Machinability and Surface Integrity Part 7 pptx

The 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

Figure 172 Typical temperature distributions (isotherms) during machining, illustrated across the: chip, insert and

work-piece; at relatively low cutting speed

.

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

The frictional stress along the tool/chip interface

can be assumed to be constant along the first half of

the contact region, then linearly decreasing to zero at

its end The 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 The 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 cU) Moreover, the values

of ‘β’ calculated from the temperature distributions

closely-agreed to those obtained some years earlier

by Boothroyd (1963) The 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

The 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 affinity

be-tween the work-hardened chip – often 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 affinity The

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

‘flood-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 effects in the cuttconcern-ing process Friction is very complex subject which relates not only to: chip

flow-stress and ‘stiction’ 0 problems at the chip/tool

in-terface, but concerns the tribological conditions along this interface Cutting tool rake and flank faces are never perfectly smooth, as even when faces and edges have been either been ground, or super-finished the abrasive nature of the super-finishing process pro-duces an abraded surface that approaches the grit size

of the abrasive medium Therefore, to the naked eye the insert’s surface looks smooth, but at the ‘micron-level’ of surface magnification (i.e 1 × 10– m), the cut-ting insert’s surface has localised ‘high-spots’ , or as-perities present These asas-perities significantly 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 coeffi-cient of friction here, but the asperities may be either

‘plastic’ , or ‘elastic’ in nature 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

0 ‘Stiction’ , is sometimes confused with its ‘close alternative’ this being: ‘stick-slip’ These terms are worth stating, to

ex-plain their respective differences and have been defined in the

following manner: ‘Stiction’ is: ‘The phenomenon at an

inter-face where the frictional stress is equal to the shear yield stress of the softer material.’

‘Stick-slip’ is: ‘A jerky motion between sliding members due to

the formation and destruction of junctions.’ (Kalpakjian, 1984)

 ‘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 flow field Specifically, the lower the

hydro-static stress in the bulk flow field, the less effort 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)

cutting insert surface conditions, clearly illustrating

that even when ‘super-finishing’  an insert’s face it still

has asperities present

7.7.1 Tool Wear

Introduction

On a single-point turning tool’s cutting insert, the

main regions of wear are normally confined to the:

rake face; flank; trailing clearance face; together with

 ‘Super-finishing’ , is based on the phenomenon that a lubricant

of a given viscosity* will establish and maintain a separating

film between two mating surfaces, if their roughness does not

exceed a specific value and, if certain critical pressure –

keep-ing them apart – is not exceeded Thus, 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 The maximum stock removed from the insert will be

ap-proximately 50 µm.(Degarmo et al., 2003)

*Viscosity relates specifically to oils, which will vary with

temperature Different 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 effectiveness of the overall machining operation Considerable time and effort has been spent by both researchers and tooling companies, ensuring that tool wear mechanisms and their respective classifications for specific machining operations are understood So,

by knowing the anticipated wear behaviour for a cut-ting insert for a specific 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 influence tool wear when component machining, these are: material removal rate; efficient chip control; machining eco-nomics, precision and accuracy demanded; plus the machined surface texture requirements

If one magnifies 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* [°] →

Ra [µm] ∆q [°]

* 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 ≥ 10klocal/E*.

NB ‘s’ = Shear strength and ‘k’ = local shear stress.

[Source Childs, et al., 2000]

.

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]

.

Figure 174 Typical wear patterns that could be present on a cemented carbide (uncoated) cutting insert, utilised

under ‘steady-state’ turning conditions

.

• 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 flood 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

These factors offer 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 fluid 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 difficulties, 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 specific wear patterns can be found, so long

as the actual wear regime, or composite wear

behav-iour can be singularly identified With this in mind,

the following classifications 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 flanks, usually the result of an abrasive

wear mechanism Both of the clearance faces –

lead-ing and traillead-ing 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 after chip formation Such

a wear mechanism is considered normal

tribologi-cal behaviour and a progressive form of flank wear

can be tolerated and subsequently dealt with, by an

efficient tool-changing strategy, based upon

antici-pated tool life expectancy

NB Toward the end of the steady-state and

progres-sive flank wear regime, it could lead to several un-desirable factors, such as: increasing friction, which can possibly change the insert’s profile – leading to poor machined surface texture, or dimensional

in-accuracies as the ‘tool drifts’  – 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 diffusion wear mechanism

 ‘Tool drifting’ , is a term used to describe the fact that having

initially set the tool to a particular dimensional size, the tool’s flank will progressively wear – under steady-state machin-ing conditions The variability in dimensional size can be the

subject of both random and systematic errors – even when

the operation is behaving normally This 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-drifting The 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 This is why it is usual practice to utilise ‘Statistical control techniques’ to show any significant changes in output Therefore, ‘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 signified by establishing a ‘straight-line’ (i.e plotted) relationship on the

‘Probability paper’.

 ‘Diffusion wear’ , was initially proposed in 1858 by the

Ger-man physiologist Adolph Fick (1829–1901), where he enun-ciated laws governing the diffusion 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  st Law stating: ‘That 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  s] = −  D  [m /s](∂c/∂x)[atoms/m 1/m] Fick’s   st   Law Where: J = flux, net flow of atoms; D = diffusion coeffi-cient; ∂c/∂x = concentration gradient.

NB Assuming that X-axis is parallel to direction in which concentration gradient is operating Fick’s   nd  Law was de-rived from the 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  nd  Law (General case) By differential calculus, this 2 nd

Law changes to: ∂c/∂t) = D ∂c/∂x.

