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Wear 249 (2001) 473481

Workpiece surface integrity considerations when finishturning gamma titanium aluminide

A.R.C. Sharman a,, D.K. Aspinwall b,c, R.C. Dewes b, P. Bowen a,ca School of Metallurgy and Materials, University of Birmingham, Edgbaston, Birmingham B15 2TT, UK

b School of Manufacturing and Mechanical Engineering, University of Birmingham, Edgbaston, Birmingham B15 2TT, UKc Interdisciplinary Research Centre (IRC) in Materials for High Performance Applications, University of Birmingham,

Edgbaston, Birmingham B15 2TT, UK

Received 15 September 2000; received in revised form 14 March 2001; accepted 16 March 2001

Abstract

Gamma titanium aluminides (-TiAl) are currently being evaluated by aeroengine manufacturers as possible replacements for conven-tional titanium alloys and nickel based superalloys in gas turbine engines. Unfortunately, even when machining using operating parametersselected to minimise workpiece damage, turned surfaces contain cracks and microstructural alterations. The paper initially reviews priorwork on the surface integrity of -TiAl produced by various machining processes. Following on from this, surface integrity data from a fullfactorial experiment are presented when turning Ti45Al2Mn2Nb+0.8 vol.% TiB2 employing extremely fine finishing cuts. Workpiecesurfaces/subsurfaces were evaluated in relation to microstructural alterations, strain hardening/microhardness changes, and 2D surfaceroughness (Ra).

All the machined surfaces contained cracks, cracked TiB2 particles, deformed microstructure and microhardness increases regardless ofthe operating parameters. However, when comparing the surfaces with those obtained in previous work using more abusive parameters, asignificant reduction in surface damage was obtained. 2001 Elsevier Science B.V. All rights reserved.

Keywords: Surface integrity; Gamma titanium aluminide; Turning

1. Introduction

Research on titanium aluminides (TiAl) began in the1950s, however, it is only over the past decade that sig-nificant industrial interest has occurred primarily in theaerospace and automotive sectors, due to the developmentof two phase gamma titanium aluminide (-TiAl) [1]. Incomparison to nickel based superalloys, the strength andductility of TiAl materials puts them at a disadvantage,however, when the properties are normalised by density, thepicture is more promising, see Table 1 for selected mechan-ical/physical property data [14]. In terms of weight, thedifference between a conventional titanium alloy such asTi6Al4V and a representative -TiAl alloy is marginal,however, the ability to operate at significantly higher tem-peratures is attractive. From a mechanical design stand-point, the low room temperature ductility of -TiAl, whichtypically ranges between 0.3 and 4% depending on the al-loy composition and microstructure, together with its low

Corresponding author. Tel.: +44-121-414-3541;fax: +44-121-414-3958.E-mail address: [email protected] (A.R.C. Sharman).

fracture toughness are of concern. The highest ductilityis associated with duplex, as opposed to fully lamellarmicrostructures which have lower ductility, lower tensilestrength but higher fracture toughness (2035 MPa m1/2)[1,5] and lower fatigue crack growth rates [6]. Conse-quently, much of the current research effort has focused onlamellar materials due to their more favourable balance ofproperties. On a more positive note, the majority of metalshave fatigue run out strengths around 5060% of their yieldstrength, however, -TiAl alloys exhibit values approachingtheir yield value [7].

All the worlds leading aeroengine manufacturers, whichincludes RollsRoyce, General Electric (GE) and Prattand Whitney, have research programmes evaluating -TiAlalloys. Although applicable for engines used in both com-mercial and military applications, much of the presentinterest relates to commercial aircraft in order to reducefuel consumption. A great deal of the research work is atan advanced stage and reported demonstrator componentsinclude missile fins [8], low pressure (LP) turbine blades[9,10], transition duct beams, blade dampers, high pressure(HP) compressor blades, LP turbine side plates, nozzleguide vanes and other more esoteric engine components

0043-1648/01/$ see front matter 2001 Elsevier Science B.V. All rights reserved.PII: S0 0 4 3 -1 6 48 (01 )00575 -0

474 A.R.C. Sharman et al. / Wear 249 (2001) 473481

Table 1Properties of competing aerospace alloys [14]Property Material/condition

Ti45Al2Mn2Nb+ 0.8 vol.% TiB2(near lamellar)

