A quantitative investigation of tool crater wear was carried out in free cutting steels with and without lead addition (commercial grade AISI 12L 14 and AISI 1215 respecfively ) at moderately high cutting speeds ( 140-200 m min - I ) using cemented carbide cutting tools. Crater wear was quantitatively measured by determining the amount of tungsten carried into the chips using instrumental neutron activation analysis. The bulk of tungsten in the chips occurs as soluble tungsten dissolved in the steel matrix rather than as tungsten carbide confirming that dissolution of the tool into the workpiece is the dominant mechanism of tool crater wear. Experimental results have confirmed that lead decreases the cutting force and the contact length but is ineffective in suppressing tool dissolution wear. Since dissolution of the tool occurs by a diffusion mechanism, it should be possible to design a diffusion barrier at the tool-chip interface to suppress dissolution wear. It is demonstrated that deformable oxide inclusions (CaO--AI203-2SiO2) engineered into the workpiece (AIS! 1215 IE) form a glassy layer at the tool-chip interface that suppresses dissolution wear. Alternatively a HfN coating pot on the tool acts as an effective diffusion barrier, as the solubility of HfN is seven orders of magnitude ( 10 million times) less than that of tungsten carbide in the austenite phase of the steel at the tool-chip interface temperature. Thus, inclusion engineering of the workpiece and coating of the tool are identified as two viable and attractive options to replace lead in free cutting steels. Theoretical analysis of the above experimental observations constitutes the subject of Section 4. The effect of tribology of seizure occurring at higher cutting speeds on the tool-chip interface temperature is analysed using finite element modelling. The shear flow of the chip material under the compressive stress of the seized region is described using Bowden and Tabor's equation. The effect of temperature distribution of the seized region on the di ffusional transport is analysed. A comparison of the experimentally measured tungsten transported to the chip with the theoretical prediction suggests that an enhanced diffusion operates at the tool-chip in~fface. High diffusivity paths contribute to an enhancement in the diffusion coefficient that is two orders of magnitude greater than the lattice diffusion coefficient.

1. ELSEVIER Wear197(1996)45-55
WEAR
Micromechanisms of tool wear in machining free cutting steels
K. Ramanujachar, S.V. Subramanian
DepartmentofMaterialsScienceandEngineering.McMasterUniversity.Hamilton.Ontario.Canada
Received21 September1994;accepted5 October1995
Abstract
A quantitative investigation of tool crater wear was carried out in free cutting steels with and without lead addition (commercial grade
AISI 12L14and AISI 1215 respecfively) at moderately high cutting speeds ( 140-200 m min- I) using cemented carbide cutting tools. Crater
wearwasquantitatively measured by determining the amount of tungsten carried into the chipsusinginstrumental neutron activation analysis.
The bulk of tungsten in the chips occurs as soluble tungsten dissolved in the steel matrix rather than as tungsten carbide confirming that
dissolution of the tool into the workpiece is the dominant mechanism of tool crater wear. Experimental results have confirmed that lead
decreases the cutting force and the contact length but is ineffective in suppressing tool dissolution wear.
Since dissolution of the tool occurs by a diffusion mechanism, it should be possible to design a diffusion barrier at the tool-chip interface
to suppressdissolution wear. It is demonstrated that deformable oxide inclusions (CaO--AI203-2SiO2) engineered into the workpiece (AIS!
1215 IE) form a glassy layer at the tool-chip interface that suppresses dissolution wear. Alternatively a HfN coating pot on the tool acts as
an effective diffusion barrier, as the solubility of HfN is seven orders of magnitude ( 10 million times) less than that of tungsten carbide in
the austenite phase of the steel at the tool-chip interface temperature. Thus, inclusion engineering of the workpiece and coating of the tool
are identified as two viable and attractive options to replace lead in free cutting steels. Theoretical analysis of the above experimental
observations constitutes the subject of Section 4.
The effect of tribology of seizure occurring at higher cutting speeds on the tool-chip interface temperature is analysed using finiteelement
modelling. The shear flow of the chip material under the compressive stress of the seized region is described using Bowden and Tabor's
equation. Theeffect oftemperature distribution ofthe seized regiononthediffusionaltransport isanalysed. Acomparison oftheexperimentally
measured tungsten transported to thechip with the theoretical predictionsuggests that anenhanced diffusion operates atthetool-chip in~fface.
High diffusivity pathscontribute to an enhancement in thediffusion coefficientthat is twoorders of magnitude greater than the lattice diffusion
coefficient.
Keywords: Toolwear;Freecuttingsteels;Diffusionalwear;Inclusionengineering;Coating;Machinability
1. Introduction
A quantitative understanding of the micromechanisms of
tool crater wear and chip formation during the machining of
free cutting steels is essential in order to design non-toxic
free cutting steels as a replacement for toxic leaded free cut-
ting steels. The functional role of lead in free cutting steels
used at low cutting speed with high speed steel tools is well
established [ 1]. Lead inclusions lower the critical accumu-
lated damage that aids ductile fracture in chip formation. At
low cutting speeds, the tribology of sliding wear operates at
the 'tool-chip interface, where lead lowers the tool-chip inter-
face friction. Shaw et al. have carried out extensive research
over the years on free cutting steels. They have shown that
lead lowers the cutting force and the tool-chip contact length.
Lead forms a fluid layer at the tool-chip interface that lowers
the coefficient of friction. Lead increases the shear plane
0043.16481961515.00 © 1996ElsevierScienceS.A.Allrightsreserved
SSD10043-1648(95)06810-4
angle [ 1]. At low cutting speeds the tribology operating at
the tool-chip interface is one of sliding. This point has been
confirmed by Doyle and Tabor [2]. In their work. planing
experiments were conducted with transparent sapphire tools
on copper, indium, aluminium and lead at cutting speeds
ranging from 5 mm s- t to 500 mm s- t. Relative motion
between the chip and the tool was inferred. At the high cutting
speeds used in CNC machining with carbide tools, the tri-
bology of seizure operates at the tool-chip interface [3].
However the micromechanisms of tool wear are not well
established. The objective of this work is to elucidate the
micromechanisms of tool wear in machining free cutting
steels with and without lead during moderately high curing
speeds using cemented carbide cutting tools and to identify
design strategies to eliminate lead in free cutting steels.
