Wear, 104 (1985) 157 - 171 157

PROCESSES I~OLVED IN THE WEAR OF CEMENTED CARBIDE TOOLS

GWENDOLYN DIXON and ROGER N. WRIGHT

materials engineering Department, Rensselaer Folyte~hnic Institute, !Froy, NY 12180 (U.S.A.)

MINYOUNG LEE

General Electric Corporate Research and Deuelopment, Schenectady, NY 12301 (U.S.A.)

(Received June 30,1983; revised March 13,1985; accepted June 24,1985)

Summary

Studies were made of the initial cratering wear and subsequent wear propagation of WC-(Ti, Ta, W)C-Co tools used to cut carbon steel. Using rnech~i~~ly polished tools and short cutting times (1 - 11 s), microstruc- tural and chemical changes involved in the initial wear process were identi- fied and used as a basis for evaluating several wear models. Making use of iron-coated and diffusion-annealed tools, the wear propagation process was similarly evaluated in tests of up to 6 min duration.

Microscopic examination shows that WC grains fracture intragranularly, con~quen~ly generating fine WC wear debris. Analysis of chip material reveals the size of WC wear debris to be an order of magnitude less than the original grain size. The (Ti, Ta, W)C wear debris size was found to be of the order of the original grain size. Adhesion seems to be the most, likely mech- anism of wear. Abrasive wear seems to be inconsistent with the observed similarity between results for ~clusion-conning 52100 steel and 1045 steel relatively free of inclusions. Despite evidence of subs~tial diffusion, diffusional wear models seem inapplicable owing to the relative wear rates of coated and uncoated tools and to the absence of intergranular failure. Dis- solution models seem inconsistent with the observed coexistence of the iron coating and the WC particles.

1. Introduction

The wear of commercial cemented carbide tools during turning oper- ations has been the subject of a number of major investigations, and abra- sion, adhesion, diffusion and d~solu~ion have all been cited as playing a role in the wear process. No single mechanism seems satisfactory. Abrasive wear theory cannot account for the association of short tool life with the tenden- cy for some tools to react chemically with the workpiece [ 1 - 31. Also, abrasive wear theory tends to predict a larger number of grains of wear

0043-1648/85/$3.3~ @ Eisevier ~~uoia/~inted in The Netherlands

158

debris on the underside of the chip than are found [4]. Adhesive wear theory is not consistent with observations in quick-stop experiments. Quick- stop experiments have shown that the tool usually separates from the chip rather cleanly, thus discrediting the theory of a strong bond between the tool and the chip [4]. Diffusion wear theory still lacks solid data to support the proposed wear rates [ 3, 5, 61. Dissolution wear theory is thermodynam- ically feasible, but it remains to be seen whether it adequately describes the cratering process [ 71.

Beyond this, most of the experimental work has involved long-time cutting. Cutting beyond the time at which a crater has been developed, and a stagnant material layer has been built up, poses a three-body problem. That is, periodic removal and reconstruction of the built-up layer, interaction between the tool material and the built-up layer and different sliding condi- tions between the chip and the built-up layer mutually interfere and several competing actions occur at the chip-tool interface. Observations after long- time cutting reveal only the total consequences or some partially preserved phenomenology that is difficult to interpret.

The ambiguities of long-time cutting tests have motivated a study of short-time cutting wear with the emphasis on careful delineation of chemical composition and microstructure. The short cutting times were achieved with a grooved workpiece providing interrupted tool contact. Discrete cutting intervals from 1 to 11 s were studied. A WC-(Ti, Ta, W)C-Co tool was used with 52100 and 1045 steel workpieces. Initial cratering wear was observed by means of scanning electron microscopy (SEM). Beyond this, wear propagation conditions were simulated in tests of up to 6 min cutting time, undertaken with WC-(Ti, Ta, W)C-Co tools that were coated with a 0.2 E.rm layer of iron and were subsequently annealed to allow the iron to diffuse into the tool. The wear of the iron-coated tools was observed with SEM and secondary ion mass spectrometry (SIMS) techniques. This metallurgical treatment approximates tool surface conditions during mid-lifetime wear propagation, without introducing the uncertainties of long-time test history. It has been proposed [8] that diffusion alone cannot account for total tool material loss because such diffusion activates secondary mechanisms. That is, if iron diffuses into cobalt, it enhances the formation of the embrittling r) phase, and formation of this phase at the grain boundaries can facilitate intergranular fracture of carbide grains on the tool surface. Because the proposed process periodically repeats, the diffusion depth required to cause grain segregation would be very shallow. Thus, if iron is diffused a priori into the tool surface layer and a short-time test is undertaken, the tool can be expected to behave like a tool that has been in use for an extended time.