Inconel 718(solution treatedand aged)

Ti6Al4V(solution treatedand aged)

Density (g/cm3) 3.76 8.4 4.46Specific modulus (GPa/Mg/m3) 45 25 26Tensile strength (MPa) 672 1430 1035Specific strength (MPa/g/cm3) 179 170 232Yield strength (MPa) 548 1190 965Ductility (%) 1.6 21 8Fracture toughness (MPa m1/2) 20 61.5 4466Thermal conductivity (W/mK) 22 11.4 7.5Maximum operating temperature (C) 750 1000 600

[10]. Engine tests conducted by GE, Volvo, Allison andPratt and Whitney have reported successful -TiAl results,indeed in the GE programme, it was found that cracks inLP turbine blades resulting from mishandling did not prop-agate during cycling [10]. Despite this and other favourableoutcomes, -TiAl has yet to fly in passenger service.

On the automotive front, potential use of -TiAl applica-tions include engine valves and turbo impellers. Understand-ably, in view of the more stringent safety requirements ofthe aerospace industry, the transition from research materialto production component has been faster in the automotivesector, atleast with Formula 1 engines. The authors are awareof only one production vehicle containing a -TiAl tur-bocharger manufactured by a Japanese automotive company[11], possibly as a consequence of the significantly highermaterial and processing costs than are currently encountered.

1.1. Machinability of -TiAl

Numerous studies have shown conventional titaniumalloys to be significantly more difficult to machine thanthe vast majority of workpiece materials regardless of themeasure employed [1217]. This is due to the fact that ti-tanium has low thermal conductivity, retains its strength toelevated temperatures and has a high chemical affinity forall tool materials. Cutting speeds are generally limited to

A.R.C. Sharman et al. / Wear 249 (2001) 473481 475

confined to 100m below the workpiece surface [23]. Incontrast, crack free HSM surfaces have been found to con-tain significant lamellae deformation and strain hardened re-gions [24], however, the deformation has been shown to bebeneficial to fatigue life. Four point bend fatigue test resultsfor HSM surfaces have produced run out peak stress val-ues of 700 MPa compared to 500 MPa for ground sur-faces [25]. The greater deformation in the subsurface of theworkpiece was found to correspond with highly compres-sive residual stresses that would be expected to retard crackinitiation during fatigue cycling.

Previous work on -TiAl has revealed that regardless ofalloy composition, all turned workpiece surfaces containeddefects of a similar magnitude, although the first genera-tion 4822 alloy, had cracks running down the interfacebetween adjacent lamellae laths [29]. Splitting of lamellaeis often seen in fracture toughness and fatigue testing and isattributed to low adherence between adjacent lamellae. Thecracking problem appears to have been resolved in secondgeneration alloys such as 45220.8, which employs ad-ditions of boron. This forms in situ titanium diboride parti-cles and causes grain refinement and enhanced mechanicalproperties.

The following experimental work was undertaken to eval-uate the effects of extremely fine finish turning parameterson workpiece surface integrity when machining 45220.8.

2. Experimental

2.1. Workpiece materials and equipment

The workpiece material was a grain refined -TiAl ofcomposition Ti45Al2Mn2Nb + 0.8 vol.% TiB2 XDTMcast by Howmet, USA, into bars 150 mm diameter300 mmlong. This was HIPped at 1260C/170 MPa for 4 h, then heattreated at 1010C for 50 h to improve mechanical proper-ties. The material had a near fully lamellar microstructurewith a 50100m grain size. All the machining trials wereconducted on a MHP MT80 CNC turning centre which hadan internal 20 bar6 l/min cutting fluid supply. A high pres-sure environment was supplied by an external ChipblasterRV1100 unit with an output of 65 bar26 l/min. The toolsused were uncoated WC (K10 ISO grade) indexable insertswith a +5 side and back rake angle, 5 clearance angleand 45 tool cutting edge angle (ISO classification SNMG120408-23).

