Astakhov V. Tribology of Metal Cutting

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Improvements of Tribological Conditions 377

• Coating does change the dimensions of the cutting tool. Change depends on the coat-

ing, its speciﬁed thickness and the coating process. Typically, PVD is recommended

for high-tolerance tools and CVD for loose tolerance tools. Most PVD coatings add

2–3 µm per side to a tool or component. CVD and PVD CrN are thicker and can

add 10 µm or more in some cases. Processing temperatures may grow or shrink

some substrate materials. CVD temperatures, in particular, affect the heat treatment

conditions of tools and components and can cause dimensional changes.

• Coatings are often applied in multiple alternating layers. This is because the hardness

increases as its grain size decreases, and the grain size decreases simultaneously

with the decrease in coating thickness [76]. This is especially true for alumina

coatings; thinner alumina coating layer are harder. Harder coatings provide better

wear resistance. The desirable maximum thickness: minimum grain size was not

attainable before the introduction of alternating multi-layer coating [76]. The most

common method of achieving the reduction of grain size in Al

layers is to

periodically interrupt their deposition by applying a thin layer of TiC, TiCN or TiN.

Growth of grains inhibited in each succeeding Al

layer. The application of more

layers in this way can be used to build up to the desired overall coating thickness

while keeping grain growth in Al

layers to a minimum.

• The effectiveness of various coatings depends on the type of machining operation

and machining regime. In low-speed end milling, wear resistance is determined by

the presence of chromium in the coating while at high-speed end milling TiAlN coat-

ings assures high wear resistance. In turning, the effectiveness of different coatings

depends on the cutting speed, as shown in Fig. 6.17.

• The best results are achieved when multi-layer coatings like (TiAl)N, (AlTi)N and

(AlTiCr)N coatings are used. Their use: (a) reduces the strength of adhesion bonds at

n(m/min)

T(min)

400

300200100

Fig. 6.17. Inﬂuence of cutting speed on tool life for cutting inserts with various coatings.

Longitudinal turning, no cutting ﬂuid. Work material: AISI steel 4140, HB200, cutting feed

f = 0.26 mm/rev, depth of cut d

= 3 mm. Tool geometry: cutting edge angle κ

= 60

◦

normal ﬂank angle α

= 7.5

◦

, normal rake angle γ

=−7.5

◦

Substrate – carbide M30. Coatings:

(1) TiCrN, (2) AlTiN, (3) no coating, (4) TiN and (5) TiCrN.

378 Tribology of Metal Cutting

the tool–chip interface and thus reduces the severity of the friction at this interface,

(b) improves tool life; (c) increases machining superﬁcial and in-depth residual

stresses because greater portion of the thermal energy generated in machining ﬂows

into the workpiece [77].

In recent years, new diamond coatings were developed for machining nonmetallic or

nonferrous work materials. Earlier, polycrystalline diamond (PCD) tools and “hard”

carbide grades were used for this purpose. Now, a variety of carbon-based cutting tool

coatings are available. They are termed CVD diamond, diamond-like carbon (DLC),

amorphous diamond, hard carbon and a host of brand name designations. In reality,

these coatings can be grouped into three generic categories: true diamond, DLC and

hard carbon.

Basic properties of DLC are listed in Table 6.10. Amorphous (a-C) and hydrogenated

amorphous carbon (a-C:H) ﬁlms have high hardness, low friction, electrical insula-

tion, chemical inertness, optical transparency, biological compatibility, ability to absorb

photons selectively, smoothness and resistance to wear. For a number of years, these

economically and technologically attractive properties have drawn almost unparalleled

interest towards these coatings. Carbon ﬁlms with very high hardness, high resistivity

and dielectric optical properties, are now described as diamond-like carbon or DLC [78].

Several methods have been developed for producing diamond-like carbon ﬁlms:

1. Primary ion beam deposition of carbon ions (IBD).

2. Sputter deposition of carbon with or without bombardment by an intense ﬂux of

ions (physical vapor deposition (PVD)).

3. Deposition from an RF plasma, sustained in hydrocarbon gases, onto substrates

negatively biased (plasma-assisted chemical vapor deposition (PACVD)).

Until recently, the work on DLC worldwide has not yielded the expected beneﬁts in the

ﬁeld of wear resistance and general mechanical performance. Most of the success has

Table 6.10. Some basic properties of diamond and DLC materials.

Property Thin ﬁlm Bulk

CVD diamond a-C a-C:H Diamond Graphite

Form Faceted crystals Smooth or rough Smooth Faceted crystals

Hardness (HV) 3000–12 000 1200–3000 900–3000 7000–10 000

Density (g/cm

) 2.8–3.5 1.6–2.2 1.2–2.6 3.51 2.26

Refractive index – 1.5–3.1 1.6–3.1 2.42 2.15

Electrical

resistivity

(Ω/cm

)

> 10

–10

> 10

0.4

Thermal

conductivity

(W/m

◦

1100 – – 2000 3500

Chemical stability Inert Inert Inert Inert Inert

Improvements of Tribological Conditions 379

been in applications for magnetic storage media and optical coatings. The reasons for

this are:

• only thin coatings (<1 µm) have been used,

• the 2D aspect of most of the deposition routes,

• the difﬁculty in gaining good adhesion to metallic substrates.

One of the main problems with DLC deposition at low temperature is the creation of

very high internal stress levels in the ﬁlms. This, combined with the ensuing lattice

mismatch when DLC is applied to a wide range of substrates, commonly leads to poor

adhesion. In high mechanical stress applications, the adhesion of the ﬁlms is of paramount

importance. This problem has now been overcome by ensuring that there are no stress

concentrations near the coating–substrate interface. The magnetron sources are used to

deposit a series of multi-layer compounds prior to the deposition of DLC. The layers have

graded interfaces. This ensures that there are no abrupt changes in composition, and that

the stress is introduced into the ﬁlm gradually. The optimum multi-layer structure series

is: titanium, titanium nitride, titanium carbonitride, titanium carbide and then the DLC.