In this work the micromechanisms of tool crater wear arc
investigated quantitatively during the machining o.~an AISI

2. 12L14 and AISI 1215 grades of free cutting steels using
tungsten carbide tool at various cutting speeds. The total
amount of tungsten transported from the tool to the chips is
determined by neutron activation analysis. The area of tool-
chip contact was experimentallydetermined as a function of
cutting speed and the temperature at the tool--chip interface
was derived from the cutting force measurements. The
amount of tungsten transported by diffusion into the chip for
the measured tool-chip contact area and tool--chip interface
temperature derived from cutting force measurement was
compared with the total amount of tungsten carried into the
chlps. By choosing a coating (HfN) which has the least
ti',ermodynamicpotential for dissolution into the workpiece,
it is shown that the amountof dissolutionwear can be reduced
drastically. It is demonstrated that deformable oxide inclu-
sions can be engineered into free cutting steels that suppress
dissolution crater wear. Theoretical analysesare carried out
1. to relate the contact length to parameters relating to the
mechanicsof metal cutting and material parameters of the
workpiece,
2. to predict the tool--chipint.~rfacetemperatureof the seized
region using thermal finite element modelling, and
3. to predict the diffusional transport of tungsten from the
tool across the tool--chip interface.
2. Experimental
2.1. Cutting conditions forfree cutting steels AlSl l2Ll4
andAlS11215
The experimental programme involved a comparative
study of three free cutting steels: AISI 12L14 (leaded free
cutting steel), and AISI 1215 (non-leaded free cutting
steel), AISI 1215 IE (deformable oxide inclusionengineered
non - leaded free cutting steel). The cutting tool is cemented
tungsten carbide K-i tool (85% WC, 4% TaC-NbC, 11%
Co) in the uncoated and HfN coated condition. The compo-
sition of these steels used in the study are summarised in
Table 1.
The metal cutting conditions used in this study are:
depth of cut= 1.27 mm, feed rate=0.2 mm rev-~, rake
angle = + 5°, cutting speed varied from 140 to 200 m min - ~,
triangular insert (Kennametal TPG 432), clearance
angle= 11°, nose radius =0.794 mm ( 1/32 in), no coolant
used.
K.Ramanujachar,S.V.SubramanianI Wear197(1996)45-55
2.2. Neutron activation analysis
Tool crater wear was quantitatively measured by deter-
mining the tungsten content of the chips generated during
machining, using neutron activation analysis. All the chips
generated at a given cutting speed for a given time were
collected and irradiated in a neutron flux of 10t2 neu-
trons cm- 2s- ~..Thetotal amountof tungstencarried into the
chips was determinedusingstandardsof known composition.
The total amountof tungstenis made up of tungstenas atomic
tungsten and tungsten as tungsten carbide. The chips were
then dissolved in concentrated hydrochloric acid and the
tungsten carbide retained in the undissolved residue deter-
mined. The tungsten in solution in the iron matrix was taken
as the difference between the total tungsten in the chips and
the tungsten tied up as tungsten carbide [4].
2.3. Force measurements
Forces exerted on the tool were measured in order to cal-
culate the tool- chip interface temperatures. Cutting force,
feed force and radial force measurements were carried out
using a piezoelectric dynamometer. The signals acquired
from the dynamometer were processed using a computer.
2.4. Confactareameasurements
Tool - chip contact areas for increasing cutting durations
were measured from photographs of the tool tips. A magni-
fication of 50 X was used.The contact length was determined
from these photographs.
Fig. 1 shows a typical optical micrograph of the tool rake
face in which the contact length can be clearly distinguished.
Optical micrographs were used to measure the stick region
of the contact length involving seizure.
Table I
CompositionofA.I.S.I.12LI4,A.I.S.I.1215,inclusionengineeredsteel
Grade C (%) Mn (%) P (%) S (%) Si (%) Pb (%)
12L14 0.07 1.05 0.08 0.3 0.003 0.15
1215 0.08 0.95 0.08 0.3 0.006 0
1215 I.E. 0.096 1.08 0.06 0.32 0.18 0 Fig. I.An opticalmicrograph ofrakefaceoftool,showing theseizedregion
ofcontactlength.

3. K.Ramanujachar,S.V.Subramanian/ Wear197(1996)45-55
Table2
Neutronactivationanalysisoftungsteninchipsgeneratedduringmachiningfreecuttingsteelswithcementedcarbidetools
Grade Tool Speed(m rain-t ) W" total (ppm) WasWCt,(ppm) SolubleW~(ppm)
12LI4 KI 160 6.3+0.1 0.31 5.99
12LI4 KI 200 7.14-0.1 0.16 6.94
1215 KI 160 6.84-0.1 0.27 6.53
1215 KI 200 7.3::1:0.I 0.22 7.08
1215LE KI 160 0 0 0
12151.E KI 200 1.4+ .25 0.2 i.2
1215 KI CHfN 160 0.74-.2
1215 KI CHfN 200 1.64-.3
aCraterwear.
bMechanicalwear.
cChemicalwear.
Percentagemechanicalandchemicalwear
Grade Tool Speed Percent Percent
(m rain- :) chemical mechanical
12L14 K-I 160 95 5
12L14 K-I 200 98 2 p,
1215 K-I 160 96 4
1215 K-I 200 97 3 *~
1215I.E. K-I 200 86 14 ~ ~l
3. Results
3.1. Experimental results onfree cutting steels (AISI
12L14, AIS! 1215,1215 IE)
Table 2 summarizes tungsten content of the chips collected
over a duration of 15 s at two cutting speeds (160 and
200 m min- ~) for unleaded steel (AISI 1215), leaded steel
(AISI 12L!4) and inclusion engineered steel ( 1215 IE). Of
the total tungten picked up by the chip, the amount oftangsten
contributed by toogten carbide was quantitativelydetermined
and the difference is taken as the dissolved amount that is
atomicallytransported from the tool into the chip. The amount
of tungten dissolved in the chip is taken as a quantitative
measure of chemical wear. The amount of tungsten present
as tungsten carbide particles is considered as a quantitative
measure of physical or mechanical wear. Table 3 summarises
the percentage mechanical and chemical wear for uncoated
cemented carbide tool during the machining of three free-
cutting steels. Clearly the chemical wear dominates over
mechanical wear in each case, constituting in excess of g5%
of wear. Cutting experiments were performed with uncoated
cemented carbide K-1 tools and with HtN coated K-I tools
and the results are compared. Fig. 2 is a bar chart that shows
the increase in total tungsten concentration in the chip (total
crater wear), tungsten dissolved in the chip (chemical or
dissolution wear) and tungsten present as WC (mechanical
wear) in the chip during machining three free cutting steels
with uncoated tungsten carbide (K-I) tool at 160 m rain- J.
A~,.t~14' .,um.121efAlSl-121Sa)'l t~l~V6Comd toc~j
[mToT~wm~w mwc ]
WEAR' WEAR WEAR
Fig. 2. The bar chart for each steel shows the incw.asein total tungsten
concentration(crater wear), tungstendissolved (dissol,tion wear), and
tungstenpresentas WC (mechanicalwear) in the chipduring machining
the steelwithK-I toolata cuttingspeedof 160m ~in- t Theresultsfrom
neutronactivationa:talysiscomparetoolwearforthe~ ofleaded(AISI
121.14) and non-lcaded (AISI 1215), inclusionengineeredAIS! 1215
(I.E.) d~,ring machiningwith uncunledK-I tool and a Hff~co~ed tool
duringmachiningAIS11215.
