Figure 172. Typical temperature distributions (isotherms) during machining, illustrated across the: chip, insert and work- piece; at relatively low cutting speed . Machinability and Surface Integrity secondary deformation zone tends to be linear in na- ture from: (˙γ int ) – at the interface, → zero – at the boundary of the triangular secondary zone. e frictional stress along the tool/chip interface can be assumed to be constant along the rst half of the contact region, then linearly decreasing to zero at its end. e frictional heat source distribution at this interface, can be obtained from stress and velocity dis- tributions at this location. In Fig. 173a, the basic ‘FEM mesh’ is shown, with typical temperature distributions obtained from this being illustrated in Fig. 173b. e accuracy of this particular example for the ‘Tay-model’ for the total sum of all heat sources was within 2.6% of actual mea- sured power consumption (F c U). Moreover, the values of ‘β’ calculated from the temperature distributions closely-agreed to those obtained some years earlier by Boothroyd (1963). e FEM approach to machin- ing data capture and analysis covers these and other related parameters and clearly indicates the power of simulation – more will be mentioned on this subject later in the chapter. 7.7 Tool Wear and Life Introduction e working environment for most machining pro- cesses is extremely harsh, with pressures exerted onto a minute area of tool tip being of the order of >1600 MPa, with localised temperatures reaching over 750°C creating a sterile surface at the tool/chip inter- face, making this an ideal state for a pressure-welding condition. In attempting to minimise this anity be- tween the work-hardened chip – oen this plastic de- formation making the chip >5 times harder than that of the parent workpiece material, means that there are several ways of relieving this tool/chip anity. e ob- vious one is to use a cutting tool material that is in- ert to the workpiece such as a either a: ceramic, or mixed-ceramic cutting insert composition, or some- thing similar, but this may not prove to be satisfactory, particularly if interrupted cutting conditions are antic- ipated. In this situation above, perhaps by utilising a multi-coated cemented carbide insert this may reduce this ‘adherence-tendency’. Lastly, the correct grade of ‘ood-coolant’ may: lower the interface temperature, reduce friction here, while somewhat improving the machined surface texture. When only partial success is achieved by employing the above tooling strategies, the last resort may be to adjust the cutting data to en- hance and provide a ‘less-abusive machining regime’ , while simultaneously improving the ‘steady-state’ wear conditions. So far, no mention has been made here concern- ing frictional eects in the cutting process. Friction is very complex subject which relates not only to: chip ow-stress and ‘stiction’ 60 problems at the chip/tool in- terface, but concerns the tribological conditions along this interface. Cutting tool rake and ank faces are never perfectly smooth, as even when faces and edges have been either been ground, or super-nished the abrasive nature of the super-nishing process pro- duces an abraded surface that approaches the grit size of the abrasive medium. erefore, to the naked eye the insert’s surface looks smooth, but at the ‘micron- level’ of surface magnication (i.e. 1 × 10 –6 m), the cut- ting insert’s surface has localised ‘high-spots’ , or as - perities present. ese asperities signicantly reduce the contact area produced between the forming chip and its contact at the interface on the tool’s rake face. Not only can these asperities considerably decrease the ‘real area of contact’ and as a result increase the coe- cient of friction here, but the asperities may be either ‘plastic’ , or ‘elastic’ in nature 61 . In Table 11 (i.e. exper- imental data extracted from: Childs, et al., 2000, con- cerning surface texture assessment of cutting insert faces), comparison is made between a small sample of 60 ‘Stiction’ , is sometimes confused with its ‘close alternative’ this being: ‘stick-slip’. ese terms are worth stating, to ex- plain their respective dierences and have been dened in the following manner: ‘Stiction’ is: ‘e phenomenon at an inter- face where the frictional stress is equal to the shear yield stress of the soer material.’ ‘Stick-slip’ is: ‘A jerky motion between sliding members due to the formation and destruction of junctions.’ (Kalpakjian, 1984) 61 ‘Plastic asperities – on a plastic chip’ , these are ‘high-spots’ that will sink into the chip and how they achieve this action, does not depend on local conditions at interface contact, but on the bulk plastic ow eld. Specically, the lower the hydro- static stress in the bulk ow eld, the less eort is required for these asperities to sink.