Sections of the machined workpiece were cut out of thebars using either a Charmilles Technologies Robofil 200or an Eroda DCNC 300 electrical discharge wire machine(EDWM). These were used for 2D topographical, machinedsurface, microstructural and microhardness analysis. Sec-tions were hot mounted in Bakelite in a Buehler Metaservmounting press, ground using SiC paper and polished withSiO2 solution. After polishing, they were etched for 7 s in

Table 2Test matrix

Depth of cut (DOC, mm) Cutting speed (m/min)25 40

0.05 0.1 0.05 0.1

Cutting environment (barl/min) 6526 206

Krolls reagent (composition 3% hydrogen fluoride, 6% ni-tric acid and balance water). Knoop microhardness mea-surements were carried out on a Mitutoyo MVK-G3 micro-hardness measuring machine. Microhardness measurementswere conducted with a load of 25 g for 15 s. One of theproblems associated with microhardness measurement con-cerns its sensitivity to hard particles just below the surface.For this reason, a series of three readings were taken at eachdepth from the machined surface and an average obtained.The microhardness measurements in the bulk of the sampleand at 15m depth, were taken five times to improve theaccuracy of the results. Subsurface microstructural analysiswas conducted using a Leitz Wetzler optical microscope.Selected surfaces were examined in a JOEL 5410 scanningelectron microscope (SEM). 2D surface roughness measure-ments (Ra and Rt) were conducted on a Rank Taylor HobsonForm 10 surface profilometer with a cut-off and evaluationlength of 0.8 and 5 mm, respectively. Each measurement wasrepeated five times at a different position on the surface andan average taken.

Cutting force was measured with a Kistler 9257A threecomponent piezoelectric dynamometer and associated 5011charge amplifiers connected to a PC employing KistlerDynoware force measurement software. Measurements weretaken within the first 10 s of cut with a new tool (


Fig. 1. Lamellae deformed in the direction of cutting.

Fig. 2. Cracked TiB2 particles.

cracks which ran in an arc from the machined surface intothe main body of the workpiece and back out (see Figs. 3and 4). Both Zhang et al. [29] and Mantle and coworkers[2022] have reported similar results.

Examination of the workpiece surface showed that thecracks were perpendicular to the feed direction and werecontained within a single feed band (0.05 mm). When com-paring the surfaces produced with those obtained in previ-ous work at higher feed rates and depths of cut [2022], itis clear that a reduction in crack size (from 150 to 50m)and depth (from 15 to 5m) was obtained. Fig. 5 shows thatthe crack density was greater at 25 m/min cutting speed andincreased as the depth of cut increased. ANOVA showed thatdepth of cut had the largest percentage contribution ratio(PCR) and was the only factor to be statistically significant

Fig. 3. Sectional view of crack produced in the workpiece.

Fig. 4. Workpiece surface cracking.

Fig. 5. Crack density with operating parameters.

at the 5% level, see Table 3. The PCR gives the total per-centage that each factor contributes to the total variation inthe results and is a measure of how much the performancecould be improved if the factor was controlled exactly. Thefigure of 5% (chosen by convention) denotes the level of riskin incorrectly assuming that a factor is significant when it isnot. Cutting fluid supply did not appear to affect the levelof cracking to any great extent. The lowest level of crackingwas seen at 40 m/min cutting speed and 0.05 mm depth ofcut.

The form of surface cracking observed in the present workwas described by Bailey and Sadat who correlated its in-cidence to the formation of discontinuous chips for a wide

Table 3ANOVA results (5% level).Factor PCR (%)

Numberof cracks

Cuttingforce

Surfaceroughness(Rt)

Surfaceroughness(Ra)

Depth of cut 53 67.6 20 0Cutting speed 19 2.4 0 0Cutting fluid supply 0 0 67 84Error 28 30 13 16


Fig. 6. Elastic stress pattern existing at the start of a cut [37].

range of workpiece materials [3036]. Bailey [3032] pro-posed that crack nucleation in the primary shear zone, oc-curs when either the strain, strain rate and/or temperature,are such that the ductility of the chip material in that area isexceeded. Primary cracks are produced in the region of thetool nose and propagate down into the surface following ashallow angle under a combination of high shear and ten-sile deformation. The crack then rotates to follow the pathof maximum stress beneath the surface as shown in Fig. 6[31]. If propagation does not completely reach the surface,a crack or cavity is left behind. In addition, the material be-tween the workpiece surface and the upper surface of theprimary crack is subjected to bending because of the con-tinual advance of the tool. Therefore, a second crack initi-ates and propagates towards the workpiece surface forminga cavity. When examining quick stop chip/workpiece surfacesections, cracks seen in the discontinuous chip could be di-rectly matched to cracks left behind in the workpiece surface[33]. This model fits the surface damage seen with -TiAl.

