On Chip Morphology, Tool Wear and Cutting Mechanics in Finish Hard Turning

On Chip Morphology, Tool Wear and Cutting Mechanics in Finish Hard Turning

On Chip Morphology, Tool Wear and Cutting Mechanics in Finish Hard Turning M.A. Davies, Y. Chou, C. J. Evans (Z), National Institute of Standards and ...

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On Chip Morphology, Tool Wear and Cutting Mechanics in Finish Hard Turning M.A. Davies, Y. Chou, C. J. Evans (Z), National Institute of Standards and Technology, Gaithersburg, USA Received on January 8,1996

A b s t r a c t Topography of surfaces produced in finish hard turning using cubic boron nrtride (CBN) tools is affected by a large number of factors including tool wear and the mechanics of the chip formarion process This paper shows first that tool wear rates are affected by interactions between the work material and the binder phase of the CBN tool For finish hard turning, low CBN content, ceramic binder tools give longer lives and better finish than high CBN content metallic binder tools For low CBN tools, wear rate is directly related to the rnicrostructure of the work material and to the CBN grain size SEM studies suggest that chip morphology is independent of work material microstructure, but varies with tool wear Orthogonal cutting tests show that, above a critical speed, segmented chips are formed by catastrophic localized shear and that chip segmentation spacing may be reflected in a modulation of the machined surface Segment spacing IS a function of depth of cut, rake angle, and surface speed, approaching a limiting value with speed Specific cutting energies decrease with speed, also approaching an asymptote A simple mechanical model gives reasonable predictions of segment spacing along the original surface, although a full thermo-plastic model will be required to account for other aspects of the chip formation process

Keywords: cubic boron nitride(CBN), tool wear, chip formation

1. I n t r o d u c t i o n Finish turning of hard steels ( S O Rc) is an economically attractive alternative to grinding in some applications' . The enabling technology for this process has been the availability of polycrystalline cubic boron nitride (CBN) tools, although some ceramic tools have also been used. A body of literature already exists reporting tool wear data in a range of materials; the focus of this paper is fundamental mechanisms of material removal and tool wear. In a series of preliminary experiments, through hardened AISI 52100 bearing steels were CBN turned on a conventional lathe2 and on two different diamond turning machines. These experiences indicated that tool wear rates and surface finish were essentially independent of the choice of machine, suggesting that tool/work material interactions dominate, at least over this range of machine characteristics. Konig indicated that some minimum machine stiffness is required, without givin the data on which that observation is based. Also Sen et al recently published data showing relatively low wear rates obtained with a high stiffness machine and low CBN content tool materials; however, these authors do not give comparative wear rates for the same process parameters on a machine of lower stiffness or for a range of tool materials. Clearly, a full understanding of finish hard turning will require knowledge of the complete system and of the mutual interactions between machine characteristics, material properties and process parameters. This goal first requires fundamental understanding of all the interactions.


2. Tool wear in CBN hard t u r n i n g CBN tool materials may be broadly classified into two groups high CBN content materials (-90%) typically with a metallic binder phase, and low CBN content (50 to 75%) usually in a ceramic binder Data from several sources show that for high stock removal rate operations (roughing), high CBN tools give longer lives For precision finish hard turning, however, low CBN tools give longer tool lives and consistently better surface finishes4 This surprising result, given the higher bulk hardness and fracture toughness of high CBN tool materials, has provoked a flurry of alternative hypotheses including' differences in bond strength', differences in tool thermal conductivity leading to increased thermal soften-

Annals of the ClRP Vol. 45/7/1996

ing in the shear zone6, and differences in defect densrty wrthin the CBN grains7 None of these hypotheses provide a selfconsistent explanation for all the observed behavior In particular, rt appears that insufficient attention has been given to possible effects due to the microstructure of tool and/or work material or to tribochemical effects, although both these areas have received increased attention in precision grinding and diamond turning'