The effect of HfN coating on the tool on machining unleaded
AISI 1215 on tool wear as measured by tungsten concentra-
tion in the chip is shown for comparison. Fig. 3 compares the
neutron activation results for the above case~ at a higher
cutting speed of 200 m rain- :. Clearly chemical or dissolu-
tion wear dominates at both cutting speeds in both leaded and
unleaded steel, However there is a ~.-astic reduction in tool
dissolutionwear both in the case of inclusionengineered steel
and coated tool. Clearly engineeringdeformable oxide inclu-
sions into the work piece is very effective in reducing tool
crater wear. Coating the tool with Ht'N is also very effective
in suppressing dissolution wear.
Table 4 summarizes the neutron activationanalysis results
for Hi content of AISI (1215) steel chips collected over a
duration of iSs at two cutting speeds (160 and
200 m rain- t).
Table 5 summarizes experimental data for free cutting
steels comprising of cutting force, contact length, the shep,r
angle, the ratioof feed force to cutting force and the maximum

4. ;i 200 M/MIN
K.Ramanujachar.5.1/.Subramanian/ Wear197(1996)45-5.5
lance over the seized contact length obtained from the finite
element analysis outlined in Appendix B.
A
I with HfN Coated tool)
imToT~.wm ~w rowe I
c~Nn o,~L~nON~EC.A.,eAt
WEAR WEAR WEAR
Fig. 3. The bar chart for each steel showsthe increasein total W (crater
wear), W dissolved(dissolutionwear) andWproductasWC (mechanical
wear) inthechipduringmachiningthesteelwithK-I toolata cuttingspeed
of200 m rain- 1.The resultsfromneutronactivationanalysiscomparetool
wearforthe casesof leaded(A1SI12LI4) andnon-leaded (AISI 1215),
inclusionengineeredA1SI1215(I.E.) duringmachiningwithuncoatedK-
I toolanda Ht'NcoatedtoolduringmachiningAISI 1215.
Table4
Neutronactivationanalysisof Hf in chips duringmachiningA.I.S.I 1215
steelwithHfNcoatedcementedcarbidetool
Grade Tool Speed Concentration3fHf
(m rain-~) (ppm)
1215 K-I + Hfl~ 160 0.414-0.1
1215 K-I + HfN 200 0.614-0.51
temperature at the tool-chip interface calculated using
Boothroyd's model [5]. For comparison, theoretically pre-
dicted contact lengths are also included in Table 5. Fig. 4 is
a schematic of elemental configuration used for thermal finite
element .malysis. Fig. 5 is a typical temperature distribution
along the tool- chip interface expressed as fractional dis-
Table5
4. Discussion
4. !. Theoretical analysis
In order to understand the functional role of lead as a free
cutting additive, it is essential to reach a quantitative under-
standing of the following effects noted in free cutting steels
due to lead addition:
1. the reduction of cutting forces during the machining of
leaded steel;
2. reduction of contact lengths during the machining of
leaded steel;
3. marginal reduction of the diffusional crater wear during
machining with carbide tools at raoderately high cutting
speeds.
4.1.1. Contact lengths
Various models for the contact length have been proposed,
for example, by Abuladze [6], Lee and Shaffer [7], Thom-
sen and Kobayashi [8]. All these models are applicable at
low cutting speeds, where the tribology of sliding operates at
the tool-chip interface. But, at higher cutting speeds, where
the tribology of seizure occurs, these models are not appli-
cable. The chip material of the seized region in the secondary
shear zone is sheared under the action of compressive and
shear forces acting on the rake face of the tool. Owing to the
large normal forces, the apparent contact area approaches the
real contact area. In their analysis of junction growth during
metallic friction Bowden and Tabor [9] used an equation
first developed by Nadai [ 10] to relate the normal and shear
stresses to the material flow stress in a seizure situation. This
equation:
Cuttingforce,contactlengthandshearplaneanglemeasurementsonfreecuttingsteelsalongwiththepredictedcontactlengthandmaximumtool.chip interface
temperature
Grade Speed Tool Fc(N) Xc(measured) Xc(predicted) ~ (tad) Fr/F~ Maximum
(m rain- ~) (ram) (ram) temperature(K)
1215 '.40 K-I 547 0.84
1215 160 K-I 538 0.85
1215 180 K-I 530 0.84
1215 180 K-I 530 0.84
1215 200 K-I 518 0.94
12LI4 140 K-I 397 0.71
12LI4 160 K-I 393 0.74
12LI4 180 K-I 367 0.74
12LI4 200 K-I 357 0.79
1215 160 K-I 440 0.612
(I.E.)
1215 200 K-I 438 0.61
(I.E.)
1215 1(30 KI + Hfi'q 477 0.49
1215 200 KI + HfN 477 0.50
0.77 0.41 0.62 1136
0.78 0.43 0.60 1155
0.76 0.44 0.57 !196
0.76 0A4 0.57 1196
0.85 0.45 0.55 1216
0.64 0.51 0.58 1036
0.51 0.57 10670.67
0.67 0.51 0.57 1101
0.71 0.51 0.55 1138
0.55 0.48 0.50 1134
0.55 0.53 0.52 1167
0.45 0.45 0.53 1015
0.45 0.50 0.56 1070

5. K. Ramanujachar,S.V.Subramanian/ Wear197(1996)45-55 49
2
r-I o '~=~" .~
o s~,, ,~ :~.~._.~
-6 0
O.OO0! 0.000~' 0.0003 0.0004 O.O00S0.0006 0.0007
X--CO~N
Fig.4. TypicalelementalconfigurationusedintheFEM.Theinsetshows
thedomainwithrespecttothesecondaryshearzone.
1400.0
1800.0-
1200.0-
1100.0
g ~ooo.o®
g00.0
600.0
~" 7oo.o
600.0
600.0
,~O0.O 0:1 0:2 0:~ 0:4 O:S O:S O;7 O:S 0:0
l~r~.bnd did.z~e.