‘ Asperities – on an elastic foundation’ , this situation is ex- tremely complex phenomena and put simply, in conditions of low contact stresses, the chip beneath these asperities is elas- tic. (Childs, et al., 2000) Chapter cutting insert surface conditions, clearly illustrating that even when ‘super-nishing’ 62 an insert’s face it still has asperities present. .. Tool Wear Introduction On a single-point turning tool’s cutting insert, the main regions of wear are normally conned to the: rake face; ank; trailing clearance face; together with 62 ‘Super-nishing’ , is based on the phenomenon that a lubricant of a given viscosity* will establish and maintain a separating lm between two mating surfaces, if their roughness does not exceed a specic value and, if certain critical pressure – keep- ing them apart – is not exceeded. us, as minute peaks occur on the cutting insert’s surface, they are then cut away by the abrasive (e.g. minute diamond abrasive in a lubricant – oil – suspension) this being applied with a controlled pressure – until a required level of smoothness has been achieved. NB e maximum stock removed from the insert will be ap- proximately 50 µm.(Degarmo et al., 2003) *Viscosity relates specically to oils, which will vary with temperature. Dierent oils vary by dissimilar amounts for the same temperature, this is why the ‘viscosity index’ (VI) has been developed. the actual nose radius (Fig. 174). Likewise, the type of wear pattern provides important information as to the eectiveness of the overall machining operation. Considerable time and eort has been spent by both researchers and tooling companies, ensuring that tool wear mechanisms and their respective classications for specic machining operations are understood. So, by knowing the anticipated wear behaviour for a cut- ting insert for a specic machining operation, this al- lows the user to optimise productivity by ensuring that the ‘ideal’ tool grade and its associated geometry, will produce the desired machining conditions with the correct type of cut for the chosen workpiece materi- al’s composition. A range of factors can inuence tool wear when component machining, these are: material removal rate; ecient chip control; machining eco- nomics, precision and accuracy demanded; plus the machined surface texture requirements. If one magnies then inspects the wear pattern on a worn cutting edge, then it is reasonably straightfor- ward to establish both the cause and remedy for the indicated type of wear (i.e see Appendix 11), this will allow subsequent tooling to be more adequately con- trolled during following machining operations. In order to ensure that the correct tool has been selected, it is really only down to basic ‘good engineering prac- tices’ , namely: • that the initial selection of criteria for the cutting data is sound; Table 11. Cutting insert surface texture and contact stress severity data. ← 10k local /E* [°] → Tool finish: Surface texture data: Al/HSS: Cu/HSS: Brass/WC Steel/WC Ra [µm] ∆q [°] CVD-coated 0.2-0.5 3-7 1.2 1.9 2.8 1.8 Ground 0.1-0.25 2-4 1.2 1.9 2.8 1.8 Super-nished 0.03 0.4 1.2 1.9 2.8 1.8 * When s/k is <0.5, an asperity is totally elastic – if the plasticity index is <5 and totally plastic if its >50. As s/k increases to 1, these critical plasticity index values reduce. In large s/k conditions of metal machining, an asperity would normally be ‘fully-plastic’ , if: ∆q ≥ 10k local /E*. NB ‘s’ = Shear strength and ‘k’ = local shear stress. [Source Childs, et al., 2000] . Machinability and Surface Integrity Figure 173. Finite Element Method (FEM), to obtain simulated, but realistic data on isother- mal temperatures within the cutting region. [Source: Tay et al., 1993] . Chapter Figure 174. Typical wear patterns that could be present on a cemented carbide (uncoated) cutting insert, utilised under ‘steady-state’ turning conditions . Machinability and Surface Integrity • good quality and consistent workpiece material is to be utilised; • that the condition monitoring of machine tool en- sures that it is in an optimum state for use; • any ood coolant supply and quality – if it is to be used – is of the correct grade and dilution concen- tration; • work-holding/support is both rigid and precise/ac- curate; • expert support is available – if necessary – along with the user’s own practical experiences. ese factors oer a good ‘start-point’ in ensuring that the ‘ideal’ tool wear development takes place. Classification of Tool Wear Types Tool wear depends on several inter-related factors, some of these have been mentioned above, but are worth restating, such as: the cutting insert and work- piece material combination – plus their physical, mechanical and chemical properties; cutting insert ge - ometry; as well as cutting uid properties and pressure – if applied; together with various other operational parameters – cutting data selected, stability of the cut- ting process and work-holding application techniques. Any knowledge