Fig. 7. Cutting forces.

The larger cracks and the lack of plastic deformation seenin the cavity are indicative of the low ductility of the mate-rial. Bailey [32] also noted that when cutting speed was in-creased, chip formation with AISI 4340 steel changed fromdiscontinuous to continuous and surface cracking was pre-vented. Higher cutting speeds generate higher cutting tem-peratures and so increased the ductility of the chip materialpreventing crack initiation in the shear zone. To obtain asimilar effect with -TiAl, the temperature generated wouldhave to exceed the brittle to ductile transition temperature(BDTT) of around 700C. Attempts to generate high temper-atures in the shear zone and so increase the ductility of thechip material have been conducted previously [20]. Whenusing polycrystalline cubic boron nitride (PCBN) tools athigh cutting speeds (200 m/min), analysis of surfaces pro-duced showed less cracking compared to when using un-coated WC, however, the cracks were large 0.20.6 mmlong. These were believed to have been created by relax-ation of thermal stresses in the surface layers upon cooling,caused in part by the materials poor thermal conductivity(26 W/mK), giving rise to high thermal gradients.

Material removal when machining ceramics occurs whenstresses build up ahead of the advancing tool around grainboundary interactions, dislocation pile ups, inclusions andother stress raisers. These stresses initiate cracks which growrapidly causing fracture and chip formation. Consequently,cracks and cavities are left in the machined surface. Duc-tile regime machining of ceramic materials has been investi-gated by a number of authors [3841]. It has been suggestedthat there is a fracture threshold loading limit below whichcracks will not develop and the material will be removedby plastic flow [38]. Both the theory proposed by Bailey[3032] for ductile materials and that for ceramics put for-ward by Zhang et al. [38], state that to prevent crackingthe stresses developed in the cutting zone must be reduced.Crack free surfaces have been produced in -TiAl by grind-ing [23] and HSM [24]. A characteristic of both processes isthe low cutting forces generated. Cutting force can be used


as a measure of the stress developed in the cutting zone dur-ing cutting. Fig. 7 shows that in the present work, depth ofcut had the largest effect on all three cutting forces with aPCR of 67.6% (Table 3). Depth of cut also had the largesteffect on the number of cracks produced with an averagereduction from 21 to 13.25 cracks/mm2 when the depth ofcut was reduced to 0.05 mm, see Fig. 5. Varying the cuttingspeed (2.4% PCR) and cutting fluid supply (0% PCR) withinthe range investigated had no major effect on either cuttingforce or the number of cracks produced. In addition, thetrend for cutting speed having a greater (albeit insignificant)effect than cutting fluid supply follows for both the level ofcracking and the force generated, see Table 3. An increase incutting speed for the majority of workpiece materials wouldlead to a reduction in cutting force due to a combination offactors, including an increase in the shear plane angle and acorresponding reduction in chip thickness/contact area [42],together with an increase in cutting temperature and thus areduction in shear strength of the workpiece material in theshear zone [37]. However, as the BDTT for -TiAl is around700C and the chips produced remained needle like, it isunlikely that the present operating parameters caused tem-peratures to reach this level.

The use of high pressure cutting fluid supply has beenshown to reduce cutting force by preventing seizure of chipson the rake face, reducing friction and altering the chipcontact length and shear plane angle [43]. This is becausethe cutting fluid partially penetrates the tool/chip interface.However, when turning -TiAl the chip contact area is verysmall and any further reduction is unlikely to significantlyinfluence cutting force. Trent [44] states that the tangentialcutting force is usually the largest of the three force compo-nents, followed by feed and radial values. Fig. 7 shows thatthe radial force was larger than the feed force. The low valueof feed force, was a function of the very small chip contactarea and limited chip flow across the rake face seen withthis material. In addition, the high effective approach anglecaused by cutting solely on the nose radius would contribute

Fig. 9. Workpiece surface roughness results (a) 65 bar26 l/min and (b) 20 bar6 l/min.