Fig.5.Thetemperaturedistributionalongthetool- chipinterfacecalculated
usingFEManalysisduringmachininganAI$11215withuncoatedK-Itool
ata cuttingspeedof 160mmin-=.
p2 + 3S2= p2o (1)
Where P = normal stress, S = shear stress, Po = material flow
stress in compression, is used to describe the plastic flow in
the secondary shear zone and is the starting point of the
analysis carried out in Appendix A. It should be noted that
while the forces are derived with a flow rule, the contact
length is derived by calculating the total work, which is the
sum of the work done in the primary shear zone and the work
done in the secondary shear zone. The experimentallydeter-
mined shear plane angle was however used for the tempera-
ture calculations. Thus both the mechanics of tool-chip
contact as well as the energetics of metal cutting have been
taken into consideration in this analysis. At the high cutting
speedsthat have been used in this study the singleshearplane
theory of Merchant is considered to be a good approximation
and strainrates on this plane are used in calculatingthe plastic
work in the primary shear plane [ 11,12]. It should be noted
that shear occurs over a narrow zone but in the theoretical
analysis, the thickness of this narrow zone does not occur
explicitly in the expression. The resulting formula for the
seized contact length is given by:
X
7~Ip cos a tlsee(4'- a) cosec 4'= : o (2)
= l/g(a,k)- (l12Vr3) sin ~cos(4'-a)
where
~-p== shear flow stress of the chip material in the primary
shear zone, a= rake angle, 4'ffishear plane angle, tt ffifeed,
Po fficompressiveflowstressofchipmaterial in thesecondary
shear zone, k= feed force/cutting force and g(a,k) is given
by
g(cz,k) = 1/2 sin2a+ 1+2k2cos2a + k2+ 2k sin(2a)
(3a)
Since the work of shear in the primary shear zone consti-
tutes in excess of 80% of the total work done even under
conditions of seizure, it is instructive to evaluate the contact
length by approximatingthe work done as that ofthe primary
shear zone. The approximation leads to the following sim-
plified expression, given by
ep~cos a ttg(a,k)
X~=po cos(4'-a) sin 4'
(3b)
Table 5 shows a good agreement between the predicted
and measured values of contact lengths. This approach in
spite of the simplifying assumptions involved serves as the
basis for identifyingthe key parametersthat influencecontact
length. According to Eq. (3), the contact lengthXc:
1. increases with the feed tt;
2. decreases with the ratio feed force/cutting force (k); and
3. decreases with the increase in the shear plane angle.
•The k valueis lower in thecaseof leaded steel.Additionally
the valueoftheshearplane angle, 4'is largerwhile machining
leaded steel. Consequently the reduction in contact length
that occurs while machining leaded steel is linked to the
reduction in the value ofk and the increasein the value oftbe
shear plane angle, 4', Physically k can be considered analo-
gous to a coefficient of friction in a simple sliding situation.
The cutting forces in thecase of leaded steel are reduced with
respect to the unleaded steel as a result of the reduction in the
contact length, X=.
The geometry of the secondary shear zone is roughly tri-
angular in shape with its base as the seizedcontact lengthand
its height as the secondary shear zone thickness.The dimen-
sions of this triangle are established from optical metallog-
raphy. For this triangular geometry of secondary shear zone,
a first principles thermal finite element analysis is carried out
to establish tool-chip interface temperature.
4.1,2. Thermalfmiteelementanalysisofthesecondaryshear
~.one
In order to calculate theoretically the temperature at the
tool-chip interface we have to solve the partial differential
equation:
K ~Z(T) +Q/pCp=O (4)

6. 50 K.Ramanujachar,S.V.Subramanian/ Wear197(1996)45-55
where K= thermal diffusivity, T= temperature, Q = rate of
heat generation,p = density of thechip material, Cp= specific
heat.
The heat generation term in the above equation arises from
the heatgeneratedas a resultofplastic shearingof thematerial
in the secondary shearzone immediatelyadjacentto the tool-
chip interface. A schematic diagram of the secondary shear
zone is shown in Fig. 4. The temperature was assumed to
vary linearly over the upper boundary of the region (OC)
from the primary shear plane temperature at O to a maximum
calculated from the Boothroyd model [5] at the end of the
seized tool-chip contact lengthC. The flux at tool-chip inter-
face BC was calculated from the steady state condition. The
flux boundary condition was converted to an integral over the
domain area with the aid of the Gauss-Green theorem. The
integral was then evaluated by discretizing the domain into
linear variational triangular elements whose typical configu-
ration is shown in Fig. 4. This process converts the partial
differential equation into a system of linear equations [ 13].
Th,~~ystemof ~quations was then solved with a Gauss direct
elimination which duly took advantage of the sparse nature
of the global stiffnessmatrix. For the temperature profile, the
local equilibrium concentrations of tungsten were calculated
from solubility data for WC in appropriate chip phases. The
diffusional mass transport of tungsten into the chip was then
calculated.
The program input consists of:
Experimental input:
1. Contact length
2. Cutting force
3. Feed force
4. Chip thickness
5. Shear plane angle
6. Depth of cut
7. Cutting speed
Material parame:cr~:
1. Density of workpie:e material
2. Specific heat of workpiece material
3. Thermal conductiv.ty of workpiece material
4. Shear yield strength of chip material in the secondary
shear zone
5. Chemical diffusivity of solute in the chip
Output:
1. Temperature distribution at the tool--chip interface
2. Tooi material transported into the chip by diffusion
A typical temperature profile generated is shown in Fig. 5.
The temperature increasescontinuously over the seized con-
tact length in qualitative agreement with Tay et al. [ 14].
Appendix B shows a schematic diagram of the algorithm.
This model is used to predict tool-chip interface temperatures
during the machining of leaded and unleaded steel. Table 5
compiles the maximum temperatures during the machining
of leaded and unleaded steel at the four different cutting
speeds. Experimentally determined forces and geometrical
parameters, the contact length, shear plane angle were used
to compute these temperatures. Clearly the temperatures in
leaded steel are lower than those of unleaded steel. This is in
agreement with the limited thermocouple measurements
reported by Shaw et al. [ 1]. Having established that temper-
atures are lower in the case of leaded steel we can use the
temperature profiles generated through the model to obtain
the amountof tungstentransported by dissolution wear to the
chip material by diffusion.
4.1.3. Quantitative modelling of diffusion wear
FEM analysis has shown that the temperature along the
seized contact length does not vary linearly. At the tool tip
the temperature is that of the primary shear zone. At the end
of the seized contact length, the temperature gets raised to a
maximum temperature predicted by Boothroyd's model.
Thus a volume element of the chip traversing through the
seized contact length starts offat the tool tip as ferrite at the
low primary shear zone temperature and progresses through
a temperature range, transforms to austenite that gets raised
to a maximumtemperature at the end of the secondary shear
zone. The secondary shear zone undergoes severe deforma-
tion involvinga large shear strain, high strain rate under high
compressivestress. In the ferrite region, the effect of severe
deformation is accommodated by shear bands. In conse-
quence, the diffusivity is probably enhanced by dislocation
pipe diffusion. Evidence for anstenisation has been reported
by Sheibourne [ 15], Hau Bracamonte [ 16] and Ingle [4].
TEM examination of the secondary shear zone by Ingle has
revealed the presence of ultra fine grains of less than 0.5/~m
diameter in size. During deformation of austenite in the sec-
ondary shearzone, dynamicrecrystallisationprobably occurs
involving grain boundary motion. Grain boundaries act as
high diffusivity paths [ 17]. The moving boundaries during
recrystallisation contribute to enhanced diffusivity [18].