2. Procedure

2.1. Materials The tool material was Carboloy (registered mark of General Electric

Company, Schenectady, NY, U.S.A.) grade 370 cemented carbide, a tough

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wear-resistant grade used for heavy-duty roughing cuts. The nominal chemical composition of the Carboloy 370 tool material is, by weight, 72% WC, 8% TiC, 11.5% TaC and 8.5% Co. The tool material has a density of approximately 1.26 X lo4 kg me3 and a hardness specification of 90.7 - 91.5 HRA.

Two workpiece materials were used, 52100 steel and 1045 steel. The 52100 steel was hot rolled, spheroidized and annealed bar with a measured hardness of 85.7 - 91.0 HRB. The nominal composition of the 52100 work- piece is given in Table 1. The second workpiece, mill-annealed 1045 steel, had a measured hardness of 71.0 - 75.5 HRB. The chemical composition is given in Table 1.

TABLE 1

Chemical compositions (weight per cent) of the workpiece materials

Element 1045 steel 52100 steel

c 0.47 Mn 0.67 P 0.008 S 0.014 Si 0.19 Ni 0.11 Cr 0.20 MO 0.05

0.98 - 1.10

0.25 - 0.45 0.025 maximum 0.025 maximum 0.20 - 0.35

1.30 - 1.60

2.2. Materials characterization The scanning electron microscope used was an ISI-Super HIA. Micro-

photo~phs were taken of the worn surfaces and elemental analysis of worn areas was obtained using a Tracer Northern energydispersive system. In- depth SIMS analysis of the tool was obtained using a Cameca IMS 3-F ion microscope. The primary beam was O*+with an impact energy of 8 kV.

2.3. Turning technique The turning tests were carried out on a 15.5 h.p. (11560 W) Lodge and

Shipley lathe with a variable-speed spindle. The cutting speed was 600 sur- face ft min-’ (3.05 m s-l), the depth of cut was 0.050 in (1.27 X 10W3 m) and the feed rate was 0.0123 in rev- ’ (3.175 X 10e4 m rev-‘). No lubricant was used. The tool had a side rake angle of 5”, a back rake angle of 0” and a side cutting edge angle of 15”.

2.4. Initial wear tests The very early stages of tool wear were investigated using polished

cutting faces and very short cutting times. The rake face of the tool was hand polished using a diamond abrasive to obtain an optical quality finish.

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The particle size of the diamond abrasive was gradually decreased in several steps to submicron size to minimize any residuaf stress in the surface layer. The cutting times investigated were 1, 3, 5, 7,9 and 11 s. The workpiece was the 1045 steel.

The length of each cut was predetermined by the axial length of full- radius segments of a circumferentially grooved workpiece. For example, a 1 s cut on a bar of 6.487 in (16.48 cm) diameter turning at 600 ft min I (353 rev mini) with a feed rate of 0.0125 in rev-’ (0.0317 cm rev ‘) requires a segment of workpiece only 0.0735 in (0.187 cm) long. A deep groove can be cut 0.0735 in (0.187 cm) from the entry end of the workpiece to disengage the tool from cutting after 1 s. At this cutting speed a 1 s cut only covers about 6 rev of the workpiece. Steps were taken to minimize the effect of entry and exit transients on such a short cut. A cleaning tool with the same nose radius as the test tools (l/32 in or 0.0793 cm) was used to prepare the entry side of the workpiece so that the entire length of cutting edge becomes engaged at once and the maximum chip thickness is reached at the end of the first full revolution of the workpiece. The groove at the end of the cut was prepared with a parting tool and has essentially square edges. The nose portion of the tool therefore continues cutting about one more revolution after 1 s because of the 15” side cutting edge angles used, but the major portion of crater area investigated was not in contact with the chip during this extra revolution.