Fig. 8. Cutting force and tool wear against time.

to the higher than anticipated radial forces. Fig. 8 showsthat as wear of the cutting tool increased cutting force roserapidly. Likewise, it has also been shown that as tool wearincreases the level of cracking increases.

The use of high pressure cutting fluid supply resulted inincreased workpiece surface roughness (from 0.2 to 0.8mRa), as detailed in Fig. 9. Statistical analysis of both Rt andRa (see Table 3), showed that cutting fluid supply had a largeinfluence on the results (67 and 84% PCR, respectively).High pressure cutting fluid supply has been shown to causehigh levels of built up layer (BUL) on the tool [45]. Randomfracture of the BUL, results in an uneven workpiece surfaceand BUL from the tool is often left adhering to the surface,this can clearly be seen when comparing Fig. 10a and b.

Analysis of the microhardness depth profile showednear surface (15m) hardness values of around doublethe bulk value. Fig. 11 shows that the microhardness in-crease was confined to a depth of around 200250m from


Fig. 10. SEM micrographs of the workpiece surface produced at 25 m/mincutting speed, 0.05 mm depth of cut and both cutting fluid environments.

the machined surface. Examination of the peak hardnessvalues showed that the application of high pressure cut-ting fluid supply resulted in increased peak microhardness(610660 HK0.025), possibly due to the increased coolingeffect reducing cutting temperatures and thus the ductilityof the workpiece in the shear zone.

Fig. 11. Microhardness depth profile.

4. Conclusions

Under all the conditions used in these trials, the workpiecesurface damage consisted of deformed lamellae, crackedTiB2 particles, a highly strain hardened surface layer withcracks running into the body of the workpiece and backout towards the machined surface.

When using very small depths of cut and low feed rates,a reduction in the size of the cracks was obtained com-pared to previous work. In addition the number of cracksdiminished when the depth of cut was reduced from 0.1 to0.05 mm and the cutting speed was increased to 40 m/min.The type of cutting fluid supply had no effect upon sur-face cracking.

Microhardness readings at the workpiece surface wereslightly lower than previously found, however, they werestill approaching twice the bulk hardness. The use of highpressure cutting fluid supply led to a slight increase insurface microhardness, most probably because the mate-rial was subjected to lower cutting temperatures. No othertrend for microhardness was observed with respect to cut-ting conditions.

Depth of cut had most effect on all three cutting forceswith a PCR of 67.6%. Reducing the depth of cut from0.1 to 0.05 mm reduced cutting forces significantly.Cutting speed and cutting fluid environment had nosignificant effect on cutting force within the range inves-tigated.

In general, the cutting forces were low, the maximumforce being 126 N (tangential), at 40 m/min cutting speed,0.1 mm depth of cut and with the high pressure cuttingfluid supply.

For most workpiece materials, the tangential force isusually the largest of the three force components, fol-lowed by the feed and radial values. In the presentwork, the radial force was larger than the feed force.


The low value of the feed force reflected the very smallchip contact area and limited chip flow across the rakeface.

Acknowledgements

The authors would like to thank Prof. A.A. Ball, Head ofthe School of Manufacturing and Mechanical Engineeringand Prof. M.H. Loretto, Director of the IRC in Materialsfor High Performance Applications, for provision of facili-ties and funding. Thanks also go to the UK Engineering andPhysical Sciences Research Council (EPSRC), RollsRoycePlc (Wayne Voice and Colin Sage), De Beers Industrial Di-amonds (UK) Ltd. (Matthew Cook and John Collins), andSandvik Coromant UK (Andy Smith) for funding and tech-nical support.

References

[1] A. Partridge, M.R. Winstone, -TiAl alloys: current status and futurepotential, in: Proceedings of the 4th International Charles ParsonsTurbine Conference on Advances in Turbine Material, Design andManufacture, Newcastle, UK, 46 November 1997, pp. 358377.

[2] D.M. Dimiduk, Gamma titanium aluminide alloys an assessmentwithin the competition of aerospace structural materials, Mater. Sci.Eng. A263 (1999) 281288.

[3] E.F. Bradley, Superalloys: A Technical Guide, ASM International,1988, ISBN 0-87170-327.

[4] J. Horrocks, RollsRoyce, Derby, personal communication, 44(1992).

