Cutting tool wear in the machining of hardened steels. part i alumina.ti c cutting tool wear

Wear 247 (2001) 139–151

Cutting tool wear in the machining of hardened steelsPart I: alumina/TiC cutting tool wear

J. Barry∗, G. ByrneDepartment of Mechanical Engineering, University College Dublin, Belfield, Dublin 4, Ireland

Received 27 July 2000; accepted 31 August 2000

Abstract

A study was undertaken to investigate the mechanisms of alumina/TiC cutting tool wear in the finish turning of hardened steels withparticular cognisance of the work material inclusion content. A six-fold variation in tool life was observed when machining different heatsof BS 817M40 steel (similar to AISI 4340) of 52 HRC. In machining steels containing Ca-bearing mixed oxide inclusions, a reactionbetween the alumina phase of the tool and oxide inclusionary deposits is the dominant wear mechanism. In machining steels containingvery low levels of Ca or steels with a very low inclusion content, tool wear appears to be largely based on superficial plastic deformation ofthe tool surface. The rate of tool wear appears to be determined by the hard inclusion content or alloy carbide content of the work material.Impingement of hard particles against the tool surface are thought to result in the generation of transient localised stresses which exceedthe average contact pressures and, thus, either facilitate the operation of additional slip systems or overcome the increases in the criticalresolved shear stresses on active slip systems due to prior strain. The influence of saw-tooth chip formation on cutting tool wear is alsoconsidered. © 2001 Elsevier Science B.V. All rights reserved.

Keywords:Tool wear; Machining; Alumina

1. Introduction

The machining of hardened steels of 45–65 HRC, withgeometrically defined cutting tools is a process of consider-able industrial importance. ‘Hard machining’ offers certainadvantages over the more traditionally employed grindingprocesses [1,2], however, there is considerable concern re-garding the metallurgical structure of machined surfaces andin particular the surface ‘white layer’ [3,4]. Even with anunworn cutting tool, microstructural alterations are evidentas a partially transformed nanocrystalline layer, less thanone micron in depth from the geometrical surface [5]. As aflank wear land develops, the thickness of the microstruc-turally altered layer increases, typically to several micronsat the end of the tool life. Thus, minimising the rate of cut-ting tool wear is a key objective in the manufacture of highprecision hardened steel components.

Among the earliest reports of the machining of hardenedsteels with geometrically defined cutting tools is that of Kingand Wheildon [6] in which steels of up to 61 HRC were ma-chined with alumina cutting tools. However, it was not until

∗ Corresponding author. Tel.:+353-1-7-06-19-83;fax: +353-1-2-83-05-34.E-mail address:[email protected] (J. Barry).

after the development of TiC reinforced alumina and in par-ticular, CBN and CBN/TiC composites, that hard machiningbecame an established process. These materials remain the‘enabling technology’, however, the mechanisms throughwhich they wear are not adequately understood. This, thefirst of two papers on tool wear in hard machining is con-cerned with the wear of alumina/TiC cutting tools. The wearmechanisms of CBN/TiC cutting tools are discussed in partII [7].

Several mechanisms have been highlighted as being ofimportance in the wear of alumina and alumina-compositecutting tools, among the most important of which are‘superficial’ plastic deformation [6,8–12] and chemical in-teractions between the tool and work materials [13–17]. Themajority of the studies of the wear mechanisms of aluminaand alumina composite cutting tools have been undertakenusing steels of low to moderate hardness, e.g. 150–300BHN. Qualitatively, the findings are considered relevant tothe machining of hardened steels in that the primary dif-ference between hard machining and ‘soft machining’ isgenerally, though not necessarily, the evolution of highercutting temperatures, e.g. 200◦C [18,19].

It is well known that single crystal alumina (a-Al2O3)can exhibit considerable ductility at elevated temperaturesand even at relatively low temperatures under hydrostatic

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140 J. Barry, G. Byrne / Wear 247 (2001) 139–151

stress states [20]. The temperatures and pressures at whichthe basal and prism slip systems and possibly the pyrami-dal slip system are operative are easily attained during thecutting of steels, in which for example, temperatures of upto 1400◦C have been measured [21] and stresses greaterthan 1 GPa [5]. Indeed, transmission electron micrographsof dislocation structures (generated during plastic deforma-tion) within bulk alumina and alumina coatings have beenpresented by Kim et al. [8] and Larsson et al. [9], respec-tively. Plastic deformation is usually described as superficialor discrete in nature as it is limited to the outermostmm orso of the tool surface [11,12]. An exception to this has beenpresented by Goh et al. [11], who describe a transition inthe flank wear mechanism of alumina tools in which bulkplastic deformation proceeds superficial deformation.

