
xni10958p-4 277.21 2.13 2.51
TiCN
xcd10956p-1 1 1 105.04 2.78
xod1095sp-2 1 1 170.14 2.33
xod10955p-3 1 231.48 2.01 Average Area =
xod1095sp—4 1 279.45 2.01 2.26
AlgO;
xoo109ssp-1 1 107.73 3.56
xod1095sp-2 1 171.11 2.60
xod10955p-3 234.40 2.47 Average Area =
xod10956p-4 ' 274.74 2.54 2.64

 

 

87

 

 

 

 

 

 

 

 

 

 

 

 

 

Wsignation Work Material Coating Material 1 Cutting Speed Crater Area 1W? 1
1 (m/min) (mmz) (J)
1018(unsph) TiN 1
xnl1018un-1 ‘ 90.90 3.69 *m'"
xnl1018un-2 154.40 3.21
xnl1018un-3 214.30 2.17 Average Area =
xnl1018un-4 275.60 2.36 2.86
TiCN
xcl1018un-1 91.00 3.56
xcl1018un-2 153.00 2.75
xcl1018un-3 214.00 2.41 Average Area =
xcl1018un-4 275.70 2.01 2.68
AI203
xol1018un-1 92.00 3.49
xol1018un-2 153.30 2.97
xol1018un—3 1 214.80 2.54 Average Area =
xol1018un-4 1 275.30 2.35 2.84
1 1045(unsph) TiN
xnm1045un-1 ‘ 91.60 2.02 "'mm'
xnm1045un-2 164.30 2.59
xnm1045un-3 221 .90 2.03 Average Area =
xnm1045un-4 274.50 2.45 2.27
TiCN
xcm1045un-1 91 .60 2.75
xcm1045un-2 152.60 2.34
xom1045un-3 13 220.60 2.06 Average Area =
xcm1045un-4 1 276.70 1.82 2.25
1 AI203
1 First cut interrupted
due to BUE
xom1045un-2 152.20 2.91
xom1045un-3 221.00 2.49 Average Area =
xom1045un-4 284.10 2.46 2.62

 

 

88

 

 

 

 

 

 

 

 

TNTDesignation Work Material Coating Material Cuttinngpeed Crater Area W1
(mlmin) (mm’) (J)
1070(unsph) TIN
xnn1070un-1 90.30 2.99 “”
xnn1070un-2 152.30 1.59
xnn1070un-3 214.50 2.28 Average Area =
xnn1070un-4 272.40 2.26 2.28
TiCN
xcn1070un-1 92.00 1.90
xcn1070un-2 152.80 1.66
xcn1070un-3 214.00 1.53 Average Area =
xcn1070un-4 271.90 1.45 1.64
Al203
xon1070un-1 92.90 2.54
xon1070un-2 152.50 2.25
xon1070un-3 213.50 2.14 Average Area =
xon1070un-4 275.00 1.99 2.23

 

 

Appendix D

CURVE FITTED PEAK TEMPERATURE

D.1 Inverse Estimation Scheme for the Flank Temperature

IR pyrometry was adopted to measure the temperature of the tool at a fixed point on
the rake face. One criticism of the method has been the effect of the Chip coming in the
field of view of the pyrometer [3 8]. Previous attempts at this approach were made by
Wang et al. [42]. However, adopting the assembly shown in Figure 4.3 resolved this
problem. To estimate the cutting interface temperature, Yen and Wright [26] suggested a
method using a remote sensing technique. Since the coating thickness is exceedingly
small to offer any thermal resistance, the same model was capable of estimating the
interface temperature [39]. The flank temperature can be estimated using the scheme
suggested by Oxley [29], which states that the tool work interface temperature is 0.82-
0.95 times the mean chip-tool interface temperature. The average value of 0.89 times the
chip-tool interface temperature was used for this study (All temperatures being in
Kelvin).

Determining the steady state cutting temperature from the knowledge of the transient
temperature at a point remote to the interface involves solution of the 3-D transient heat
conduction problem. Since the time derivative in the heat diffusion equation is of the first
order, it was inferred that the time growth in the remotely measured temperature is of the

form in equation D. 1.

7(1) = Ti + 70(1— e‘a’) Equation DJ

89

90

where T; is the initial temperature in the temperature history of the pyrometer and To and
a are constants derived from a curve fitting process on the temperature data of the
pyrometer. Then T ,-+ To would represent the steady state temperature of the pyrometer.
The temperature measurements showed less than 2% difference between the empirical
result and that obtained from the curve fitting process. In addition, the difference between
the steady state temperature obtained by the above equation and the peak temperature

measured by the pyrometer was less than 7°C.

The 1-D ellipsoidal mapping model simplifies the inverse problem of estimating the
interfacial temperature [26]. The modeling and the governing differential equations are
discussed in the referred paper. The final form of the inverse solution [40] can be

expressed as in equation D.2.

T

 

Equation D.2

 

Tr _
where

T, is the steady-state cutting interface temperature

T is the steady state remotely measured rake face temperature

Tao is the temperature in the far field of the tool (Taken to be 25°C)

x is the distance of the point of measurement from the origin of the axes

a is the radius of the circular tool-chip contact area.

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