2.5. Wear propagation tests VVear prop~ation testing consisted of turning 1045 and 52100 steel bar

stock with iron-coated tools. For control purposes, uncoated tools were used. The tests were run for up to 6 min, with stops at intervals of 30 s for SEM examination. Before and after wear propagation testing, the tools were subjected to SIMS analysis.

Prior to these cutting tests each tool was mechanically polished and then coated with 0.2 pm of iron using a sputtering technique. The sputtering time for each tool was 6 min at 500 W forward power with 2250 V applied to a target 0.1524 m in diameter. FoIlowing the sputtering process, the tools were annealed to allow the iron to diffuse into the bulk of the tool. Two tools were annealed at each of three temperatures, namely 973, 1073 and 1173 K. Each tool was put in a high purity alumina crucible and placed in a vacuum induction furnace. After 6 min at temperature, the power was turned off and helium was introduced to facilitate cooling and to minimize further diffusion on cooling.

After annealing each tool was subjected to SIMS analysis and represen- tative profiles are shown in Fig. 1. Figure 1 is a plot of the relative counts of iron versus depth, where the depth is estimated from the SIMS sputtering rate and time. The extent of iron diffusion after 6 min is clear, with penetra- tion to about 0.4 pm at 973 K, about 1.8 pm at 1073 K and about 2.3 pm at 1173 K. An unannealed “Fe control” sample was analyzed and is also shown in Fig. 1. The 0.2 ym sputtered layer is quite apparent in the “Fe control” results.

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Depth (A-x lo-?

Fig. 1. Ion probe data showing the relative intensity of iron as a function of depth in iron- coated and variously annealed tools. The “Fe control” tool was sputter coated but not annealed.

Profilometer traces were taken on the tools to reveal the general wear geometry and/or the accumulation of iron.

3. Results

3.1. Initial wear tests Figure 2(a) shows an area of a cratered tool surface after 1 s cutting

time. Figure 2(b) shows the area in Fig. 2(a) after the accumulated iron has been removed by etching in hydrochloric acid. It should be noted that the triangular WC particles have undergone a degradation that manifests itself in the form of grooves or ridges that occur intragranularly. The rounded (Ti, Ta, W)C particles appear to be unaffected. It should be noted also that the grooves on the WC particles do not appear to be in the direction of chip flow. The (Ti, Ta, W)C particles appear to project above the WC particles. The carbide particles are the order of 1 - 2 E.trn in diameter.

A detailed examination of the worn tools after machining for 1, 3 and 5 s revealed that iron bonds preferentially to certain grains and not to others. Point-by-point analysis of the worn areas using energy-dispersive spectros- copy (EDS) showed that iron adheres to the WC particles and not the (Ti, Ta, W)C particles.

The chips generated during machining were examined for wear debris from the worn tools. Chip material was dissolved for 30 min in a 10%

Fig. 2. (a) Iron accumulation after 1 s cutting time and before etching. The chip flow is in the direction of the arrow. (b) The area shown in (a) after removal of the iron by etching

with HCl: A, WC; B, complex carbide.

sulfuric acid-distilled water solution. Each solution was filtered through a 0.8 pm polycarbonate filter. The filters were coated with gold and analyzed using SEM and EDS. Figure 3 shows the results of the analysis. Tungsten was found and the implication is that the WC debris is perhaps an order of mag- nitude finer than the particle size in the tool.

Using a direct SEM and EDS examination of the chips, (Ti, Ta, W)C particles were found on the underside of the chips. The (Ti, Ta, W)C par- ticles found were of the order of the original grain size. No WC wear debris was detected using this method.

A summary of the basic observations seen in the photomicrographs of the work tools after machining for 1 - 11 s is as follows.