obtained on analytical studies of wear mechanisms, is largely based upon the results from ex- perimental trials. Simply obtaining wear data presents considerable diculties, then simply analysing these results can be somewhat onerous, due to isolating the major cause of this particular wear regime. Neverthe- less, having stated these problems, many potential so- lutions to specic wear patterns can be found, so long as the actual wear regime, or composite wear behav- iour can be singularly identied. With this in mind, the following classications for tool wear are given be- low (i.e. see Fig. 174 for of several these wear patterns), which include: • Flank wear – as its title suggests, occurs on the cut- ting edge’s anks, usually the result of an abrasive wear mechanism. Both of the clearance faces – lead- ing and trailing edges, together with the tool nose radius are subject to a parallel land wear, created by the workpiece travelling past the contact regions of the tool both during and aer chip formation. Such a wear mechanism is considered normal tribologi- cal behaviour and a progressive form of ank wear can be tolerated and subsequently dealt with, by an ecient tool-changing strategy, based upon antici- pated tool life expectancy. NB Toward the end of the steady-state and progres- sive ank wear regime, it could lead to several un- desirable factors, such as: increasing friction, which can possibly change the insert’s prole – leading to poor machined surface texture, or dimensional in- accuracies as the ‘tool dris’ 63 – creating variability in tolerances of successive parts. • Crater wear – this is present on the rake, or chip face and is normally the result of a combination of an abrasion and diusion 64 wear mechanism. 63 ‘Tool driing’ , is a term used to describe the fact that having initially set the tool to a particular dimensional size, the tool’s ank will progressively wear – under steady-state machin- ing conditions. e variability in dimensional size can be the subject of both random and systematic errors – even when the operation is behaving normally. is dimensional variabil- ity, causes for example: turned diameters to get larger, while drilled holes get smaller – as successive components are ma- chined, this is the essence of tool-driing. e term process capability* has been coined to explain the stochastic process output from a normally-operating production process – see Chapter 2, Footnote 26, for more information regarding this subject. *Process capability (C p ) can change during consecutive pro- duction output of components, being the result of the ‘vari- ables’ (i.e. as each singular part dimension is known), pro- ducing either random, or systematic errors, or both, as the production run progresses. is is why it is usual practice to utilise ‘Statistical control techniques’ to show any signicant changes in output. erefore, ‘Shewart charting techniques’ in combination with ‘Probability paper’ are employed, to esti- mate the: C p value and to determine if the process is behaving/ operating ‘normally’ – usually a ‘normal output’ is signied by establishing a ‘straight-line’ (i.e. plotted) relationship on the ‘Probability paper’. 64 ‘Diusion wear’ , was initially proposed in 1858 by the Ger- man physiologist Adolph Fick (1829–1901), where he enun- ciated laws governing the diusion of substances generally on a quantitative basis. Today, we are concerned with ‘atomic migration’ within metallic solid solutions. Fick produced two laws, with Fick’s 1 st Law stating: ‘at the amount (J) of a ma- terial moving across a unit area of a plane in unit time is pro- portional to the concentration gradient (∂c/∂x) at the same time but of opposite sign’. It can be expressed as follows: J[atoms/m 2 .s] = − D [m 2 /s](∂c/∂x)[atoms/m 3 .1/m] Fick’s 1 st Law Where: J = ux, net ow of atoms; D = diusion coe- cient; ∂c/∂x = concentration gradient. NB Assuming that X-axis is parallel to direction in which concentration gradient is operating. Fick’s 2 nd Law was de- rived from the 1 st Law and from the fact that matter is con- served, relating the change in concentration with time (∂c/∂t) and it can be expressed as: (∂c/∂t) = ∂/∂x (D∂c/∂x) Fick’s 2 nd Law (General case) By dierential calculus, this 2 nd Law changes to: ∂c/∂t) = D ∂ 2 c/∂x 2 . Chapter e crater can be formed either via a hard-particle grinding action, which mechanically-removes rake face surface layers, or by a complex ‘atomic diusion process’ 65 interacting between the chip and the tool material (ie see Fig. 174 – top right). NB If a cutting insert has high bulk hardness, combined with ‘hot-hardness’ 66 , plus minimum af- nity between these two materials, this will dimin- ish any crater wearing tendencies. Moreover, crater wear changes the cutting insert geometry of the edge, which may impair chip formation and modify cutting forces, or lead to a weakened edge strength. Many of today’s multi-coated