2.1 Facing AISI 52100 Constant surface speed dry facing experiments were performed on a conventioral lathe on rings of through hardened (62Rc) bearing grade AlSl 52100 steel For each combination of parameters, five tools (0 8 m m nose radius, -30" rake, 5" clearance) were used at a fixed feed rate of 12 5 p-drev After each 260 m of cutting distance, maximum flank wear was measured in an optical microscope and land V()B, part surface finish measured using a profilornete? Morphology of the worn areas on the tools was examined in a scanning electron microscope (SEM) and chemical composition of selected tool surfaces investigated using x-ray photoelectron spectroscopy (XPS)" Four different tool materials from two suppliers were used Most of the experiments used either a high CBN (0 92 volume fraction CBN, cobalt binder, coded "Hl")" or a low CBN (0 70 vol CBN, titanium nitride binder, coded " L l " ) tool material both having the same nominal CBN particle stze distribution Some data were also taken with two other low CBN, ceramic binder tools ('L2" and 'L4"), the main difference between these latter materials is the CBN grain s u e 2.1. I Results Fig 1 shows surface finish and flank wear as a function of cutting distance for 2 m/s surface speed (V,) and 50 pn depth of cut Finish deteriorates with tool wear for all the experiments, as would be expected, finish obtained with the metal bonded high CBN tool is always worse (for identical conditions over the range tested) than the low CBN tools After some initial rapid wear, the wear rate is reasonably constant over the range tested In agreement with previously reported results, the data show that the high CBN metal bond tools showed more rapid wear than the low CBN ceramic


binder tools. The difference between low CBN tools is much less than between the tool types (ie high vs low CBN). Fig 2 shows flank wear rate as a function of surface speed and depth of cut for H1 and L1. Flank wear rate increases monotonically with speed for both tool materials. The effect of changing depth of cut is more pronounced at low depth of cut Low magnification scanning electron micrographs show a crater zone, flank wear, and a transferred layer on the flank of every tool. At higher magnifications it is clear that the crater and flank wear zones have an essentially polished character, with no evidence of large scale fracture. The character of the transferred layer, however, is significantly different between the high and low CBN tools, and that character vanes with process conditions. After approximately 2.6 k m cutting distance, for L1 the transferred layer is smooth and uniform while on H I the layer is rougher and has a grooved structure. With increasing cutting speed the layers get thicker, with the groove structure becoming more pronounced on H1. Only at the highest speed were any grooves detected on L1. High magnification shows that the transferred layer on L1 tools has a flake like structure, suggesting that the layer is less well bonded to the tools than the H1 layers which show no such structure. Etching in concentrated sulphuric acid ( z 55%) easily removed the L1 layers; even after long leaching times, residual layers are found on H I tools. The surface structures below the leached off layers are also different. The L1 tool surfaces were characterized largely by smooth CBN grain surfaces, with occasional shallow pockets, presumably where worn grains have been plucked out due to loss of bonding. This suggests adhesive wear. The grooved surface seen in H1 below the transferred layer, however, suggests abrasive wear mechanisms. XPS spectra of the transfer layer show strong iron and oxide peaks, as expected. After leaching, there is no evidence of the iron peak; the H1 surface indicates the presence of silicon dioxide which is not detected on L1.

0.7 i


0.6 0.5


#---. E 0.4 3.



0.2 0.1 n l






1000 1500 2000 Cutting distance (m)


500 --e---

-E A




- L2

400 300


iz 9 200 100

0 (b/



i000 1500 2000 Cutting distance (m)


Fig 1 (a) Surface finish and (6) flank wear for high CBN (Hl) and low CBN content tools on 62Rc AlSI 52100 78











Surface speed (mls)