Clearly, the lattice diffusion is enhanced by the high diffusiv-
ity paths available in the fcrrite and the austenite regions of
the secondary shear zones as well as the region undergoing
ferrite to austenite phase transformation.
In the model originally proposed by Bhattacharyya [ 19],
the mass transported is given by the equation
M = 1.1284CoA(~-~) (5)
whereM = numberof moles transported, Co= interfacialcon-
centration in moles per volume, A = cross-sectional area,
D = diffusion coefficient, and ¢= diffusion time.
The interracial concentration Co and the diffusivity D are
critical material parameters that determine diffusion wear.
However in Bhattacharya'smodel, the equilibrium solubility
product of tungstencarbide that determineslocal equilibrium
of the tungsten concentration at the interface was not taken
into account. Ingle [4] has developed a thermokinetic model
that takes into consideration the local equilibrium concentra-
tion of tungsten that is determined by the equilibrium solu-
bility of tungsten carbide in austenite. However, Ingle's
model assumesan average interface temperature and austen-

7. K. Ramanujachar.S.V.Subramanian/ Wear197(1996)45-55
Table6
Calculateddiffusionenhancementfactorto reconcilewiththeexpetimen-
tallymeasureddissolution
Speed Tool Diffusionenhancementfactor DEFforGrade1215
(DEF)forGrade12L14
160 K- I 9654• 336•
200 K-1 536• 160•
•Thisistheaveragevaluebywhichthelatticediffusioncoefficientshould
beenhancedinordertoreconcilewiththeexperimentallydetermineddis-
solutionwear.
Database:
logD~=A~+BIIT, AI ffi -3.5, B~ = - 12850 [20]
logDr=A2+ B2/T. A2= --2.5,B2ffi- 16300[20]
logWaC~=A~ +BJT, AI =3.3,Bj=6404.3 [21l
logW~Cy=A2+B21T,A2=5.3,B2=6404.3 [20]
isation over the entire seized contact length. In the present
model, the temperature variation along the seized contact
length has been considered and the diffusional model is cou-
pled to the phase transformation resulting from temperature
variation along the contact length. The equilibrium solubility
product of WC and the lattice diffusion coefficient of W
appropriate to anstenite and the fcrrite regionsare considered.
In the absence of quantitative information on the diffusivity
enhancement in the ferrite and austenite regions, the lattice
diffusion coefficient is used to predict the dissolution wear.
An average enhancementfactor for lattice diffusion has been
calculated for each case to reconcile the predicted value with
the experimentally determined value of tungsten dissolution.
Table 6 summarizesthe back-calculatedenhaneement factors
for the leaded and non-leeded steels for the cutting speeds
160 m rain- i and 200 m min- ~respectively. As the cutting
speed increases, the tool-chip interface temperature increases
and the enhancement factor decreases. The diffusion
enhancemen: factor is greater for leaded steel than that for
unleaded steel, suggestingthat there are morehigh diffusivity
paths operating in the leaded steel than that in the non-leaded
steel.
4.1.4. Micromechanismsoftoolwearinmachiningfree cut-
ring steels with carbide tools
Both leaded free cutting steel (AISI 12L14) and
non-leaded free cutting grade steel (AISI 1215) exhibit
pronounced crater wear at cutting speeds 160 and
200 m min- i respectively, when machined with K 1 grade
cemented carbide tool. Bulk of the tungsten lost from the tool
is dissolved atomically in the chips, confirming that dissolu-
tion wear is the dominantmechanismof tool wear, see Figs. 2
and 3. The effect of addition of Pb to free cutting grade of
steels is to decrease the cutting force (27-41%) and the
contact length ( !3-16%) as can be seen fromTable 5 and as
a result, decrease the predicted maximum tool- chip inter-
face temperature by 80 °C. However, the tool dissolution
wear as measured by dissolved tungsten in the chips is com-
parable to that of non - leaded steel. Clearly, neither lead nor
sulphide inclusions are effective in preventing or minimizing
tool dissolution wear at highercutting speeds during machin-
ing with cemented carbide tools.
4.1.5. Effect of engineered deformable oxide inclusions on
tool dissolution wear
The effect of engineering deformable oxide inclusions of
anorthitic composition CaO. AleO~.2S;.Oeis to suppresstool
dissolution wear. Extensive research hasbeencarded outover
the years on inclusion engineering with respect to thermo-
dynamic phase stability of inclusions relevant to deoxidation
of steel, deformability of inclusions during steel hot rolling
[29-31 ]. Fanlring and Ramalingam [32], have investigated
the beneficial effect ofcalcium modification ofaluminainclu-
sions on tool wear. Gatellier et al. [33] have mapped the
viscosity of inclusions on the CaO--AI203-SiO2 phase stabil-
ity diagram and have demarcated the glassy anorthite phase
region as being deformable because of low viscosity, lower
than the plasticity of steel. Pierson et al. [34]. Unimetal
Research have developed a new process for the production
of glassy phase anorthite inclusions in free cutting steels that
improve the tool life while cutting with carbide tools. The
exact mechanism by which glassy phase anorthite inclusions
improve the tool life has not been resolved and this is the
subject of further discussion.
It is instructive to examine the diffnsivity of tungsten in
the glassyanorthite phase. The equation governing the trans-
port ofcationic species (W) in an anion network is described
by the Nemst-Planck equation:
whereJ~= ionic flux,D~= selfdiffusivity, Ci= coneenwation,
Z~fcharge, FfFaraday constant, ~bfelectrical potential,
T= temperature, R = gas constant, and X= distance.
An additional condition is that charge neutrality has to be
maintained which results in
~z~',=0 (7)
The resulting flux equation depends not only on the con-
centration gradient of the cation concerned (W), but also on
the concentration gradient ofconcomitantly diffusing anions:
0c,
Yi=DI I-~"x"1"DI20x (g)
where
Di~-"DsCs(Zs)2- DIC,~ Z,) 2
DH= /" ~-"DjCs(Zs)2 (9)
jffil
~D,DTNC,
DI2•Jn ! -- (10)
E ~)jCj(Zj) 2
j=l

8. K.Ramanujachar.S.V.Subramanian/ Wear197(1996)45-55
proposed for the case of diffusional wear. By choosing a
coating that is stable enough to resist dissolution or decom-
position, wear due to chemical instability can be suppressed
[36].
The effect of machining non-leaded free cutting steel AISI
1215 with a HfN coated K- 1 tool is to increase the shear plane
angle and decrease the feed force to cutting force ratio, see
Table 5. In consequence, the contact length decreases in
accordance with the foregoing theoretical analysis. Temper-
ature distribution along the seized contact length was calcu-
lated using thermal finite element modelling. Temperature
distribution at the tool-chip interface was lowered as a result
of coating. Quantitative analysis of diffusional wear for HfN
was carried out using the equilibrium solubility product of
HfN as a function of temperature for the appropriate stable
phase involved i.e. ferrite or austenite. Typical results
obtained for machining with HfN coated tool at a cutting
speed of 200 m min- t are summarized in Table 7. Clearly
the effect of low equilibrium solubility product of HfN is to
suppress the diffusional flux to rather low values even at the
maximum tool-chip interface temperature. The diffusional
wear is dramatically reduced by HfN coating.