Superficial plastic deformation usually results in the for-mation of narrow ridges (of severalmm in width) on thecrater wear surface, running parallel to the chip flow direc-tion, as a result of either necking of surface asperities oradherence and smearing of wear debris generated upstream[9,11,12]. Brandt and Mikus [10] have suggested that workmaterial alumina inclusions may adhere to the tool surface,forming protective layers and reducing the rate of cuttingtool wear. The work material inclusionary deposits formedteardrop-shaped ridges, tapered in the direction of chip flowand between which ‘rivers’ of Mn-S-Fe-O compounds wereobserved to flow. The formation of course ridges on theflank wear land of alumina composite tools was attributedby Brandt [12] to intergranular fracture due to dislocationpile-up following plastic deformation.

Brandt and Mikus [14] noted that surface deformationof the cutting tool may be greatly affected by chemical re-actions between the work and tool material. Examinationof alumina cutting tools revealed increased concentrationsof Fe and Mg. It was noted that the spinel phases formedbetween alumina and the oxides of these elements possessgreater ductility and higher yield strengths and that suchphenomenon may account for the observed reduction in therate of tool wear. Tönshoff and Bartsch [13] also presentedevidence which shows the reaction between alumina andFe2O3 to occur during machining. Upon subsequent coolingto room temperature, the high temperature phase decom-poses resulting in the formation of ultra-fine alumina grainsand an amorphous iron oxide.

Table 1Composition of steels used for machining tests, in wt.%a

Steel Heat C Si S Cr Ni Mo Al Ti O Ca

BS 817M40 1 0.40 0.163 0.034 1.176 1.371 0.283 0.029 0.003 30 252 0.38 0.193 0.021 1.104 1.350 0.254 0.035 0.003 50 33 0.44 0.204 0.012 1.163 1.653 0.189 0.004 0.006 20 64 0.44 0.165 0.030 1.220 1.546 0.247 0.026 0.003 – 195 0.41 0.166 0.031 1.214 1.510 0.246 0.023 0.003 – 25

Tool steel – 0.797 0.199 0.007 1.71 0.144 0.416 0.026 0.004 20 24

a Only one heat of the low alloy tool steel was used. The O and Ca contents are in parts per million.

Another predominantly chemical wear mechanism is thatoriginally proposed by Ham and Narutaki [15] in which areaction between CaO work material inclusions and the alu-mina tool result in the formation of soft glassy compounds,thus, accelerating tool wear. In contrast, in the machining ofCa-treated steels with P-type carbide tooling, a reduction inthe rate of tool wear is effected by the formation of an ad-herent layer or ‘belag’ which limits diffusion/dissolution ofthe tool [17,22]. Narutaki [16] and Narutaki and Murakoshi[17] have also presented evidence that pure alumina toolswear more rapidly in the machining of Ca-treated steels com-pared to the machining of conventional steels. However, itwas also noted by these authors that an increase in the sul-phur content of the Ca-treated steel effectively negated thedeleterious effects of Ca, through, it was suggested, the for-mation of a Mn-rich layer which inhibited the reaction be-tween CaO and the (alumina) tool. Yamane et al. [23] havenoted that oxides other than CaO may react with alumina,examples being MnO and SiO2; it was noted that the low-est melting point in the ternary system composed of theseoxides and alumina is only 1200◦C.

Diffusion/dissolution wear of the TiC phase in alu-mina/TiC composites has also been suggested to be animportant wear mechanism [12]; the solubility of TiC iniron being three orders of magnitude greater than that ofalumina. Katsumura et al. [24] have shown that while highcontent TiC tools exhibit greater flank wear resistance thanlow content TiC tools, they are subject to greater rates ofcrater wear. This was attributed to the fact that TiC was‘more apt’ to react with the work material, resulting in agreater degree of adhesive wear.

The aim of the current study was to substantiate and/orto clarify the role of the above wear mechanisms in themachining of hardened steels with particular cognisance ofthe work material inclusion content.

2. Experimental

Finish turning and facing tests were undertaken using fiveheats of BS 817M40 steel and a low alloy tool steel. Thecompositions of these steels are given in Table 1. The geom-etry of the workpieces used for both cutting arrangements isshown in Fig. 1. The workpieces were austenitised at 850◦C

J. Barry, G. Byrne / Wear 247 (2001) 139–151 141

Fig. 1. The geometry of the workpiece used for facing and turning tests.

for 1 h, oil quenched and tempered for 2 h. For the majorityof tests with the BS 817M40 steels, the workpieces weretempered at 250◦C, resulting in a hardness of 52 HRC. Testswere also conducted with workpieces of 40 HRC (480◦Ctemper) and 49 HRC (320◦C temper). Tests with the lowalloy tool steel were conducted with workpieces of 52 HRC(450◦C temper) and 60 HRC (180◦C temper).