E --. 2d





a, L




60 c



30 0





150 Depth of cut (pn)




f i g 2 flank wear rate as a function of (a)surface speed and (b) depth of cut when machining 62Rc AlSI 52100 2.1.2 Discussion The resutts given above suggest that interactions between the binder and the work matenal can dominate wear rates in finish hard turning The grooves in H? (which should not be confused wtth Pekelharing grooves at the feed rate) suggest that abrasive wear also affects tool wear rates This abrasion may arise either from plucked out CBN grains or as a result of hard phases in the work material Similarly, the systematic vanation in wear rate among different low CBN content tool materials suggest a microstructural effect 2.2 Workpiece and fool micrasfructural effects Surprisingly little attention has been paid to microstructural effects in CBN turnin For high stock removal rate processes, Chryssolouris” showed that tool wear depended on percentage martensite and the pe, see, and composition of IY the hard phases Takatsu et al tested a range of combinations of tool and work materials and indicated that workpiece microstructure affected adherent layer formation and hence wear rates, although no explanation for the mechanism IS given EvansI4 used a powder metallurgy steel, with a highly refined microstructure, in an attempt to reduce the contribution of abrasion in diamond tool wear Here we attempt to quantify that idea by comparing tool wear rates for three steels of the same chemistry but different microstructure Through hardened (62-64 Rc) bars (25 m m nominal diameter) of VIMVAR (vacuum induction melt vacuum arc remelt), PM (powder metallurgy), and conventional AlSl M50 tool steel were turned (V, = 2 m/s, feed = 12 5 @rev, depth of cut = 50 lrn ) in the conventional lathe2 using three L3 tools for each work material Fig 3 shows the microstructures of the materials Tool wear and finish were measured (Section 2 1) and chips were collected for SEM observation To evaluate the effect of tool material microstructure conventional MSO bars were turned using L2, L3 and L4 tools under the conditions given above These tool materials are from a single vendor have 60% CBN content and the

same binder , and differ microstructurally only in the CBN grain sizes, nominally 3, 1 and 0.5 pm respectively.

500 -

2.2.7 Results Fig 4 shows flank wear as a function of cutting distance for the three work materials. After the initial rapid wear, both the VIMVAR and conventional M50s give similar, reasonably stable, wear rates. The PM steel, by contrast, shows almost an order of magnitude lower wear rate. Fig 5 shows that tool wear rate decreases with decreasing CBN grain size; note, however, that bulk hardness and transverse rupture strength also increase with decreasing grain size. The chips produced in these and other cutting experiments were studied in the SEM at a range of different magnifications. All chips showed a banded structure, and the lamellar spacing varies across the chip. Chip segments have virtually identical spacing at the start of the experiment. With increasing cutting time, chips from PM M50 show no noticeable difference. However, the segment spacing of conventional and VIMVAR M50s change substantially. This implies that cutting mechanics do not change with workpiece microstructure, but chip morphology alters with tool wear presumably due to changes in tool geometry or tool wear induced changes in thermal conditions

400 300 -

200 loo




2000 3000 4000 5000 6000 Cutting distance (rn)

Fig 4 Tod wear rate for various M50 microstructres

5001 400


E rt





SF 2001

0' 0

1000 2000 3000 4000 5000 6000 Cutting distance (in)

Fig 5 Wear as a function of CBN grain size

f i g 6 Delamination on tool flank

f i g 3 Microstructure of 3 M50 steels (a) conventional, (b) VIMVAR, and (c) PM

2.3 Discussion Tool wear in precision finish hard turning is affected by a number of probably interacting variables. We have not been able to separate the effects of binder material and CBN content; neither low CBN metal matrix or high CBN ceramic matrix tools were tested. More work is required to clarify the chemical interactions between the matrix and work material. Workpiece microstructure and CBN grain size have both been shown to be important factors in tool wear rate. Typical microstructure for M50 is ultrahard molybdenum rich carbides (M&) in a martensite matrix. The conventional and VIMVAR materials have comparable carbide sizes, surprisingly similarly segregated. The uniformly distributed carbides in the PM M50 are an order of magnitude smaller. Compton et all5 showed the role of fatigue and microfracture in diamond tool wear under some circumstances. Our results, combined with the lower forces from interactions with smaller MsC particles, and the higher fradure toughness of smaller CBN particies all point to a similar mechanism in 79

finish hard turning with low CBN tools This idea is supported by delaminations observed in the flank wear area (Fig 6) The observed variation in chip lamellar spacing wtth tool wear may be a useful diagnostic of the machining process, particularly as residual stressesI6, surface finish, and part tolerances are all related to tool condition Note that the chips are segmented and that, around the arc of a radiused tool, segmentation frequency changes A simpler cutting geclmetry is clearly required to understand these effects