Quantitative analysis of diffnsional wear has shown the
effect of coating is to decrease the maximum temperature at
the tool-chip interface by about 140 °C. The effect of reduc-
ing the maximum tool-chip interface temperature by 140 °C
on the diffusional transport of tungsten in uncoated AISI 1215
will be to decrease the tungsten concentration in the chip
from 0.17 ppm to 0.11 ppm. On the other hand the local
interface concentration of Hf is decreased by seven orders of
magnitude because the equilibrium solubilityproduct of HfN
is seven orders of magnitude less than that of WC in the
anstenite phase. In consequence, the dissolution wear of Hf
is dramatically decreased to 10-6 ppm of HfN.
In the diffusional wear analysis, the local equilibrium con-
centration of solute and the square root of the diffusion coef-
ficient of the solute (C~ °5) are the key material parameters
52
Two diffusion coefficients enter the flux equation. Each of
these diffusion coefficients depends on the intrinsic diffusiv-
ity of the cation concerned in the glass phase as well as the
concentration of anion in the glass phase [22]. The intrinsic
diffusivities of cations like W in a glass phase like calcium
aluminosilicateare dependent on the composition of the glass,
but in general are high due to the open nature of the frame-
work structure of anorthite [23]. Anorthite exhibits a struc-
ture much like cristobalite with characteristic large
interstices. The diffusivity of Mn ion in the CaO--AlzO3-SiO2
slag phase at 1763K has been reported to be
2× 10-4 cm2 s- ~ [24]. The diffusivity of Fe ion in a slag
phase of similar composition has been reported to be
5X 10-tcmZ s -I [25].
Since the slag exhibits an open structure that favours high
diffusivity of tungsten in the anorthite glassy phase it is sur-
mised that the solubility of either WC or tungsten oxide in
me glass phase is probably too low to support significant
diffusion of tungsten through the glassy anorthite phase. It
must however be pointed out that unlike the ordinary disso-
lution of tungsten in austenite discussed with AISI 1215,
dissolution of a metallic species in a glass phase is governed
by electrochemical reactions of the kind
M,=(M~)~+Z~e (ll)
It is possible that the glassy inclusions could change the
tribology of seizure at the tool-chip interface thereby sup-
pressing dissolution wear. Further research is required to
resolve the mechanism underlying the suppression of crater
wear brought about by glassy inclusion.
4.2. The effect of HfN coating on tool dissolution wear
Kramer and Sub [35 ] have investigatedsolution wear due
to the dissolutionof the tool material in the chip at high cutting
speeds. The concept ofloeal equilibrium is applied at the tool
chip interface and the transport by diffusion mechanism is
Table7
TypicalresultsfromdiffusionalwearanalysisobtainedformachiningAIS11215withHfNcoatedK-I toolat 200m/min
Temperature(K) Solpro Itcon(%) Diffco Masstran Con/,ppm)
552 1.06X10-Ja 3.26X 10-07 1.26× 10-31 6.2× 10-~9
582 1.61X10-12 1.27X10-°~ 8.43X10-29 6.23X10-~
685 3.3X l0-H 5.75X l0-°~ 1.16X l0-2s 1.05X10-~
771 2.79X 10-Jo 1.67X 10-°~ 1.93X 10-23 3.93X l0-23
856 1.54Xl0-c° 3.93x 10-°s 1.16x l0-zl 7.15x 10-22
940 6.25x l0-~ 7.91x 1O-°5 3.28× 10-2o 7.65× l0-21
1023 2X 10-°a 0.000141 5.31x 10-19 5.51x 10-20
1105 5.34Xl0-~ 8.3X 10- 05 5.58X10- ,s 1.05× 10- j'~
1186 1.24Xl0-ca 8.32X l0-°5 4.19X l0-17 2.88X 10-19
1265 2.57X10-°~ 8.36× 10-°5 2.4× 10-16 6.91× l0-19
1.05 × 10 -°6
Solpro= solubilityproductin wt.%~.
Rcon~interface concentrationofHfNin wt.%.
Diffco= diffusioncoefficientof Hfin rn2s - '.
M~suan= amountof Hfin gramstransportedto a volumeelementofchipduringitstraverseoverthecontactlengthsegment.
Con= totalconcentrati~ ~of Hfin thecl',ipsduring I5 s ofmachining.

9. K. Ramanujachar.S.V. Subramanian/ Wear 197 (1996)4.5-55
Table8
Equilibriumsolubilityproductsofnitridesandcarbidesat1400K(1123°C)
inaustenite
[Ti]IN] = 1.34X10-6 [Til[C] =7.04× 10-6
[Zrl[N] =2.34:<10-~ [Zr][C] = 1.64X10-2
[Hf]IN] = 3.15× 10-~ [W][C] =4.08
[Nb][N]= 3.54× 10-~ [Nb][C] =5.99× 10-3
[Ta][N] = 1.26×10-4 [Ta][C] =7.94× 10-3
[V][N] =2.60X 10-3 IV][C] =8.60× 10-I
that determine the diffusional wear. In the case of the HfN
coating on WC tool, the decrease in the local equilibrium
concentration (C~) has a dominating effect over the temper-
ature effect in decreasing the diffusional wear. The local
equilibrium concentration C~ in torn is determined by the
equilibrium solubility product of carbide or nitride as thecase
may be. Table 8 summarizesthe equilibrium solubility prod-
uct of nitrides and carbides in the austenitic phase of steel at
1400 K [26].
A good adhesion between the tool and the coating, gener-
ation of thermal stressesat the tool - coating interface due to
differential thermal expansion, the possibility of formation
of brittle phases at the tool-coating interface are additional
considerations in the choice of a high integrity coating to
minimize chemical dissolution wear. For example, even
though TiN has a lower solubility product that is six orders
of magnitude less than that of WC, delamination of TiN
occurred in the presentwork rendering the coating ineffective
in suppressingdissolution crater wear.
5. Strategies for replacement of lead in free cutting
steels
The functional role of lead at low cutting speeds has been
the subject of intensive investigation by Shaw et al., who
concluded that lead forms a fluid layer at the tool-chip
interface that aids the tribology of sliding wear operating at
low cutting speeds [ ! ]. They have reported that the contact
length and cutting force are reduced by lead additions. The
role of lead inclusionsin promotingductile fracture involved
in chip formation has been analysed in terms of damage
mechanismsand the effect of lead in lowering critical accu-
mulateddamage is quantifiedby CAD measurementsby Sub-
ramanian, Kay, Stinson and Finn [28].