The crater can be formed either via a hard-particle

grinding action, which mechanically-removes rake

face surface layers, or by a complex ‘atomic diffusion

process’  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’ , plus minimum

af-finity 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

affected by crater wear than their uncoated

coun-terparts

NB From this it can be appreciated why the final stages of

dif-fusion are somewhat slow, due to the rate of difdif-fusion

decreas-ing as the concentration gradient diminishes (Higgins, 1979)

 ‘Atomic diffusion process’ , there is strong evidence – when

ferrous workpiece machining – to indicate that cratering of

WC-Co cutting inserts (i.e uncoated), occurs by diffusion 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/diffuse into the chip For WC, the most rapid diffusion 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 ‘diffusing-out’ , so Fe atoms from the chip

‘diffuse-in’ and these provide support for the carbide grains,

which in turn inhibit their removal In the chip, C atoms being

small, rapidly diffuse through the Fe matrix, however those in

the tool are strongly-bonded to W and are not free to move by

themselves Thus, the rate of diffusion of both W and C atoms

together from the tool go into the chip and thus, will control

diffusion wear with respect to its temperature – as Fick’s Laws

suggest

NB The distances for diffusion at the tool/chip interface are

between 1 nm up to 1µm Diffusion in the tertiary shear zone

(i.e flank) is normally higher than in the secondary shear

zone, due to the significantly greater workpiece surface speed

in this vicinity So, not only is attrition a mechanism for flank

wear, diffusion 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 diffusion mechanism

taking place (Armarego and Brown, 1969)

 ‘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 flow modification; 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 DOC 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 flank This ‘notching effect‘ can be reduced,

if a variable DOC 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 The notch on this flank’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 affinity associated with

tem-perature and its respective cutting speed (i.e see

Fig 28) Moreover, it can also transpire as a result

of ‘edge flagging’ , or from other wear mechanisms

This ‘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-defined, 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 effect If BUE does

oc-cur, it can create a poor surface finish 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

• The former conditions are in the main, confined

to continuous cutting and steady-state machining

conditions, albeit with single-point cutting inserts

• The latter conditions are generally restricted to

in-termittent cutting multi-point machining, or

inter-rupted cutting operations:

• Thermal cracking – is usually the result of fatigue

wear, produced by thermal cycling machining

con-ditions, such as when milling These cracks that

form are normally at 90° to that of the cutting edge

These cracks are spaced out periodically along the

cutting edge and when they propagate (i.e grow) to

 ‘Thermal 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’ (KIC) – 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–), 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 defining the ‘stress intensity factor’ (K),

spe-cifically:

K = σ √ (πc) [Nm ½]

Where: σ = fracture stress, c = half length of an internal flaw

(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 affect

tem-peratures throughout the cut A cautionary note here, concerning cutting fluid application: if used under certain conditions, the cutting fluid has a detrimental influence in some metal cutting opera-tions, as it amplifies the variations in temperature between and in- and out-of-cut

• Mechanical fatigue cracking – may be present if

cutting force shock-loads are extreme Fatigue8 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

 ‘Fatigue’ , can be defined as a: ‘Phenomenon leading to the

fail-ure of a part under repeated, or fluctuating 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

The 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 find the endurance limit for steels, or for

non-ferrous metals, alloys and plastics -the fatigue stress

‘σFS’ is specified for a finite 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= σy

σm+(σy�σe)Kσr

Where: σy = yield stress, σm = steady stress component,

σe = failure occurs – (i.e above a line drawn from this value:

σe on the ‘Y-axis’ to σu on the ‘X-axis’); Kσr = alternating com-ponent – with ‘K’ representing the ‘stress concentration factor’ and ‘σr’ representing ‘alternating stress’.

NB Most steels have an ‘endurance limit’ being about half its

tensile strength, with an approximation often utilised:

For  steels: Endurance limit = 0.5 tensile strength (i.e up to

a tensile strength of 1400 N mm–), Endurance limit = 700

N mm– (i.e above a tensile strength of 1400 N mm–).

  For Cast steel/iron: Endurance limit = 0.45 tensile strength (i.e

up to tensile strength of 600 N mm–), Endurance limit = 275

N mm– (i.e above a tensile strength of 600 N mm–).

  Non-ferrous metals/alloys: there is no endurance limit and

the fatigue stress is taken at a definitive value of stress rever-sals, e.g 5 x 10 (Carvil, 1994, et al.)

– –

NB Therefore 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 This 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 flank wear

‘Spalling’ (i.e cracking, or flaking 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 affecting

the machined part Moreover, this form of edge

fracture is more often 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 The

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

specified This ‘abusive machining regime’ , created

high flank 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 flank 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 flank 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’ This confirmed 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

7.7.2 Tool Life

Introduction

It is normal practise to assess tool life according to three mutually-influencing criteria, as any one of them could be the reason for the expensive business of sub-sequent part scrappage These criteria that significantly affect 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 drift-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 flank 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 Efficiency in chip-breaking ability – if the

cut-ting insert/tool has been operated for considerable time, there is every expectation that both flank and more importantly crater wear will be present This will have an adverse effect 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 satisfies the above wear criteria, its useful life is ended and it should be

summarily discarded The tool life’s predictability, is a

key factor in an estimation of the anticipated produc-tivity output level Approached from a different direc-tion, an CNC programmer may deliberately choose

Figure 175 Catastrophic failure of a turning insert

.