The BS 817M40 steels had a lath martensite structure withvery fine cementite precipitates. The structure of the lowalloy tool steel was a fine lath martensite with a small degreeof twinning [5]. This steel also contained a dispersion ofmicron-sized alloy carbides. The inclusion content of eachsteel is discussed in Section 3.1.

The alumina/TiC tools employed in the study contained71% a-Al2O3, 28% TiC, 1% MgO and had a grain size of1–2mm. All cutting inserts were of ISO (1832:1991) desig-nation; TNGA 160412 T01020. The nose radius of all toolswas rε = 1.2 mm, the clearance angle,α = 6◦, the edgeradius,rβ = 5–10mm and the (effective) rake angle,γ =−26◦ (cutting took place exclusively on the edge chamfer,of width, bγn = 100mm).

The tool wear tests were performed using parameters rep-resentative of those used for finish turning; cutting speed,vc = 2.5 and 4.17 m/s, feed,f = 100mm and depth ofcut, ap = 200mm. All tests were undertaken on a Dae-woo PUMA 4-3A CNC lathe under constant surface speedcontrol and were replicated six times. Heat 3 was also ma-chined on a Colchester Mastiff lathe; the results presentedhere are from the tests performed on the PUMA lathe. Thetool forces were measured with a three component platformdynamometer. Flank wear and crater wear parameters weremeasured using a toolmakers microscope and a stylus pro-filometer, respectively. The wear surfaces were examinedunder a scanning electron microscope (SEM) equipped withan energy dispersive X-ray spectrometer (EDS).

Quantification of work material inclusion size distributionwas undertaken using optical microscopy on unetched pol-ished surfaces of each steel. Sections were taken normal andparallel to the rolling direction of the bars. At least two setsof 20 adjacent, non-overlapping fields were digitally cap-tured and analysed. Inclusions in each field were identifiedand the average cross-sectional area and diameter calculated.

3. Results

3.1. Inclusion content of steels

Histograms of inclusion size for the low alloy tool steeland heats 1, 2 and 3 of BS 817M40 steel are given in Fig. 2.A qualitative assessment of the composition of the inclu-sions in each steel was undertaken using an SEM equippedwith EDS. All heats of BS 817M40 steel and the low al-loy tool steel were found to contain very fine MnS inclu-sions. These inclusions appeared elongated when viewedin cross-sections taken parallel to the rolling direction. Allheats of BS 817M40 steel also contained multi-phase or du-plex inclusions, e.g. containing phases other than MnS. Gen-erally, these inclusions were of the non-deformable type anddid not appear elongated in the rolling direction. Most of themulti-phase inclusions in heats 1, 4 and 5 of BS 817M40steel were composed of a mixed oxide phase (Ca, Al)O en-cased in a sulphide phase. Similar inclusions were observedin heat 4, though were considerably fewer in number; notethe Ca contents of the different heats of BS 817M40. (PureCaO inclusions were not observed in any heat of steel).Fig. 2b–d show backscattered electron images of multi-phaseinclusions from the different heats of BS 817M40 steel. Ofthe relatively few duplex inclusions observed in heat 2, most

Fig. 2. (a) Inclusion size distributions for heats 1, 2 and 3 of the BS817M40 steel and the low alloy tool steel. Duplex inclusion in heat 1(c), heat 3 (d) and heat 4 (b).


Fig. 3. Volume wear rate of the flank per unit engagement length as a function of tool path for alumina/TiC tools in machining different heats of BS817M40 steel and a low alloy tool steel, all of 52 HRC with cutting speed,vc = 2.5 m/s.

contained traces of aluminium and oxygen (note the rela-tively high Al and O content of this steel, Table 1). On ex-amination of a specimen of the low alloy tool steel, of areaA = 10 mm2 under SEM at 5000 X, only one multi-phaseinclusion was observed and was of similar composition andmorphology to that shown in Fig. 2c.