At higher cutting speeds using cemented carbide tools, the
tribology of seizure sets in at the tool - chip interface, result-
ing in dissolution crater wear by diffusion mechanism. The
present work has established that neither lead nor sulphide
inclusion is effective in suppressingthe dissolution wear by
diffusion mechanism. However, deformable anorthitic oxide
(CaO. AI203" 2SIO2) inclusionsengineeredin the workpiece
or coating the tool with a HfN compound are found to be an
effectivemeans of suppressingdissolutioncrater wear. Thus,
the strategy to replace lead has to consider the tribological
conditions operating at the tool - chip interface, At low cut-
ring speeds, where the tribology of sliding wear operates at
the tool- chip interface, it is essential to design soft inclu-
sions such as sulphidethat can lower the critical accumulated
damage to that of leaded steel. However, at higher cutting
speeds where the tribology of seizure operates at the
tool- chip interface, leaded steel can be outperformed by
engineeringdiffusionbarrier at the tool - chip interface. The
two viable option~ are:
1. Engineer deformable glassy phase oxide inclusions such
as CaO.Ai203-2SiO2 into the workpiece that suppress
tool dissolution wear.
2. Coat the tool with a compound such as HfN that has the
least solubility in the workpiece at the typical tool-chip
interface temperature.
Conclusions
Dissolution wear is the dominant mechanism of crater
wear in both leaded AISI 12L14 and non-leaded AISI
1215 steels during machining at high cutting speeds with
cemented carbide cutting tools.
2. Neither lead nor manganesesulphideinclusionsare effec-
five in suppressing diffusional crater wear. However,
engineered deformable oxide inclusions of the type
CaO- AI203 •2SIO2 into the workpiece are very effective
in suppressingdiffusional crater wear.
3. Coating of thecementedcarbide tool with HfN iseffective
in suppressingdissolution wear of cemented carbide tool
during machining of non-leaded free machining steel
AISI 1215 at higher cutting speeds.
4. The addition of lead results in an increase in the shear
plane angle and a decrease in the feed foree/cutting force
ratio which results in a decrease in the tool chip contact
length. The predicted tool--chip interface temperature
decreases by about 70 °C but this only has a marginal
effect on the tool dissolution wear.
5. Coating the tool with HfN increasesthc shearplane angle,
reduces the feed force/cutting force ratio and hence
decreases the tool chip contact length compared to the
uncoated tool.
6. An analytical expression has been developed relating the
seized tool chip contact length to the mechanics of metal
cutting and material properties. Bowden and Tabor's
equation is used to describe the plastic flow of the seized
material under the action of compressive and shear
stressesand the energetics of metal cutting is also taken
into consideration.Thisexpressionpredicts thattheseized
contact length is a function of the shear plane angle, the
feed foree/cutting force ratio,, feed, and the flow stresses
of the workpiece material in the primary and secondary
shear zones, This expression has been shown to be inter-
nally consistentwith the experimentaldata.
7. A first principles thermal finite element model has been
developed to predict the tool-chip interface temperature.
Quantitative modelling is user2to predict the transport of

10. 54 £. Ramanujachar.S.V.Subramanian/ Wear197(1.o96)45-55
tungsten from the tool into the chip by diffusion mecha-
nism. An enhanced diffusioncoefficientthat is two orders
of magnitude greater than the lattice diffusion coefficient
is required to reconcile the predicted and measuredvalues
of tungsten dissolution in the chips.
8. The local equilibrium concentration of solute C~and the
square root of the diffusion coefficient D°'s are identified
as the key material parameters that determine the diffu-
sional tool wear. By decreasing the temperature of the
tool--chip interface, both the equilibrium concentration
and diffusion coefficient of the solute are decreased and
in consequence the diffusional wear is reduced. But, the
local equilibrium concentration of the solute can be
decreased by orders of magnitude by choosing a coating
or a too! material that has the least dissolutionpotential in
the workpiece, and consequently,the dissolutionwear can
be decreased by orders of magnitude.Thus thedissolution
wear is more effectively suppressed by a coating that
lowers the local equilibrium concentration of solute at the
tool---chipinterface.
9. Inclusion engineering with deformable oxide inclusions
such as anorthite or coating the cemented carbide tool
with a coating that has the lowest equilibrium solubility
product in the workpiece material are two attractive and
viable alternatives to replace lead in free cutting steels.
Acknowledgements
Financial support of the research on free cutting steels by
the NSERC in the form of a Strategic Grant award and Ford
Motor Company USA is gratefully acknowledged. Helpful
discussionwi~h Drs. J.D. Embury and M. Elbestawi is grate-
fully acknowledged.
Special thanks are expressed to Mr. W.E. Heitmann and
Roger Joseph of theInland Steel ResearchUSA for the supply
of steelsused in this study, Ms Alice Pidruczny for extensive
help with Instrumental Neutron Activation Analysis and Mr
Pran Khindri, President, Weliworth Manufacturing,Oakville
for help with high speed machining.
Appendix A. An expression for the seized contact
length
The following derivation for the seized contact length
attempts to capture the mechanics of tool-chip contact as
well as the energetics of metal cutting. The mechanics of
tool-chip contact which is a situation of seizure is described
by the equation due to Bowden and Tabor:
p2 + 3Sz=p2 (Al)
where P = normal stress on the rake face, S= shear stress on
the rake face, Po = chip material flow stress.
For cutting with a positive rake angle ot
p = Fc cos a-F~/sin a (A2)
X~Z
where Xc=seized contact length, Z=depth of cut,
Fc = cutting force, F~ = modified feed force.
S= Fc sin ot+F~ cos a (A3)
x~z
Substituting Eqs. (A2) and (A3) iutu Eq. (At) and set-
ting F~/Fc =k we obtain the following equation for the cut-
ting force
F¢= PoXcZ/g( a, k) (A4)
where g(ot, k) is defined as
g (or, k) = ~/2 sinZot+ 1+ 2k2cos2a + ~ + 2k sin(200
(A5)
The work expended in machining is given by
W=FcVc (A6)
where V~= cutting velocity.
This work is expended in shearing the material in the two
zones of shear: the primary shear zone and the secondary
shear zone.
The work in the primary shear zone is
epzVslz (A7)
where ¢pz= shear strength of the chip material in the primary
shear zone, Vs=velocity on the primary shear plane,
l= Feed/sin ok,the length of the primary shear region.
The work done in the secondary shear zone is given by:
"rssVcm~12 (A8)
where ~-~= flow stress of material in the secondary shear
zone, Vchip = velocity of the chip, xc= contact length, and
z = depth of cut.
By using Eq. A4F_.q.A6Eq. A7Eq. A8 we obtain the fol-
lowing expression for the seized contact length.