(1) Iron adheres to the WC particles and not the (Ti, Ta, W)C particles. (2) Intragranular grooves or ridges are seen on the WC particles after the

accumulated iron has been etched.

(a) (b) Fig. 3. (a) Residual material from the dissolved chips (cutting time, 9 s). Ceramic material appears to be attached to a WC particle. (b) Analysis of the residual material from the dissolved chips.

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(3) The intragranular grooves on the WC particles are not observed to have a direct relation with the direction of chip flow.

(4) Complex carbide particles project above the WC particles. (5) The complex (Ti, Ta, W)C particles do not develop ridges or

grooves.

3.2. Wear propagation tests Figure 4 shows the typical appearance of the tools after several minutes

cutting time. The workpiece was 52100 steel. In addition to the crater wear per se, areas of iron accumulation (area 1) and refractory compound

accumulation (area 2) are displayed. The refractory compound accumulation was only observed when cutting the 52100 steel and was traceable to the microstructure of that stock. Specifically, the matter shown in area 2 of Fig. 4 was found by electron microprobe analysis to contain iron, calcium, aluminum and silicon. Actually, an iron layer was found to lie under the calcium, aluminum and silicon. Using the Debye-Scherrer X-ray diffraction technique (iron radiation) the calcium, aluminum and silicon layer was identified as CaAl$i,Os although the bands were broad and diffuse. X-ray diffraction also suggested the presence of CrN, and EDS analysis of the layer revealed the presence of chromium. The nominal CaAl,Si,Os is known as anorthite and exists over a wide stoichiometric range [9]. It is frequently observed as an inclusion in steels and was readily observable in the 52100 microstructure, as displayed in Fig. 5. It was not possible to observe com- parable levels in the 1045 steel microstructure and no accumulation of such refractory compound was observed on tools used to cut 1045 stock.

Figures 6 - 8 show the SIMS iron profiles for coated and annealed tools after machining with a 1045 steel workpiece. Also shown are the profiles (from Fig. 1) for the coated and annealed conditions. The machining time with the coated tool involves further iron diffusion and iron pickup. More- over, there is some iron pickup and iron diffusion on the uncoated tool. During machining, iron diffused about 0.14 I.crn for the 973 K annealed tool,

Fig. 4. Typical overview of the tool wear pattern (uncoated tool after turning for 6 min). Examination of area 1 showed a steel deposit and of area 2 showed CaAlzSizOs. The workpiece was 52100 steel.

(a) (b) Fig. 5. (a) An inclusion in steel 52100 stock. (b) EDS analysis of an inclusion in 52100 steel stock.

about 0.4 pm for the 1073 K annealed tool and about 0.97 pm for the 1173 K annealed tool. The diffusion depth for the uncoated tool was about 0.3 pm, nearly comparable with the initial profile of the tool sputter coated with 0.2 I.trn and annealed at 973 K.

lOE7

lOE6

IOES

lOE2

lOE1

lOE0 L 0

Untreated Cratered

10

Depth (A-x IO-$

Fig. 6. Ion probe data showing the relative intensity of iron as a function of depth in an iron-coated tool annealed at 973 K, a coated, annealed and wear-tested tool (973 K cratered) and an uncoated and wear-tested tool (untreated condition).

165

lOE5 -

lOE2 -

lOE1 -

L Untraated Cratered

,1073 "K Cratered

lOE0' I I

0 10 20

Depth(n'x10-3)

Fig. 7. Ion probe data showing the relative intensity of iron as a function of depth in an iron-coated tool annealed at 1073 K, a coated, annealed and wear-tested tool (1073 K cratered) and an uncoated and wear-tested tool (untreated cratered).

lOE1

I

10

I

20

Depth (Wx10m3)

I I

30 40

Fig. 8. Ion probe data showing relative intensity of iron as a function of depth in an iron- coated tool annealed at 1173 K, a coated, annealed and wear-tested tool (1173 K cra- tered) and an uncoated and wear-tested tool (untreated cratered).