[5] K.S. Chan, Understanding fracture toughness in gamma TiAl, J. Met.(1997) 3038.

[6] R. Gnanamoorthy, Y. Mutoh, Y. Mizuhara, Fatigue crack growthbehaviour of equiaxed, duplex and lamellar microstructure -basetitanium aluminides, Intermetallics 4 (1996) 525532.

[7] K.S. Chan, The fatigue resistance of TiAl base alloys, J. Met. 49(1997) 5358.

[8] K.S. Kumar, J.A.S. Green, D.E. Larsen, L.D. Kramer, XD titaniumaluminde composites, Adv. Mater. Proc. 4 (1995) 3538.

[9] P. Bartolotta, J. Barrett, T. Kelly, R. Smashey, The use of castTi48Al2Cr2Nb in jet engines, J. Met. 49 (1997) 4851.

[10] C.M. Austin, T.J. Kelly, K.G. McAllister, J.C. Chesnutt, Aircraftengine applications for gamma titanium aluminide, in: M.V. Nathal,R. Darolia, C.T. Liu, P.L. Martin, D.B. Micacla, R. Wagner, M.Yamaguchi, (Eds.), TMS, Warrendale, PA, Struct. Intermetallics,1997, pp. 413425, ISBN 0-87339-375-9.

[11] Anon, www.rallycars.com/Cars/Mitsubishi Evo V/EvoV 6.html,2000.

[12] A.R. Machado, J. Wallbank, Machining of titanium and its alloys a review, Proc. Inst. Mech. Eng. 204 (1990) 5359.

[13] M. Wang, Y. Zhang, Diffusion wear in milling titanium alloys, Mater.Sci. Technol. 4 (1988) 548553.

[14] E.O. Ezugwu, Z.M. Wang, Titanium alloys and their machinability a review, J. Mater. Proc. Technol. 68 (1997) 262274.

[15] P.A. Dearnley, A.N. Grearson, Evaluation of principle wearmechanisms of cemented carbides and ceramics used for machiningtitanium alloy IMI 318, Mater. Sci. Technol. 2 (1986) 4758.

[16] R. Komanduri, W.R. Reed, Evaluation of carbide grades and a newcutting geometry for machining titanium alloys, Wear 92 (1983)113123.

[17] A. Jawiad, C.H. Che-Haron, A. Fallah, Tool wear in machiningof titanium alloy Ti 6242, in: Proceedings of the 3rd InternationalConference on Progress of Cutting and Grinding, Osaka, Japan,November 1996, pp. 126131.

[18] A.L. Mantle, D.K. Aspinwall, M.L.H. Wise, Single point turningof gamma titanium aluminide intermetallic, Titanium 95, in:Proceedings of the Eighth World Conference on Titanium, Vol. 1,Birmingham, UK, 1995, pp. 248255, ISBN 1-86125-005-3.

[19] A. L. Mantle, D.K. Aspinwall, Temperature measurement and toolwear when turning gamma titanium aluminide intermetallic, in:Proceedings of IMC-13, Limerick, Ireland, 46 September 1996,pp. 427436.

[20] A.L. Mantle, D.K. Aspinwall, Tool wear and surface integrity whenturning gamma titanium aluminide intermetallics, in: Proceedings ofthe 3rd International Conference on Progress of Cutting and Grinding,Osaka, Japan, November 1996, pp. 6368.

[21] A.L. Mantle, D.K. Aspinwall, Surface integrity and fatigue life ofturned gamma titanium aluminide, J. Mater. Proc. Technol. 72 (1997)413420.

[22] A.L. Mantle, The machining of gamma titanium aluminideintermetallics, Ph.D. Thesis, University of Birmingham, September1998.

[23] S.A. Bentley, N.P. Goh, D.K. Aspinwall, Reciprocating surfacegrinding of a gamma titanium aluminide intermetallic alloy, in:Proceedings of IMC 16/AMPT 99, Dublin, Ireland, 36 August 1999,pp. 427435.

[24] A.L. Mantle, D.K. Aspinwall, Surface integrity of a high speed milledgamma titanium aluminide, in: Proceedings of IMC 16/AMPT 99,Dublin, Ireland, 36 August 1999, pp. 447456.