3. Orthogonal cutting and chip mechanics To better understand chip morphology and the basic mechanics of chip formation, a set of orthogonal cutting experiments was conducted. Changes in chip morphology, forces, and workpiece topography were measured 3.7 Experimental set-up Experiments were conducted on a two-axis, air bearing spindle diamond turning machine The workpiece material was through hardened (62 Rc) 52100 bearing steel Orthogonal cuts were performed on the edge of a 0 5 m m wide 94 m m diameter ring using custom ground and lapped L4 tools mounted on a three-axis dynamometer Forces parallel (F, ) and perpendicular (Ft ) to the cutting direction were measured Chips were mounted, polished in cross-section and etched using a 2% Nital solution In general, chips were found to be segmented, and the spacing between segments (denoted d in Fig 12) taken as indicative of the dynamics of the cutting process Segment spacing was measured erther from the chip cross-sections or by mounting the chips "smooth" face down in an SEM Chip cross-sections were also digitized 3.2 Experimental results Fig 7(a) shows an etched cross-section of a highly segmented chip formed at a cutting speed, V,, of 1 0 m/s, a depth of cut, t, of 30 pm, and a rake angle, cx, of -27 Fig 7(b) shows a magnified view of the region between two segments Hiqhiy localized shear flow is evident, indrcatrnq that the segmented chips form due to strain local~zation'~ rather than through fracture'' Many other materials including titanium2' , steels2' and nickel-iron alloys2' show similar behavior Catastrophic strain localization occurs when thermal softening behavior dominates over the combined effects of strain and strain rate hardeningz3 These conditions are typicalk met when the strain rate exceeds a critical value determ i i d by material properties Thus, for a given cutting geometry, the onset of strain localization and segmented chips is typically controlled by the cutting speed

I -/

the latter data were obtained following improvements in tool lapping that produces a sharper cutting edge which probably accounts for the systematic difference However, both sets of data show a similar trend Segment spacing increases monotonically with speed but appears to be asymptotic to some limiting value, and the onset of segmentation occurs at a cutting speed of between 0 3 m / s and 0 7 m/s Decreasing the rake angle and depth of cLt both cause a marked decrease in segment spacing apd may also effect the critical cutting speed for the onset of segmentation Note that chip segmentation patterns may be seen on the workpiece as well Fig 9 is a Nomarski micrograph of the surface of a workpiece followng an orthogonal cut (V,= 1 5 m/s, a = -27O and t = 30 pm) and a chip obtained at the same cutting conditions and at the same magnification Clearly the frequencies of the two are quite comparable The pattern on the workpiece indicates the frequency of the disturbance to be 250 kHz,significantly above any reasonable first mode of the structural loop





Rake = -27 t = 30 (Previous Edge Prep) Rake=-10 t = 3 0 u m m Rake = - 2 7 , t = 15 um e - Rake = -27 t = 30 urn T






2 3 Cutting Speed (m/s)



f i g 8 Chip segment spacing vanation

100 Itm


50 pm

Fig 7 SEM micrographs of etched chip cross-sections Fig 8 shows the variation of mean chip segment spacing as a function of cutting speed for two rake angles and two uncut chip thicknesses The data shown by the solid circles24 and the solid diamonds were obtained at t = 30 pm The difference between the conditions is that and a = -27