166

Figure 9 shows the surface appearance of an iron-treated tool annealed at 1173 K after machining with a 52100 workpiece for 6 min. Figure 10 shows the results of machining with a 52100 workpiece for 6 min with a standard tool. Figure 11 shows the results of machining with a 1045 work- piece for 2 min 50 s with an iron-treated tool that was annealed at 1173 K. Figure 12 shows the results of machining with a 1045 workpiece for 2 min with a standard tool.

In summary, Figs. 9 - 12 show that WC particles undergo an intra- granular degradation and that intragranular degradation of the WC particles appears not to have a direct relation with the direction of chip flow. More- over, the (Ti, Ta, W)C particles appear not to be degraded and the (Ti, Ta, W)C particles project above the WC particles.

(a) (b) Fig. 9. (a) Iron accumulation on the tool surface after 6 min cutting time and before etching. The workpiece is 52100 steel and the tool has been coated with iron and annealed at 1173 K. (b) As for (a) but after etching in HCI to remove the iron.

(4 (b) Fig. 10. (a) Iron accumulation after 6 min cutting time and before etching. The work- piece is 52100 steel and the tool is uncoated. (bf As for (a) but after etching in HCI to remove iron.

50 s cutting time and before etching. The workpiece is 1045 steel and the tool has been coated with iron and annealed at 11’73 K. (b) As for (a) but after etching in HCl to remove iron.

Fig. 12. (a) Iron accumuiatjon after 2 min cutting time and before etching. The work- piece is 1045 steel and the tool is uncoated. (b) As for (a) but after etching in HCl to remove iron.

Profilometer traces were taken on the tools after 30, 90 and 150 s. Three different areas of the crater were examined, and anomalous variations in depth with time indicated the periodic accumulation and detachment of iron. The average wear depth for all traces is given in Table 2 in relation to tool treatment and workpiece composition.

4. Discussion

The micrographs for the initial wear tests (Fig. 2) and for the wear propagation tests (Figs. 9 - 12) show similar behavior and the argument is

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TABLE 2

Average wear depth from profilometer traces taken after 30, 90 and 150 s during wear propagation testing

Tool condition Average wear depth (pm)

52100 steel 1045 steel

Uncoated 0.38 0.60 Coated, annealed at 973 K 0.27 0.19 Coated, annealed at 1073 K 0.32 0.30 Coated, annealed at 1173 K 0.32 0.52

--__ -____

made that the short-time phenomenology is relevant to much of the tool wear cycle.

In the initial stages of wear iron bonds to the WC particles and not the (Ti, Ta, W)C particles. Some WC particles appear to undergo an intragranular degradation evidenced by surface grooving. This surface grooving is peculiar to the WC particles and not the (Ti, Ta, W)C particles. The intragranular degradation of the WC particles results in their wearing away. WC wear debris is found to be smaller than the original grain size. This wearing away of the WC particles has the effect of leaving the (Ti, Ta, W)C particles projecting above the WC particles. The size of the (Ti, Ta, W)C particles that are removed from the tool surface is of the order of the original grain size. The relative appearance of the WC and (Ti, Ta, W)C grains is somewhat com- parable with the observations of Naerheim and Trent [ 41.

Adhesion is believed to be a likely mechanism for this wear. Adhesion wear develops by the bonding of the workpiece material to the tool and the subsequent removal of tool constituents when the workpiece is pulled away [ 9, lo]. Crater wear could develop according to the following sequence of events.

(1) Plastic deformation of WC particles occurs at the surface, as shown by Brainard and Buckley [ 111.

(2) The roughened surface of the WC particles adheres to the chip. The adherence of iron to these WC particles may abet the tool-chip traction and microfracture in the particles.

(3) Removal of the WC particles causes the (Ti, Ta, W)C particles to protrude above the WC particles.

(4) (Ti, Ta, W)C particle removal occurs with chip flow as surrounding material is removed.