[25] S.A. Bentley, A.L. Mantle, D.K. Aspinwall, The effect of machiningon the fatigue strength of a gamma titanium aluminide intermetallicalloy, Intermetallics 7 (1999) 967969.

[26] A.L. Mantle, S.A. Bentley, D.K. Aspinwall, A comparison oftool wear when turning Ti45Al2Nb2Mn + 0.8 vol.% TiB2 andTi44Al8Nb1B titanium aluminide intermetallics, in: Proceedingof the 4th International Conference on the Behaviour of Materialsin Machining, Stratford, UK, 1213 November 1998, pp. 300308.

[27] J.F. Kahles, M. Field, D. Eylon, F.H. Froes, Machining of titaniumalloys, J. Met. 37 (1985) 2733.

[28] S. Jeelani, K. Ramakrishnan, Subsurface plastic deformation inmachining 6Al2Sn4Zr2Mo titanium alloy, Wear 85 (1983)121130.

[29] H. Zhang, M.L.H. Wise, D.K. Aspinwall, E.M. Aspinwall,Investigation of machinability of a as HIPped gamma titaniumaluminide, in: D.A. Stephenson, R. Stevenson (Ed.), Material Issuesin Machine II, 26 October 1994, pp. 89104, ISBN 0-87339-242-6.

[30] J.A. Bailey, Surface damage during machining of annealed 18%nickel maraging steel Part 1. Unlubricated conditions, Wear 42(1977) 277296.

[31] J.A. Bailey, Surface damage during machining of annealed 18%nickel maraging steel Part 2. Lubricated conditions, Wear 42(1977) 297303.

[32] J.A. Bailey, On surface damage during machining of AISI 4340 steel,Wear 27 (1974) 161173.

[33] A.B. Sadat, J.A. Bailey, Some observations on surface damage duringmachining of a bearing bronze, Wear 108 (1986) 255268.

[34] A.B. Sadat, Surface region damage of machined Inconel 718 nickelbase superalloy using natural and controlled contact length tools,Wear 119 (1987) 225235.

[35] J.A. Bailey, Surface damage during machining of solution treated andaged 18% nickel maraging steel Part 1. Unlubricated conditions,Wear 44 (1977) 361370.

[36] J.A. Bailey, Surface damage during machining of solution treatedand aged 18% nickel maraging steel Part 2. Lubricated conditions,Wear 44 (1977) 371376.

[37] M.C. Shaw, Metal Cutting Principles, Clarendon Press, Oxford, 1997,ISBN 0-19-859020-2.


[38] G.M. Zhang, D.K. Anand, S. Ghosh, W.F. Ko, Study of theformation of macro and micro cracks during machining of ceramics,in: Proceedings of the International Conference on Machining ofAdvanced Materials, MD, USA, July 1993, pp. 465478, NISTspecial publication 847.

[39] P.N. Blake, R.O. Scattergoood, Ductile regime machining ofgermanium and silicon, J. Am. Soc. 73 (4) (1990) 949957.

[40] T. Inamura, N. Takezawa, N. Nakahara, Brittle/ductile transitionphenomena observed in computer simulations of machining defectfree monocrystalline silicon, Ann. CIRP 46 (1) (1997) 3134.

[41] P.A. Beltrao, A.E. Gee, J. Corbett, R.W. Whatmore, Ductile modemachining of commercial PZT ceramics, Ann. CIRP 48 (1) (1999)437440.

[42] P.L.B. Oxley, Mechanics of Machining, Ellis Horwood, Chichester,UK, 1989, ISBN 0-7458-0007-6.

[43] A. Smith, S. Lagerberg, P. Dahlman, J. Kaminski, High pressurejet assisted turning, Industrial Tooling 99, Southampton, UK, 78September 1999, pp. 6271.

[44] E.M. Trent, Metal Cutting, 3rd Edition, Butterworth-Heinemann,London, 1991, ISBN 0-7506-1068-9.

[45] A.R.C. Sharman, R.C. Dewes, D.K. Aspinwall, P. Bowen, Tool wearwhen turning gamma titanium aluminide using carbide and PCDtools with reduced depths of cut and high pressure cutting fluid, in:Proceedings of the 28th North American Manufacturing ResearchConference, University of Kentucky, Lexington, Kentucky. 2000,pp. 161166, ISBN 0-87263-518-X.

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