3.2. Cutting tool wear rates

Fig. 3 shows the volume wear of the tool flank per unitengagement length,Vα (on the left abscissa) and flank wearland width, VBC, (on the right abscissa) as a function of toolpath for four heats of BS 817M40 steel and the low alloytool steel, all of 52 HRC, for cutting speed,vc = 2.5 m/s(f = 100mm, ap = 200mm). The wear rate (1Vα/1lc)of the tools used to machine heat 5 of BS 817M40 steelwas intermediate to that of the tools used to machine heats1 and 4. The flank wear rate of the tools used to machinethe low alloy tool steel of 60 HRC was very similar to thatof the tools used to machine the low alloy tool steel of 52HRC (0.48mm2/m compared to 0.46mm2/m, respectively).It may be noted that the tools used to machine heat 3 of BS817M40 steel exhibited the lowest wear rate, at least onequarter that of the tools used to machine the other heats ofthis steel. The wear rate of the tools used to machine the lowalloy tool steel was intermediate of that of the tools used tomachine heat 3 and 1 of BS 817M40 steel.

With regard to the crater wear rate, several issues arenoted. Firstly and as is discussed below, the morphology ofthe tool crater surfaces after machining heats 1, 4 and 5 ofBS 817M40 steel was quite complex, being composed of aseries of ridges and grooves running approximately parallelto the chip flow direction. Determination of the standard

crater wear parameter, KT (maximum crater depth) was notpossible as the groove width was of similar dimensions asthe stylus tip radius (5mm). Also, the included angle of thestylus was 90

◦and as such, with a deepening of the grooves

(due to wear), the stylus was found to traverse along thepeaks of neighbouring ridges rather than traversing alongthe bottom of the central groove. A further issue was that thestylus traverse direction was not necessarily parallel to thegroove/ridge direction such that frequently, the stylus wouldtraverse across a number of ridges and grooves. Nonetheless,it may be noted that for those heats of BS 817M40 steelfor which there was a rapid rate of flank wear, there was acorrespondingly rapid rate of crater wear.

3.3. Morphology of tool wear surfaces

Fig. 4 shows several micrographs of the tools used tomachine heat 1 of BS 817M40 steel with cutting speed,vc = 2.5 m/s. In Fig. 4a, ridges and grooves are clearlyevident on the crater and flank wear surfaces. The ridges onthe flank wear lands were noted to form only after the flankwear land width exceeded approximately 150mm, i.e. forVBC ≤ 150mm, the flank wear land was relatively smooth.In contrast, the ridge/groove structure on the crater wearsurface, Fig. 4b, was present over the entire tool life, thegrooves deepening with increased cutting time. Fig. 4c andd show a secondary electron image (SEI) and back scatteredelectron image (BEI) of the front of the crater wear surface,respectively. Noting that atomically heavier phases appearbrighter in back scattered electron imaging, it is evident thatthe TiC phase stands proud of the alumina matrix.

Fig. 4e and f show a SEI and a BEI, respectively, ofa region on the flank wear land of the tool, indicated in


Fig. 4. Scanning electron micrographs of the wear surfaces on an alumina/TiC tool used to machine heat 1 of BS 817M40 steel with cutting speed,vc = 2.5 m/s; after a tool path of,lc = 2.94 km.

Fig. 4a. The built up layers evident in the BEI, Fig. 4f canbe seen to run between the ridges, i.e. along the bottomof the grooves. The layers contain networks of fine crackswhich are thought to have initiated when the tool cooledto room temperature due to different coefficients of thermalexpansion. The EDS analysis of the built up layers revealthem to contain Mn, S, Ca, Al and O. Layers similar inappearance and composition to those on the tool flanks werefrequently observed on the crater wear surfaces. As on thetool flank, these layers were invariably deposited along thebottom of the grooves or on the sides of the ridges; layerswere never observed on the upper surfaces of the ridges. Thecomposition of the layers on both the flank and crater wearsurfaces was found to vary considerably, though invariablycontained those elements listed above.

Examination of the tools used to machine heat 1 of BS817M40 steel with cutting speed,vc = 4.17 m/s and heats4 and 5 with cutting speeds,vc = 2.5 and 4.17 m/s revealedalmost identical wear features as those observed on the toolsused to machine heat 1 with cutting speed,vc = 2.5 m/s;namely, the presence of ridges on the crater wear surface,the formation of ridges on the flank wear land once the flankwear land width exceeded approximately 150mm and thepresence of patchy built-up-layers containing Mn, S, Ca,Al and O. Fig. 5 shows the crater wear surface of a toolused to machine heat 1 with cutting speed,vc = 4.17 m/s.The ridge structures are somewhat more pronounced thatthose formed at the lower cutting speed,vc = 2.5 m/s. Notealso the left of Fig. 5a where the downstream end of aridge has fractured. In Fig. 5b, abrasive markings are clearly


Fig. 5. Scanning electron micrographs of the wear surfaces on an alu-mina/TiC tool used to machine heat 1 of BS 817M40 steel with cuttingspeed,vc = 4.17 m/s; after a tool path of,lc = 1.45 km.

evident and may be seen toflow to either side of the up-stream end of each ridge. Note the built up layer in the BEI inFig. 5c.