Abrasive wear theory for cutting tools is based on the concept of hard inclusions in the work material indenting the tool surface. In these studies two different workpiece materials were used, 1045 and 52100 steel. In the case of 52100 steel, hard CaAlzSizOs inclusions were in the workpiece and these appeared on the tool as an oxide aggregate layer. No similar level of inclusion was found in the 1045 steel workpiece material or on the surface of tools used to machine the 1045 workpiece. In spite of this the wear features observed on the tool were much the same for the 52100 and 1045

169

steel workpieces. Thus, the role of abrasion seems to be small. Inclusion involvement has been studied by Trent [12], Opitz and Konig 181 and Faulring and Ramalingam [ 131. Steel containing glass-like inclusions has been shown to produce minimal tool wear, whereas crystalline silicates have been associated with increased tool wear [13]. The broad diffuse bands dis- played in the X-ray diffraction analysis of the anorthite layer in the present work suggest some glassy character. To examine this further, bulk samples of CaAlzSizOs were prepared in the laboratory by melting a mixture of CaO, A1,03 and SiO* in a furnace (the melting point of CaAlzSizOs is 1873 K). It was not possible to produce glassy material by an air quench from the melt. Debye-Scherrer analysis confirmed crystalline CaA12Si20s structure. Thus there is no simple basis for expecting a glassy anorthite or reduced wear therefrom. Even so, the anorthite-containing 52100 steel shows a lower un- coated tool wear rate than does 1045 steel (Table 2).

The dissolution wear theory involves the concept that the tool is dis- solved away into the stream of the chip material [ 3,7]. The findings of these studies do not support this concept. The WC particles are shown to wear intragranularly in the form of ridges or grooves that develop on part of the WC grain and not on the entire grain. Iron adheres to the WC particles, form- ing a dead zone between the tool and the flowing chip. The dead zone should act as a barrier to any dissolution process. If the chip acted in a stream-like fashion, entire grains should be dissolved instead of preferential parts of some grains.

Apart from dissolution, diffusion has been cited as a basis for cutting tool wear. It has been suggested that iron diffuses into the cobalt binder and enhances the formation of complex carbides which reduce wear resistance [ 3, 91. This mechanism seems doubtful in the present case, however. First of all, the WC particles appear to break up intragranularly and fracture does not occur along grain boundaries or in the cobalt (or Fe-Co) matrix, as would surely be required by the above diffusion mechanism. Beyond this, SEM examination revealed the same microstructural changes in uncoated tools as in the iron-coated and diffusion-reacted tools. Actually, the wear rates cited in Table 2 show that the uncoated tools wear even faster than the coated tools, although the higher diffusion anneal temperatures are associated with higher wear rates than are the lower diffusion anneal temperatures.

The diffusion data in Figs. 1 and 6 - 8 warrant further consideration, however. Gregory [ 61 has made an extensive study of diffusion between iron and cemented carbide. He distinguishes between the diffusion of iron into a cobalt “interaction zone” and the diffusion of iron from the “interaction zone” into the tool material. Table 3 shows diffusion coefficients, for the present case, calculated from Gregory’s data and projected and measured depths of penetration. The depth of penetration is that depth where the composition of the diffusing species, iron, has fallen to two orders of mag- nitude below the concentration (100%) of the iron layer. The data in Fig. 1 are consistent with Gregory’s work. The interaction zone diffusivities given by Gregory are about one-and-one-half orders of magnitude higher than those for the simple case of iron diffusing into cobalt [lo].

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TABLE 3

Diffusion projections from ref. 6 in comparison with the data in Fig. 1

Annealing temperature

(K)

Fe diffusivity into

interaction zone (from ref. 6) (cm2 s ‘)

Calculated Fe Fc penetration

penetration from Fig. 1

(Pm) Mm)

973 4.8 x 10 l4 0.15 0.4 1073 7.82 x 10 ” 0.59 1.8 1173 7.85 x 10~ I2 1.86 2.3

The data of Figs. 6 - 8 reflect the accumulation of iron on the tool sur- face as well as diffusion of iron into the tool. Even so, it seems clear that the tool surface must reach the 900 - 1000 K range since the iron profile after 3 min turning in the untreated tools approaches the profile for 6 min anneal- ing with an iron coating. It is clear that the extent of iron pickup and iron diffusion is greater, during 3 min turning, for the higher diffusion annealing treatments. This could reflect higher turning temperatures or metallurgical changes involving increased diffusivity. The rates of wear were greater for the tools diffusion annealed at higher temperatures, even though the uncoated tool showed the highest wear rate (Table 2).