Fig 9 Nomarskl micrographs showing part and chip penodicify Fig 10 shows specific cutting energy as a function of cutting speed Specific energy initially decreases with speed and then levels out The inrtial decrease is due to the effect of the highly localized thermal softening that occurs in segmented chip formation The leveling of the force curve at higher speeds coincides with the corresponding leveling out of the segment spacing curves (Fig 8) The data also indicates slightly higher specific energy values for smaller t at u = - 2 7 , and a substantially lower energy for u = -10" and t = 30 pm During the formation of a segmented chip the forces on the tool should oscillate at high-frequency as the segments are formed The spacing between chip segments can be measured along the original free-surface of the workpiece

(or from the chip cross-sections) and hence the frequency of the expected force variation can be estimated. For the chips studied, the frequencies were 50-400 kHz. These frequencies far exceed the b a n h d t h of the dynamometer, which therefore measures only the average forces. Hence accurate prediction of cutting forces demands understanding of the dynamics of the chip formation process.


Rake = -27, t = 30 um Rake = - 2 7 , t = 15 urn - - m - - Rake = -10, t = 30 urn


7 -




loo00 8000 6000 Cutting Speed-(m/s)

thermo-mechanical behavior of the material being cut, may be necessary However, a simple mechanical model developed here offers reasonable predictions about the limiting values of segment spacing at high speeds Fig 12 shows a chip segment in formation Catastrophic shear was initiated along the line BID' when the tool tip was at €3' As the tool moved from 8' to A, the stress required to cause further deformation along BD decreases due to thermal softening, and the tool experiences rapid unloading along BE Simultaneously, the loads increase along AB as the matenal in region B'AB IS indented and sheared The thermal state of the workpiece ahead of the tool and the stresses applied along both AB and BD determine when the next chip segment wll form. If the cutting speed is significantly above the critical value for s q mentation, loads transmitted along BD decrease rapidly to nearly zero and can be ignored Additionally, shear zone formation IS nearly adiabatic and the workpiece temperature ahead of the shear zone remains essentially at ambient Thus, the formation of the next segment is determined solely by the stresses applied to the workpiece along A 6

Fig 10 Speafic cutting energy as a function of speed

Tool 1

-50' 0

250 500 750 Distance along chip, pm



-1.0 m / s

3-1 mis






Work piece

Fig 12 Simple model o f segment formation (a shown as -1 07 Consider a material element bounded by the lines

AB, BD, the free-surface of the workpiece, and a line AC in-

f i g 1I Chip proriles and their power spectra Insight into the process dynamics can be gained by considering the chip as a spatial record of the temporal dynamics during cutting Segment patterns become measurably more penodic as the cutting speed is increased Fig l l ( a ) shows profiles of serrated chip edges obtained by digitzing chip cross-sections The lower signal is from a chip obtained at v, = 3 1 d s , a = -27 ', and t = 30 jtm The upper trace is from a chip cut with the same geometry at vc = 1 0 m/s The power spectra of the two chips (Fig l l ( b ) ) clearly indicate the periodic nature of the lower chip trace and the aperiodic nature of the upper trace The trend from ordered to disordered segment spacing as vc decreased was observed for all cutting geometries Indeed, althou h not emphaslzed, similar trends are seen in the literature


3.3 Mechanical Models of Chip Segmentation It is possible that the transition from ordered to disordered segmentation patterns may result from nonlinear dynamics of the chip formation process Recently, relatively simple nonlinear dynqmic systems have been shown to exhibit various types of periodic and aperiodic behaviors that may change abruptly as system parameters are varied25 Thus to predict segment spacing and cutting forces accurately a fbil dynamic model of segmented chip formation, including the

clined at an angle y/ relative to the horlzontal For \v>\v' (see Fig 12), C wll lie on the shear zone BD, while for y/<\v'. C wll lie on the free-surface of the workpiece Fig 13 is a free body diagram of this material element for y c y ' , where G~ IS the average normal stress applied by the tool on AB, and T and G are the average shear and normal stresses respectively acting on AC For simplicity, we ignore shear stresses (fnction) along AB From geometry, we can determine dot as a function of y ~ 8, and the tool position In general, this ratio wll have local maxima at two values of w VI u,l2