The morphology of the wear surfaces on the tools used tomachine the low alloy tool steel and heats 2 and 3 of the BS817M40 steel were significantly different to those shown inFigs. 4 and 5. Fig. 6 shows a number of micrographs andEDS spectra of tools used to machine heat 2 of BS 817M40steel. Unlike heats 1, 4 and 5, ridges did not form on the flankwear land, irrespective of its width, VBC, Fig. 6a. Similarly,the crater wear surface remained relatively smooth; comparefor example, Figs. 4b, 5a and 6b. Over the rear of the crater, afine ridge-type structure may be noted, Fig. 6c. On a numberof tools, a very thin layer of MnS was observed at the frontof the crater as shown in Fig. 6d. However, this is thoughtnot to be a stable built-up-layer as many tools did not exhibitsuch layers, Fig. 6f and g.

Fig. 7 shows a number of micrographs of a tool used tomachine heat 3 of BS 817M40 steel. The wear rate of the

tools used to machine this heat were at least one quarter thatof the tools used to machine the other heats of BS 817M40.Fig. 7a and b show the same cutting tool after a tool path of2.9 and 5.29 km, respectively. It is interesting to note that themost severe wear initially occurs over the rear of the crater.The wear surface is composed of some relatively smoothridges of material, running parallel to the chip flow direction(from the bottom to the top of each image), between whichthe wear surface is considerably rougher. The EDS analysisof the wear surfaces revealed traces of iron. Elements Mn,Ca and S were not present in detectable quantities (severalhundred parts per million (ppm)). In Fig. 7c, a small adheredfragment of steel may be seen, indicated by the arrow. Theflank wear surfaces were smoothly worn and as for the cratersurfaces, contained only traces of iron.

Fig. 8 shows micrographs of the tools used to machinethe low alloy tool steel of 52 HRC (8a–c) and of 60 HRC(8d). The rear of the crater surfaces are characterised by thepresence of narrow ridges, Fig. 8b, similar to those observedon the tools used to machine heat 2 of BS 817M40 steel.The influence of work material hardness on the rate of craterwear may be noted by comparing Fig. 8b and d; also note theinset profiles in the bottom right of each image. As for heats2 and 3 of BS 817M40 steel, the flank wear land was evenlyworn, Fig. 8c. The EDS analysis of the flank and cratersurfaces revealed only the presence of iron, (in addition tothe tool material elements).

4. Discussion

4.1. Influence of work material Ca content on alumina/TiCtool wear

It is considered significant that a near-identical wear sur-face morphology was observed on the tools used to ma-chine heats 1, 4 and 5 of BS 817M40 steel and that eachof the heats contain relatively high levels of Ca; 25, 19 and25 ppm, respectively, and also Ca-bearing mixed oxide in-clusions. The morphological wear features on these tools,which indicate that the dominant wear mechanism is basedon a reaction between Ca-bearing oxide inclusions and thealumina phase of the tool may be identified as:1. The presence of Ca-containing deposits on the wear sur-

faces.2. The manner in which the TiC phase stands proud of the

alumina matrix (Fig. 4c and d) indicating that the aluminaphase is subject to more rapid wear. As noted above, thesolubility of alumina in steel is three orders of magnitudeless than that of TiC. Also, according to the calculationsof Helle [25], the reactions between TiC and CaO arethermodynamic unfavourable.

3. (a) The formation of ridges over the rear of the craterwear surface and on the tool flank after a critical flankwear land is attained and (b) the fact that layers of workmaterial inclusionary compounds are only observed on


Fig. 6. Scanning electron micrographs and EDS spectra of the wear surfaces on alumina/TiC tools used to machine heat 2 of BS 817M40 steel withcutting speed,vc = 2.5 m/s; after a tool path,lc = 2.67 km. The micrographs in (f) and (g) are of crater wear surfaces on different tools, the locationcorresponds to that of (d).

the bottom of the grooves, between ridges, or on the sidesof the ridges.That the latter phenomena support the above reaction may

be appreciated by noting that the maximum temperature onthe rake face of the tool occurs a small distance back fromthe edge [21], Fig. 9 and that the temperature on the flankof the tool is likely to increase with an increase in flank landwidth (due to the greater level of energy dissipation at theflank/workpiece interface [5]). If it is assumed that depositedoxide inclusions react with the alumina phase of the tool,the reaction products will (thermally) soften as they traveldownstream along the flank and rake surfaces. Once thesecompounds soften sufficiently such that their flow stress isless than that of the steel chip, grooves will start to form.While the phenomenon which determines the initial locationof the grooves cannot be conclusively identified, it is clearthat once grooves of sufficient depth form, wear will largelyconcentrate there due to; (i) continued reaction between thereaction productsand the alumina phase of the tool and (ii)

the lower flow stress of the reaction products. The latter maybe likened togrouting in which a viscous plaster is smearedover a tiled area, filling the gaps between the individual tiles.