cutting inserts are less aected by crater wear than their uncoated coun- terparts. NB From this it can be appreciated why the nal stages of dif- fusion are somewhat slow, due to the rate of diusion decreas- ing as the concentration gradient diminishes. (Higgins, 1979) 65 ‘Atomic diusion process’ , there is strong evidence – when ferrous workpiece machining – to indicate that cratering of WC-Co cutting inserts (i.e. uncoated), occurs by diusion of the C atoms into chip at the interface (i.e see Fig. 174 – top right schematic diagram). Remembering that solid-state dif- fusion depends upon the rate at which the tool’s atoms dis- solve/diuse into the chip. For WC, the most rapid diusion is by the tool’s Co atoms – of the carbide bond and, the Fe atoms from the chip. Hence the carbide grains are undermined and swept-away for two reasons:With WC tool material, carbide grains are not isolated and constitute the bulk of the mate- rial, so support each other in a ‘rigid framework’ ,Due to Co atoms from the tool ‘diusing-out’ , so Fe atoms from the chip ‘diuse-in’ and these provide support for the carbide grains, which in turn inhibit their removal. In the chip, C atoms being small, rapidly diuse through the Fe matrix, however those in the tool are strongly-bonded to W and are not free to move by themselves. us, the rate of diusion of both W and C atoms together from the tool go into the chip and thus, will control diusion wear with respect to its temperature – as Fick’s Laws suggest. NB e distances for diusion at the tool/chip interface are between 1 nm up to 1µm. Diusion in the tertiary shear zone (i.e. ank) is normally higher than in the secondary shear zone, due to the signicantly greater workpiece surface speed in this vicinity. So, not only is attrition a mechanism for ank wear, diusion is also partly responsible – even when the rake face is hardly worn. In appearance, when the grains look to be smooth, this is a good indication of a diusion mechanism taking place. (Armarego and Brown, 1969) 66 ‘Hot hardness’ , this is the ability of a cutting insert to retain its relative bulk hardness and hence geometry at elevated tem- peratures. • Plastic deformation – occurs when high pressures (i.e. compression) are exerted on the cutting edge in combination with elevated temperatures. Con- ditions likely to create plastic deformation on the cutting insert are when high speeds and feeds are utilised on workpiece materials that are prone to work-hardening. Tool materials must have the re- quired mechanical properties to withstand plastic deformation during machining. Typically, bulging of the edge in the tool nose region, leads to: geom- etry deformation; chip ow modication; greater localised temperatures – until a critical juncture is attained. So cutting insert ‘hot-hardness’ is a vital characteristic. NB In order to combat cutting insert plastic defor- mation, a large tool nose radius, plus more robust tool geometry adds greater strength in this ‘exposed region’ of the tool. • Notch wear on insert’s leading edge – is the result of mechanical action, promoted by either machining workpiece materials that may easily work-harden, so each successive longitudinal turning pass at the same D OC leads to the previous surface condition being harder, resulting in a more abrading-action here – hence a notch will wear at this point on the insert‘s ank. is ‘notching eect‘ can be reduced, if a variable D OC is employed, to ‘even-out’ the con- tact region along the leading edge of the insert. NB ‘Black-bar stock’ having been hot-rolled from its primary processing route, tends to have a hard and abrasive oxide scale to its periphery, which may contribute to insert notching when only the surface is ‘skimmed’ by a longitudinal turning operation. • Notch wear on insert’s trailing edge – occurs by in the main, by adhesion wear, but to a lesser extent, may be the result of an oxidation wear mechanism. e notch on this ank’s trailing edge is formed where the cutting edge and the workpiece material separate. NB Notch wear here, tends to be very localised to- ward the end of the cut, enabling air to reach this cutting vicinity, which has a high temperature pres- ent, so adhesion/oxidation can be expected. • Built-up edge (BUE) formation – is usually the re- sult of tool/workpiece anity associated with tem- Machinability and Surface Integrity perature and its respective cutting speed (i.e. see Fig. 28). Moreover, it can also transpire as a result of ‘edge agging’ , or from other wear mechanisms. is ‘cold’ pressure-welded workpiece material be- ing attached to the tool as a BUE, changes the cut- ting insert’s geometry – to its detriment. Hence, this BUE is both severely work-hardened and ‘unstable’ – it will break-away from the tool mate- rial thereby potentially ‘frittering’ the insert’s edge. NB BUE machining data conditions have been reasonably well-dened, so fortunately, these