The surface grooving noted in the WC grains suggests slip or shear band development. Brainard and Buckley [ll] showed that WC wear involves frac- ture initiated by plastic deformation. Hibbs and Sinclair [14] and Green- wood et al. [ 151 have recently evaluated slip mechanisms in monolithic WC and WC-Co composites.

5. Summarizing remarks and conclusions

(1) In the initial stages of crater wear, iron accumulated from the work- piece adheres readily to the WC particles. Some WC particles undergo an intragranular degradation in the form of microfracturing. The (Ti, Ta, W)C particles do not degrade and thus project above the WC particles.

(2) WC wear debris smaller than the original grain size is generated. (Ti, Ta, W)C particles are found to leave the tool surface at a size of the order of the grain size.

(3) The primary mechanism of crater wear is believed to be adhesion and the related micromechanical fracture of the WC particles.

(4) Abrasive wear is inconsistent with the observed similarity between results for inclusion-containing 52100 steel and 1045 steel relatively free of inclusions.

(5) Despite evidence of substantial diffusion, diffusional wear models are inconsistent with the relative wear rates of coated and uncoated tools and with the absence of intergranular failure.

171

(6) Dissolution models are inconsistent with the observed coexistence of the iron coating and the WC particles.

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N. P. Suh, N. H. Cook et al., Enhancement of cemented carbide tool properties, NTIS Publ. PB-221908, Vol. 2,1973 (National Technical Information Service, U.S. Depart- ment of Commerce). B. Kramer and N. P. Suh, Tool wear by solution: a quantitative understanding, J. Eng. Ind., 102 (1980) 303. N. H. Cook and P. N. Nayak, Development of improved cutting tool materials, Tech. Rep. AFML-TR-69-1857, June 15, 1969 (Air Force Materials Laboratory). Y. Naerheim and E. M. Trent, Diffusion wear of cemented carbide tools when cutting steel at high speeds, Met. Technol., 4 (1977) 548. W. Dawihi and M. K. Mal, Contribution to the study of the deformation behavior and structure of WC-Tic-TaC-Co alloys, Cobalt (Engl. Ed.), 26 (March 1965) 25. B. Gregory, Surface interaction of cemented carbide tool material and Armco iron, Br. J. Appl. Phys., 16 (1965) 689. B. Kramer, An analytical approach to tool wear prediction, Ph.D. Thesis, Department of Mechanical Engineering, Massachusetts Institute of Technology, Cambridge, MA, 1979. H. Opitz and W. KSnig, On the wear of cutting tools, Proc. 8th Int. Machine Tool Design and Research Conf., Birmingham, 1967, Pergamon, Oxford, 1968, p. 173. The Making, Shaping and Treating of Steel, U.S. Steel Corporation, Pittsburgh, 9th edn., 1971, p. 53. H. W. Mead and C. E. Birchenall, Diffusion of Co6’ and Fe55 in cobalt, Trans. Metall. Sot. AZME, 203 (1955) 994. W. A. Brainard and D. H. Buckley, Dynamic SEM wear studies of tungsten carbide cermets, ASLE Trans., 19 (1976) 309. E. M. Trent, Cutting steel and iron with cemented carbide tools, J. Iron Steel Inst. London, Ser. A, 10 (1979) 1781. G. Faulring and S. Ramalingam, Oxide inclusions and tool wear in machining, Metall. Trans. A, 10 (1979) 1781. M. K. Hibbs and R. Sinclair, Room temperature deformation mechanisms and the defect structure of tungsten carbide, Acta Metall., 29 (1981) 1645. R. M. Greenwood, M. H. Loretto and R. E. Smallman, The defect structure of tungsten carbide in deformed tungsten carbide-cobalt composites, Acta Metall., 30 (1982) 1193.