Goh et al. [11] and Brandt [12] observed ridges of similardimensions to those observed in the current study on thedownstream boundary of the flank wear lands. Goh et al.[11] attributed ridge formation to ‘stacking and bonding ofsmeared material’. Profile measurements of the flank wearland, Fig. 10a, revealed the lower ridged portion of theland to protrude into the workpiece. It was noted that suchbehaviour could only be accounted for by bulk plastic defor-mation of the tool. The ridges observed in the current studyalso protrude into the workpiece, however, their formationis not based on plastic deformation but rather, preferentialwear which accompanies the channelling of inclusionarydeposits/reaction products. The progression of flank wearaccording to this mechanism is illustrated in Fig. 10c. Theabsence of bulk plastic deformation has been confirmed byexamining the grinding marks (from tool manufacture) on


Fig. 7. Scanning electron micrographs of the crater wear surface on analumina/TiC tool used to machine heat 3 of BS 817M40 steel with cuttingspeed,vc = 2.5 m/s. (a) tool path,lc = 2.9 km, (b) and (c),lc = 5.29 km.

the unworn regions of the flank for evidence of distortion.The grinding marks are perpendicular to the direction ofsliding motion as illustrated in Fig. 10c.

Regarding the possibility of a chemical wear mechanism,Goh et al. [11] noted that EDS analysis of the wear surfacesdid not reveal significant quantities of Si, Mn, Ca or Fe whichmay have ‘diffused from the steel’ into the tool surface; theCa content of the work material was reported as ‘<0.01wt.%’, i.e. 100 ppm. However, it is noted that the possibilitythat different wear mechanisms are responsible for the ridgesobserved in the current study and in that of Goh et al. issupported by the fact that the progression of flank wear isdifferent in each study as illustrated in Fig. 10b.

Another point of interest in the study of Goh et al. [11] istheir suggestion that seizure occurs on the tool flank face asa result of the ridges protruding into the workpiece. In thecurrent study, seizure occurred neither on the flank face northe rake face. As the ridges on the flank may be considered to

protrude into the workpiece, so may the ridges on the crater.The latter has been confirmed by taking cross-sections of thechips, normal to the direction of chip flow. Fig. 10d showsone such section; the lower white band is the secondaryshear zone which forms on the tool side of the chip (see alsoFig. 12a). The spacing of the protuberances on the lowerside of the chip are in good agreement with the spacingbetween the ridges on the rake of the tool used to producedthe chip, Fig. 10e. The absence of seizure on the tool rakeface virtually assures its absence on the flank wear land.

In the study of Brandt [12], ridge formation on the flankwear land was attributed to intergranular fracture due to dis-location pileup. It was suggested that such fracture did notoccur on the rake face due to the high temperatures there.However, close examination of the micrographs of Brandtreveals the grooves to extend down the flank, between un-worn regions of the clearance face, similar to that depictedin Fig. 10c. The inclusion and Ca content of the workmaterial was not specified.

4.2. Plasticity based wear of alumina/TiC cutting tools

According to the chemical wear mechanism discussedabove, the relatively low wear rate of the tools used to ma-chine the low alloy tool steel, containing 24 ppm Ca, mightbe considered somewhat contradictory. However, noting thatthe inclusion content of this steel is less than one-sevenththat of heat 1 of the BS 817M40 steel and furthermore, thatof these inclusions, the vast majority are MnS, it followsthat the frequency of inclusionary deposits in the machin-ing of the low alloy tool steel is significantly less than thatin machining the high-Ca heats of BS 817M40 steel (theinclusion contents of heats 1, 4 and 5 are similar).

The very rapid wear rate of the tools used to machine heat2 of BS 817M40 steel which contains 3 ppm Ca, might alsobe considered contradictory to the findings regarding workmaterial Ca content and the corresponding tool wear rate.However, these findings pertain only to one specific wearmechanism. The morphology of the wear surfaces of thetools used to machine the low alloy tool steel and heats 2and 3 of BS 817M40 steel, being very different to the mor-phology of the tools used to machine heats 1, 4 and 5 ofBS 817M40 steel, suggest the operation of a different wearmechanism. In considering the nature of this wear mecha-nism (pertaining to the tool steel and heats 2 and 3 of BS817M40 steel), the variation in the morphology of the wearsurface at the front and rear of the crater is thought to beinsightful.