re- spective cutting speeds can be avoided, particu- larly, as most CNC machining operations happen at much higher speeds and modern insert grades and coatings, minimise this BUE eect. If BUE does oc- cur, it can create a poor surface nish on the ma- chined surface. In any BUE machining condition, if it continues without attention, then the result can be rapid edge breakdown, or even result in insert fracture. • e former conditions are in the main, conned to continuous cutting and steady-state machining conditions, albeit with single-point cutting inserts. • e latter conditions are generally restricted to in- termittent cutting multi-point machining, or inter- rupted cutting operations: • ermal cracking – is usually the result of fatigue wear, produced by thermal cycling machining con- ditions, such as when milling. ese cracks that form are normally at 90° to that of the cutting edge 67 . ese cracks are spaced out periodically along the cutting edge and when they propagate (i.e. grow) to 67 ‘ermal fatigue cracks’ , are usually termed ‘comb-cracks’ – due to their appearance is not unlike that of a hair comb. When these cracks propagate to a critical length which can be ex- plained in terms of ‘Fracture mechanics’* and in particular the ‘stress intensity factor’ (K IC ) – with the ‘C’ standing for ‘critical’. Such cracks will fracture quickly around the ‘Speed of sound’ (i.e. Mach 1, or in a steel workpiece @ 5050 ms –1 ), so little, if any warning is given of the likely failure condition as it arises – when the tool’s edge eventually catastrophically fails. *In 1957, G.R. Irwin and his co-workers, laid the foundations for ‘Fracture mechanics’ and were particularly noted for the mathematics for dening the ‘stress intensity factor’ (K), spe- cically: K = σ √ (πc) [Nm ½ ] Where: σ = fracture stress, c = half length of an internal aw. (Shaw, 1984) a critical size, bulk tool material will be pulled-out of the tool’s edge – leading to a very rapid type of cutting insert edge failure. NB Varying the chip thickness will also aect tem- peratures throughout the cut. A cautionary note here, concerning cutting uid application: if used under certain conditions, the cutting uid has a detrimental inuence in some metal cutting opera- tions, as it amplies the variations in temperature between and in- and out-of-cut. • Mechanical fatigue cracking – may be present if cutting force shock-loads are extreme. Fatigue 68 is a form of fracture which is promoted by continual variations in load, but where the load in itself, is not great enough to cause fracture. 68 ‘Fatigue’ , can be dened as a: ‘Phenomenon leading to the fail- ure of a part under repeated, or uctuating stress below the ten- sile strength of the material.’ Failure usually occurs suddenly as a result of crack propagation without plastic deformation at a stress level well below that of the elastic limit for the material. e stress can be either an: ‘alternating’; ‘repeated’; or a combi- nation of these types. At a discontinuity such as a notch, hole, or step, the stress is considerably greater and is termed a ‘stress concentration factor’ (K). Graphs can be plotted , such as: SN curves (i.e. to nd the endurance limit for steels, or for non-ferrous metals, alloys and plastics -the fatigue stress ‘σ FS ’ is specied for a nite number of stress reversals), Soderberg diagram – for steel, with alternating stress plot- ted against steady stress. Moreover, a ‘safety factor’ (FS) can be applied to the graphical result, as follows: (Safety factor) FS = σ y σ m +(σ y �σ e )K σ r Where: σ y = yield stress, σ m = steady stress component, σ e = failure occurs – (i.e. above a line drawn from this value: σ e on the ‘Y-axis’ to σ u on the ‘X-axis’); Kσ r = alternating com- ponent – with ‘K’ representing the ‘stress concentration factor’ and ‘σ r ’ representing ‘alternating stress’. NB Most steels have an ‘endurance limit’ being about half its tensile strength, with an approximation oen utilised: For steels: Endurance limit = 0.5 tensile strength (i.e. up to a tensile strength of 1400 N mm –2 ), Endurance limit = 700 N mm –2 (i.e. above a tensile strength of 1400 N mm –2 ). For Cast steel/iron: Endurance limit = 0.45 tensile strength (i.e. up to tensile strength of 600 N mm –2 ), Endurance limit = 275 N mm –2 (i.e. above a tensile strength of 600 N mm –2 ). Non-ferrous metals/alloys: there is no endurance limit and the fatigue stress is taken at a denitive value of stress rever- sals, e.g. 5 x 10 7 . (Carvil, 1994, et al.) – – Chapter NB erefore at the initiation of a cut, the varia- tions in the magnitude of the cutting force and its direction, may not be too great for both the tough- ness and strength of the cutting insert. With con- tinual usage however, these fatigue cracks grow – in the main – parallel to the cutting edge and may eventually be the cause for premature tool failure. • Cutting edge chipping – this transpires