= d4 + ([email protected]' = d 4 +cx/2

As the tool moves forward from B',dot will initially be maximlzed at V / I the larger of the two angles Failure of the material may occur along this line forming a secondary shear zone that intersects the primary zone BD As the tool moves forward and the length of A 6 increases, the shear stress along the line defined by y/ = y12 will grow and eventually exceed that along v = y~~ At this point, the shear zone wll jump and a new segment will be formed with an angle \v = \ v 2 relative to the horlzontal Using 0 = qt2, the spacing between segments is given by, d = d' = t [sec( d4 - al2) - tan (d4- u/2)y(1+ sin u)


This indicates that at high speed, segment spacing IS a linear function of the uncut chip thickness, as has recently been observedx 27 Eq(l)givesd=d'= I36 ~ ~ f o r a = - 1 0 ° a n d t = 3 0 itm and d = d' = 15 5 p m for = -270 ,t = 30 pm Thls com81

pares to experimental measurements of d = 26 5 pm and d' = 17 6 pm for u = -1O", and d = 58 p m and d' = 21 2 p m for a = -27" Although the predicted limiting values of d' agree reasonably well with measurements, the predicted values of d do not This ISprobably due to the flow of material from the region B'AB (Fig 12) into the region between segments as the chips form This is not considered in the model In addition, it is likely that failure in the secondary shear zone defined by y~ = \v1 is also catastrophic and that this zone may actually combine with the primary shear zone BD This would lead to a downward curvature in the shear zone in the vicinity of the tool tip, a phenomenon that has been observed in chips formed at high cutting speed

A Fig 13 Free body diagram of element ABD in Fig 12 3.4 Discussion Segmented chip formation is a highly nonlinear dynamic process that can affect cutting forces, machine deflection and surface finish. Our data indicate that chip segmentation patterns are a strong function of cutting parameters. It is likely that variations due to uncut chip thickness and rake angle are primarily geometric: however, variations due to changes in cutting speed are related to the dynamics of chip formation. W e are developing a 1D thermo-plastic model to explain some of the effects reported here. 4. C o n c l u s i o n s Throughout the experiments reported here, surface finish is correlated, but not closely, with tool wear. For high CBN metal matrix tools, interactions between the binder and the work material produce highly adherent layers; bond failure between CBN and matrix presumably leads to the pluck out of CBN grains which cause significant abrasive wear on the tool flank. With the low CBN ceramic matrix tools, less adhesive layers and better bonding of the CBN particles in the matrix result in a fine scale attrition dominating the wear process. The reduction in wear rate associated with decreasing CBN grain size and with decreasing work material carbide size both point to initiation and propagation of microfractures as the main mechanism involved in this fine scale attrition. Segmented chips are produced by catastrophic strain localization above some critical speed. As speed increases, segmentation becomes increasingly ordered, with mean spacing increasing asymptotically with speed, depth of cut, and rake angle. Specific cutting energies also decrease asymptotically with speed. A simple mechanical model predicts limiting segment spacing, but other features of the cutting process will require more complex models currently being developed. 5. A c k n o w l e d g e m e n t s The authors thank E. Whitenton, L Smith, P. Sullivan, L. lves and K. Harper for their contributions. M.A.D. acknowledges the NRC Postdoctoral Fellowship that supported his work.