It is widely accepted that the temperature and stress distri-butions on the rake face of cutting tools are non-uniform, butare similar to those shown in Fig. 9 [21,26,27]. The decreasein the stresses and temperature over the rear of the tool facewill result in the activation of fewer slip systems within thealumina (and possibly TiC) grains constituting the tool. Assuch, von Mises criteria for homogenous plastic deforma-


Fig. 8. Scanning electron micrographs of the wear surfaces of alumina/TiC tools used to machine the low alloy tool steel of 52 HRC after a tool pathof, lc = 1.8 km, (a–c) and of 60 HRC after a tool path of,lc = 1.3 km (d). Cutting speed,vc = 2.5 m/s. The traces in the lower right of (b) and (d) areprofiles of the crater at the location of its maximum depth, vertical:horizontal magnification, 2.8 X.

tion of polycrystals (i.e. the operation of five independentslip systems) is less likely to be fulfilled with the result thata greater frequency (and degree) of strain incompatibilityoccurs between neighbouring grains. Strain incompatibilitymay result in decohesion or intergranular or transgranularfracture. That the stresses and temperatures over the frontof the crater facilitate homogenous plastic deformation can-not be ascertained, however, it is clear that the likelihood ofsuch behaviour is much greater there than over the rear ofthe crater.

If plastic deformation alone were to account for the verysignificant difference in the wear rates of the tools used tomachine heats 2 and 3 of BS 817M40 steel, a correspondingdifference in the resultant cutting force (which is related tothe integral of the stresses acting on the tool faces), should beobserved; this was not the case. Rather, the difference in thewear rates of the cutting tools used to machine the differentheats may be qualitatively accounted for by considering theinfluence of hard work material inclusions.

The impingement of hard inclusions against the tool faceswill result in the generation of transient localised stresseswhich exceed the average stress at that point and, hence, theresulting superficial plastic strain. Therefore, should a toolmaterial exhibit ductility at the temperatures and stressesgenerated during cutting, its wear rate may be related tothe hard inclusion content of the work material, through amechanism other than abrasion.

In this regard, it is noted that the Al content of heats 2and 3 of BS 817M40 steel are 350 and 40 ppm, respectively,and the O contents, 50 and 20 ppm. The higher O and Alcontent are indicative of a higher alumina inclusion con-tent. In heat 3, which contains 6 ppm Ca, only mixed oxideinclusions were observed. In heat 2, which contains 3 ppmCa, only alumina inclusions were observed. Ca is the onlycommon inclusionary element which can reduce alumina to

Fig. 9. Schematic of the normal stress (σγ ), shear stress (τγ ) and tem-perature (θ ) distribution on the rake face of cutting tools [21,26,27].


Fig. 10. (a) Formation of grooves on the flank face of alumina/TiC cutting tools once VBC exceeds a critical value. (b) Progression of flank wearaccording to Goh et al. [11] and its influence on the flank wear rate (c). (d) Ridges on the tool side of the chip, the cross-section is normal to the chipflow direction. (e) Rake face of the tool which generated the chip in (d).

form a mixed oxide inclusion; invariably, Ca is added to themolten steel only after the addition of Al. While a quan-titative analysis of the oxide inclusions in each steel wasnot feasible, the higher Al and O content of heat 2 in com-parison to heat 3 correlates with the more rapid tool wearrate.

In relation to the wear rate of the tools used to ma-chine the low alloy tool steel; it was noted above thatthe mixed oxide content of this steel was extremely lowand, thus, despite a Ca content of 24 ppm, the frequencyof interaction between such inclusions and the cuttingtool was insufficient for the chemical wear mechanism tocontribute appreciably to the wear of the tools. Note thedifference in the morphology of the wear surfaces of thetools used to machine the low alloy tool steel and heat1 of BS 817M40 steel. Consequently, the relatively mod-erate rate of tool wear, approximately twice that of the

Fig. 11. Thrust force (a) and cutting force (b) as a function of tool pathin the finish turning of the low alloy tool steel of 52 HRC and 60 HRC(vc = 2.5 m/s,f = 100mm, ap = 200 mm).


Fig. 12. Cross-sections of continuous (a) and saw-tooth (b) chips produced during the orthogonal cutting of the BS 817M40 steel of 52 HRC (the sectionis taken normal to the direction of chip flow), (a)vc = 0.83 m/s, (b)vc = 5.0 m/s,f = 100mm.

tools used to machine heat 2 of BS 817M40 steel cannotbe attributed to hard oxide inclusions. However, anotherimportance characteristic of the low alloy tool steel is thepresence of micron-sized alloy carbides. In a transmissionelectron microscopy (TEM) study of machined surfaces[5], such carbides were observed to have undergone little orno plastic deformation, thus, indicating their potential forincreasing the degree of superficial plastic strain of the toolsurface.