when the edge line fractures, rather than being the result of wear. It can be considered as a form of fatigue fail- ure, because of the cycles of loading and unloading during cutting, leading to particles of tool material being removed from the insert’s surface. is type of wear mechanism is generally the result of inter- mittent cutting operations. NB An investigation into whether this edge wear is either from chipping, or the result of ank wear. ‘Spalling’ (i.e. cracking, or aking of the surface) and ‘nicking’’ are also variants of this category of edge degeneration. • Fracture – is normally catastrophic conclusion to the cutting process (i.e. see Fig. 175). Here, bulk material fracture can have serious consequences obviously to the cutting insert, but also aecting the machined part. Moreover, this form of edge fracture is more oen than not, the termination of alternative wear regimes. If Fig. 175 is investigated in more detail, it may help comprehension of the nature of the serious problems associated with such a sudden failure mode. e cut- ting insert was purposely catastrophically failed in practical trials conducted by the author, using a rea- sonably robust turning and facing geometry, longitu- dinal turning P/M ferrous compacts without coolant. Here, the cutting speed was raised by 25% above the optimum, with the feedrate 40% greater than usually specied. is ‘abusive machining regime’ , created high ank wear and plastic deformation to the cutting edge, which shortly failed – catastrophically. In Fig. 175c, detail of the fracture surface indicates both duc- tile and brittle failure modes instigated from the worn leading edge’s ank. By increasing the cutting data by just the cutting speed alone and leaving the feedrate at the optimum, tool life was reduced on other simi- lar inserts, but catastrophic failure did not occur, only very high levels of ank wear. However, if the cutting speed was kept at the optimum and the feedrate was increased – as mentioned – in-line with other insert trials, then catastrophic failure eventually occurred, well before that predicted by ‘Taylor’s tool life calcu- lation’ . is conrmed the fact that the high abrasive nature to the testpieces produced from ferrous-based P/M compacts, in combination with an increased fee- drate caused premature catastrophic failure of the cut- ting inserts during these ‘harsh’ machinability trials. As previously mentioned, Appendix 11 has a con- cise ‘trouble-shooting guide’ for some of the potential wear regimes that are likely to be experienced during many machining operations. .. Tool Life Introduction It is normal practise to assess tool life according to three mutually-inuencing criteria, as any one of them could be the reason for the expensive business of sub- sequent part scrappage. ese criteria that signicantly aect machined components and can be the reason for curtailment of the cutting tool’s life are: 1. Ability to sustain workpiece tolerances – here if the tool has been in operation for too long ‘in-cut’ , then this will increase the tendency for ‘tool dri- ing’ which will amplify machined component vari- ability, while creating inconsistency in part produc- tion (Figs. 31ci and ii), 2. Maintaining machined surface texture quality – as the tool is progressively utilised, the ank and cra- ter wearing tendencies will increase, leading to de- generation of the surface texture, below that which was demanded from the designer’s direct engineer- ing requirements (i.e. see graph in Fig. 148), 3. Eciency in chip-breaking ability – if the cut- ting insert/tool has been operated for considerable time, there is every expectation that both ank and more importantly crater wear will be present. is will have an adverse eect on chip-breaking ability, leading to either poor component surface texture, or variability in component tolerances, or both (Figs. 37 and 38a and b). If a cutting insert, or tool no longer satises the above wear criteria, its useful life is ended and it should be summarily discarded. e tool life’s predictability, is a key factor in an estimation of the anticipated produc- tivity output level. Approached from a dierent direc- tion, an CNC programmer may deliberately choose Machinability and Surface Integrity Figure 175. Catastrophic failure of a turning insert. Chapter . Figure 172 . Typical temperature distributions (isotherms) during machining, illustrated across the: chip, insert and work- piece; at relatively low cutting speed . Machinability and Surface Integrity. capture and analysis covers these and other related parameters and clearly indicates the power of simulation – more will be mentioned on this subject later in the chapter. 7. 7 Tool Wear and Life Introduction e. ≥ 10k local /E*. NB ‘s’ = Shear strength and ‘k’ = local shear stress. [Source Childs, et al., 2000] . Machinability and Surface Integrity Figure 173 . Finite Element Method (FEM), to obtain