6. R e f e r e n c e s 1 Konig, W . et al "Turning vs Grinding", 1993, ClRP Annals, Vol42, pp 39-43


2 Hardinge Superslant. Specific commercial products are identified in order to fully describe the experimental procedure. Such identification does not imply endorsement by the National Institute of Standards and Technology nor imply that they are necessarily the best for the purpose 3 Sen P.K., Keisler R., Vavrille J-L, and Shore P. "Precision Machining of Hard Ferrous Matenals with PCBN Tooling using the "ION' Machine" 1995, Proc IPES 8, Wiley 4 Chou, Y., Barash, M., and Evans C. "Finishing Performance and Wear Mechanisms of Cubic Boron Nitride in Turning Hardened 52100 Steels" 1995, Ceramic Industry Manufacturing Conference and Exposition, Pittsburgh, Pa 5 Narutaki Y. and Yamane Y. "Tool Wear and Cutting Temperature of CBN Tools in Machining of Hardened Steels", 1979, ClRP Annals, Voi 28. pp 23-8 6 Bossom P. K. "Finish Machining of Hard Ferrous Workpieces", 1990, Industrial Diamond Review, pp 228-32 7 Hooper R. M., Shakib J. I. Brookes C. A " Microstructure and Wear of Tic-Cubic BN Tools', 1988, Materials Science and Engineering, Vol 105'1 06, pp 429-33 8 Paul E., Evans C. J, Polvani R. S., McGlauflin M. L. and Mangamelli A "Chemical Aspects of Tool Wear in Single Point Diamond Turning" Precision Engineering, in press 9 Federal Instruments EMD-1500 (see Note 2) 10 Chou Y. "Wear Mechanisms of Cubic Boron Nitride Tools in Precision Turning of Hardened Steels", 1994, Ph.D Thesis, Purdue University 11 H1 = G E BZN 6000, L1 = GE BZN 8100; L2, L3 & L4 are Sumitomo BN 200, 250 and 300 respectively (See Note 2) 12 Chryssolouris G. "Turning of Hardened Steels using CBN Tools", 1982, J. Appl. Metalworking, Vol 2, pp100-6 13 Takatsu S., Shimoda H., and Otani K. "Effect of CBN Content on the Cutting Performance of Polycrystalline CBN Tools", 1983, Int J. Refract. & Hard Metals, Vol 1 2 , pp175-8 14 Evans C. "Cryogenic Diamond Turning of Stainless Steels", 1991,ClRP Annals, Vol 40, pp 571-5 15 Compton, D., Hirst, W. and Howse, M.G.W. 'The Wear of Diamond", 1973, Proc R SOCLond A 1973, 333, 435454 16 Tonshoff H. K., Wobker H-G., and Brandt D. "Tribological Aspects of Hard Turning with Ceramic Tools" ,1995, J. SOC. Tribol. & Lubric. Engineers, pp163-8 17 Recht R. F. "Catastrophic Thermoplastic Shear" (1964) J. Appl. Mechanics, Vol 31 ppl89-93 18 Konig W., Klinger M., and Link R."Machining of Hardened Steels with Geometrically Defined Cutting Edges - Field of Applications and Limitations" 1990, ClRP Annals, Vol 39, pp 614 19 Shaw M. C. and Vyas A. "Chip Formation in lylachining of Hardened Steels" 1993, ClRP Annals, Vol 42, pp 29-32 20 Komanduri R. "New Observations on the Mechanism of Chip Formation when Machining Titanium Alloys" 1981, Wear, Vol69, pp179-88 21 Komanduri R., Schroeder T., Hazra J., von Turkovich B. F. and Flom D. G. "On the Catastrophic Shear Instability in High Speed Machining of an AlSl 4340 Steel", 1982, J. Eng. for Industry, Vol 104, pp 121-31 2 2 Komanduri R. and Shroeder T A. "On Shear Instability in a Nickel-Iron Base Superalloy" 1986, J. Eng. for Industry, Vol 108, pp 93-1 00 23 Wright T. W. and Walter J. W. "On Stress Collapse in Adiabatic Shear Bands" 1987, J. Mech. Phys. Solids, Vol 35, pp 701-20 24 Davies M. A., Evans C. J., Harper K. K. "Chip Segmentation in Machining AlSl 52100 Steel" 1995, Proc. ASPE, pp 235-238. 25 Moon F. "Chaotic and Fractal Dynamics" 1992, Wiley 26 Arcona C. and Dow T. A "Development of a Tool Force/Surface Finish Model in Diamond Turning" ,1995, Proc ASPE, p p l 3 - I 6 27 Ueda K. and Manabe K. "Chip Formation Mechanism in Microcutting of an Amorphous Metal" 1992, ClRP Annals, Vol 41, pp 129-32