Another point of interest in relation to tool wear in ma-chining the low alloy tool steel is the influence of workmaterial hardness. While the flank wear is almost identi-cal in machining the steels of 52 and 60 HRC, the level ofcrater wear is dramatically different. In addition, the cut-ting and thrust forces were virtually identical over the lifeof the tool, Fig. 11. The influence of work material hard-ness is generally to increase the temperatures in the chipformation zone. In machining an AISI 52100 bearing steel,(of similar composition and structure to the low alloy toolsteel used in the current study), with cutting speed,vc =2.5 m/s, Ueda et al. [19] noted the temperature on the toolflank to increase from 820 to 870◦C with an increase inwork material hardness from 510 to 700 HV (50 to 60HRC, approximately). While there is no data available onthe corresponding increase in the rake face temperatures, itis likely to be of a similar value, though for a given workmaterial hardness, the temperatures on the rake face arelikely to be several hundred◦C higher than on the flank[19,21].

According to the plasticity based wear mechanism dis-cussed above, the fact that the flank wear rate remains thesame with an increase in work material hardness from 52to 60 HRC suggests that the accompanying temperature rise(of the order of 50◦C [19]) is insufficient to activate addi-tional slip systems on the tool flank. It cannot be specifiedwhether additional slip systems are activated on the rakeface, however, the increase in the rate of crater wear maybe accounted for by a transition in the mechanisms of chipformation. In machining the steel of 60 HRC, a saw-toothchip is formed whereas in machining the steel of 52 HRC,a transitional chip is formed which shares characteristicsof both the continuous and saw-tooth chip types, i.e. thechip exhibits a low degree of segmentation. Cross-sectionsof a continuous and a saw-tooth chip, produced by orthog-onal cutting are shown in Fig. 12, (note that in addition tohardness, cutting speed influences chip morphology in hardturning).

The essential characteristic of saw-tooth chip formationis periodic catastrophic failure within the primary shearzone [28] which results in a high frequency variation incutting force [29], of the order 20–50 kHz (a frequencywell beyond the natural frequency of conventional plat-form dynamometers). As the time averaged cutting forceas measured with a platform dynamometer, remains in-variant with respect to work material hardness and chipmorphology, it follows that the peak transient stresses arelarger during the formation of the saw-tooth chip from theharder work material. It is feasible that the higher stresses


facilitate either the operation of additional slip systemsor overcome the increases in the critical resolved shearstresses due to prior strain. The influence of chip formationmechanics on the flank wear rate are probably of secondaryimportance.

5. Conclusions

A six-fold variation in the rate of alumina/TiC cuttingtool wear has been observed in the finish machining of dif-ferent heats of BS 817M40 steel of 52 HRC. In machin-ing steels containing Ca-bearing mixed oxide inclusions, areaction between the oxide inclusionary deposits and thealumina phase of the tool can lead to rapid tool wear. Acharacteristic of this wear mechanism is the formation of aridge/groove morphology over the rear of the crater surfaceand over the downstream end of the flank wear land once itswidth exceeds a critical value; in the current study, VBC ≈150mm.

In machining steels containing very low levels of Caor steels with a very low inclusion content, tool wear islargely effected by superficial plastic deformation of thetool surface. It appears that the rate of tool wear is deter-mined by the hard inclusion content or alloy carbide con-tent of the work material. Impingement of hard particlesagainst the tool surface results in the generation of tran-sient localised stresses which exceed the average contactpressures and, thus, either facilitate the operation of addi-tional slip systems or overcome the increases in the criticalresolved shear stresses on active slip systems due to priorstrain.

The influence of saw-tooth chip formation on the rateof cutting tool wear appears to be related primarily tothe resulting high frequency cutting force variation. Whilethe transition from continuous to saw-tooth formationdoes not result in a change in the time-averaged cuttingor thrust force (and, hence, the resultant tool force), itdoes result in a force variation such that the peak tran-sient stresses are greater than the average contact stressesduring continuous chip formation. Though thought notto occur in the current study, the temperature rise in thechip formation zone which accompanies an increase inwork material hardness, may also influence the rate of toolwear should it result in the operation of additional slipsystems.

Acknowledgements

This work was funded by Enterprise Ireland, the IrishState-Agency for science and technology, under their basicresearch grants scheme. Also acknowledged is the assistanceof the Irish National Heat Treatment Centre, Department ofMechanical Engineering, UCD and the Electron MicroscopyLaboratory, UCD.

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