X¢= ('rpz/po)COSa sec(q~-a)tl cosec ~b (A9)
I/g( et, k) - ( 1/2Vf3)sin ~ksoc0k- a)
Since the work done in the primary shear zone is over 80%
of the work done in machining in the first approximation the
two can be equated yielding the following simplifiedexpres-
sion for the seized contact length.
X~= ¢p~eos¢~tlg(cc, k)/PoCOS(~b-c¢) sin ~b (A10)
where tl = feed, In a typical case the plastic work from the
secondary shear zone is about 15%.
If this is included in the analysisby multiplyingEq. A6 by
an appropriate factor the predicted value ofX~yields an even
better agreement with the experimental value than approxi-
mating the total work as that of the primary shear zone.

11. K. Ramanujachar, S.V.Subramanian / Wear197 (1996) 45-55
Appendix B, Algorithm for the FEM calculation of
tool-chip interface temperature along the seized contact
length [27]
INPUTMATERIALFLOW,~ PROPERTIBS~[
] COMPIL/~NODEGORUINATI~JELEMENTAREAS I
t
DISPLAYANDGRAPHRESULTS ]
References
[I] M.C. Shaw, P.A. Smith,E,G. Loowanand N.H. Cook, Trans ASME,
78 (1956) 1143-1165.
[2] E.D. Doyle and D. Tabor, Prac. R. Soc. London, Ser. A, 173 (1973)
336.
[3] E.M.Trent,Metal Cutting,3rfledn,, Butlerwoflh-Heinmann,London,
1984.
[4] S.S. Ingle, The micl~omechanismsof cemented carbide cutting tool
wear, Ph.D. Thesis, McMasterUniversity,1993.
[5l G. Boothroyd,Proc. Inst. Mech. Eng., 177 (29) (1963) 789.
[6] N.G.Abuledz¢,On the Plastic ComactZone. TranslatedPublications,
GeorgiaPolytechnicInstitute,Tbilisi. USSR.
[7] E.G. Thomsen and S. Kobayashi,Metal curingptinciples~I, Trans,
ASME, Set. B. 84 (1962) 71.
[8] E.H. Lee,B.W. Shatter, Trans. ASME, 18 (4) ( 1951) 405.
[9] D. Tabor,Proc. R. $oc. London, Ser. A, 251 (1959) 378.
[10l A. Nadai,Plasticity, McGraw-Hill,NewYork, 193I.
[II] K.Okushima,Hitomi,Trans. Jpn. Soc. Mech. Eng., 26 (1960) 243.
[12] D. Keeecioglu,Trans. ASME, 80 (1958) 158.
[13] D.RJ. Owen,A Simple Guide to Finite Elements, Pineridge,Swansea,
UK, 1979.
[14] A,O. "lay, M.G. Stevenson,G. de V. Davis and P.L.B.Oxley,Int. J.
Mach, Tool. Des. Res., 16 (1976) 335.
[15l A.M. Sbelboura,W.T. RobertsandE.M.T~nt, Mater. ScL Technol.,
I (1985) 220.
[16] J.L. Han-Bracamonte,Metals Technol., 8 ( 1981) 447.
[17] I. Kaur, W. Gust and L. Kozma, Handbook of Grain and Interface
Boundary Diffusion Data. VoL2, Zieglet,Stuttgart. 1989.
[18] D.R.Olanderand U.M.EI-Saied,Acto Metall., 40 (6) (1992) 1329-
1336.
[19] A. BhattacharyyaandA. Ghosh, Proc. Sth Int. Machine Tool Design
and Research Conf., 1964,Pcrgamoo,New York.
[20] B.Urehenius,in D.V.DoaooeandJ.S. Kirkaldy(~Ls.),Hardenability
Concepts with Applications to Steels, 1977,
[21] T. Nishizawa, Proc, Int. Conf. on The Physical Metallurgy of
Thermomechanical Processing ofSteels andOther Metals, Vol. l, iron
andSteelSocietyofJapan,Tokyo, 1988,p. 39.
[22] D.A.OkangwuandW.K.Lu,J. Chem. Phys., 63 (1975) 734.
[23] W.D. Kingery, H.K. Bowen and D.R. Uhlmannn, Introduction to
Ceramics,Wiley,New York, 1975,
[24] D.D.Ma,Ph.D. Thesis, McMaaterUniversity,1992.
[25] L. tang and G. Derge, Paper in the Physical Chcmist~ of Process
Metallurgy,AIME, 7 (1961) 503.
[26] K. Balasubramanian,unpublished research, McMaster University,
1985.
[27] K. Ramanuja,'har,Micrnrnechanismsof tool wearin machiningfree
cuttingsteels,M. Eng. Thesis, McMaaterUniversity, 1994.
[281 S.V. Subramanian,D.A.R. Kay, R. Stinson and M.E. Wren.in M.E.
Finn (ed.), Proc. Int. Conf. on Factors Influencing Machining and
their Controls,ASMoCincinnati, 12-14 Sept. 1989.
[29] G. Benard, P.V. Riboud and (3. Urbain, Etude de la Piasticit6
d'lnclusions d'Oxidcs, IRSID Rep. PCM-RE. 776, November 1980.
[30] G.Banard,P.V.RiboodandG. Urb~n, Rev.Met. CIT, 78(5) (1981)
421--434.
[31] G.M. Faulting and S. Ramalingarn,Inclusionprecipitationdiagrams
forthe Fc-O--Ca-AIsystems,Metoll. Trans. B, lib (1980) 125-130.
[32] G.M.Faultingand S. Ramalingam,Oxide inclusionsandtool wearin
machining,Metall. Trans.A, IOA (1979) 1781-1788.
[33] C. GatellieroH. Gaye, J. Lehman,J.N. Pontoireand F. Castro, Proc.
74th Steel Making Conf., Washington, DC, 1991, ISS, Warrcndale,
PA,pp. 827-834.
[34] G. Pierson,L. Payraudeanand J. Bellot, Soft steel for metal cutting
andmethodofproducingit, USPatent 4 978 499,1990.
[35l B.M. Kramar and N.P. Sub, Tool wear by solmion: a quantitative
understandingJournalofengincctingforindusu3,,Trans.ASME, 102
(1980) 303-309.
[36l N.P. Sab, Tribophysics, PrenticeHall,NJ, 1986,pp. 366-382.
Bibfiographies
K. Ramanujach~r: obtained his B.Tech. from the Indian Insti-
tute of Technology (Kanpur) and M. Engg from McMaster
University. He is currently pursuing graduate studies at
McMaster University for his Ph.D. degree.
S.V, Subramanian: graduated from Benarcs University
(India) and obtained his M. Met and Fh.D, degrees from the
University of Sheffield, UK. He is currently in the faculty of
Materials Science and Engineering at McMaster University.
His current research interests include basic science and tech-
nological aspects of solidification processing, thermome-
chanical processing of steels and microstructural engineering
for improved machinability.