It has also been subsequently found that the mechanical properties of the hard carbon

ﬁlms can be improved by incorporating a small percentage of metal dopant (usually ∼5%

titanium) in the ﬁnal carbon structure.

The resulting ﬁlms have excellent friction and wear properties:

• microhardness of up to 4000 HV,

• coefﬁcient of friction during dry running in air against cemented WC Tungsten–

Cobalt alloys < 0.15,

• wear rate 20% of that of titanium nitride,

• exceedingly low counterface wear.

The application of DLC coating on cutting tools resulted in signiﬁcant improvement in

tool life in some applications and allowed dry cutting in some applications. However,

the major problem is a great scatter in tool life results, although the mean tool life is

higher than that with TiN-coated tools. This is because the DSL ﬁlm lacks adhesion to

the cutting tool surface so it is peeled off at early stage of the cutting process and thus the

tool could not utilize the discussed advantages of the DSL ﬁlm. Recently, some research

teams found different ways to increase the adhesion properties of DSL coatings [79].

6.3 Improving Tribological Conditions by Modiﬁcation of Properties

of the Work Material

In the author’s opinion, one of the major hidden resource in improving tribological

conditions in metal cutting (reduced tool wear and improved quality of the machined

surface), and thus reducing the cost per unit of the machined part is in the properties

and structure of the work material. Unfortunately, this simple and self-obvious resource

380 Tribology of Metal Cutting

did not prevent sufﬁcient attention of researchers and practitioners in the ﬁeld of metal

machining although the properties of the work material deﬁne to a large extent:

• Properties of the chip contact surface which is the counterpart of the tool rake face

over the tool–chip contact length.

• Properties of the work material at the tool–workpiece interface (the ﬂank surface),

which deﬁne the contact conditions and tool wear over this interface.

It is interesting to point out that the properties of the tool materials and coating at the

discussed interfaces are subjects of a great body of theoretical and experimental studies

from the shop ﬂoor level (for example [80,81]) to high level of applied physics (for

example [82–84]), while among the properties and conditions of the work material in

many studies in the ﬁeld, only generic (not actual) chemical composition (or AISI (SAE,

ASM) designation like AISI steel 1045, for example), hardness and material dealer state

(for example, as rolled, annealed, hot rolled, cold rolled, extruded etc.) are mentioned.

In the author’s experience, these are not sufﬁcient even to the ﬁrst approximation to

characterize the machinability of the work material and thus the contact processes at the

tool–chip and tool–workpiece interfaces. As a result, a great scatter in experimental data

is a nuisance in experimental metal cutting.

When it comes to industry, one expects that high costs of poor machinability, great

scatter in tool life (that particularly hurts in production-automated lines and manufac-

turing cells with no or minimum human attendance) and great scrap rate prevented at

least some attention of practical manufacturing engineers to the properties of the work

material. Unfortunately, this is not the case even in the most advanced industries and

manufacturing facilities. This is particularly true in the automotive industry where the

losses due to misunderstanding and/or underestimation of the discussed issue result in

the losses of tens of millions of dollars. In the author’s experience, the hidden loses are

even greater than that as a number of premature failures of powertrain components (ﬁrst

of all, transmissions) are due to the burr and chips left after machining.

Manufacturers have been reluctant to hold the materials suppliers to a narrow range

of chemical composition and hardness variation in the materials supplied. The varia-

tions in the chemical composition and hardness make it very difﬁcult to specify the

optimum tool geometry, suitable grade of the tool material and optimal machining regime.

Moreover, it is next to impossible to implement the results of tribological studies under

these conditions as the modeled and experimentally obtained data cannot be relevant

for the whole range of properties of the work materials. Some common causes for poor

material speciﬁcation in manufacturing and research practices are:

• Prime cause is the lack of knowledge and readily available data on the correlation

of the properties (both mechanical and metallurgical) of work material and their

machinability (not to mention the metal cutting tribology).

• False perception that tighter speciﬁcations and control of metallurgical proper-

ties would always result in higher cost of blanks (casting, forgings etc.) and

materials. There are two major misunderstandings: (a) often, the tightened spec-

iﬁcation reduces the usage of some very expensive materials, for example in gray

Improvements of Tribological Conditions 381

cast iron; requirements to increase the hardness of gray cast iron lead to reduced

annealing time that, in turn, reduces the energy (natural gas or electricity) spent,

(b) the automotive industry rapidly shifts from the consideration of the cost of indi-

vidual components (blanks, tool, parts, etc.) to that of the cost per unit including

reliability of the processes and the ﬁnal products. In such a context, even if the cost

of blanks or raw materials grows, the overall saving on the much higher tool and

process costs, process stability, chip disposal, better quality of the machined part

and assembly structures would be many hundred times than this increased cost.

• Many automotive companies have developed standards on material speciﬁcation

more than 20 years ago, so these standards do not reﬂect the advances made in the

materials production and control. To change any particular speciﬁcation is a cumber-

some process that requires persistence and consumes a lot of time of the originator.

Moreover, since the automotive industry has outsourced the tooling management,

this originator – the tooling application specialist – is an outsider that makes this

task even more difﬁcult.

When it comes to research and development, the following causes can be listed:

• Prime cause is the lack of knowledge on the relevant properties of the work material

relevant to metal cutting in general and to the metal cutting tribology in particular.

The known books and articles on machinability of materials (for example [5,85])

are of a little help in such understanding.

• High cost of metallographic equipment. Many universities and research laborato-

ries cannot afford to have the equipment needed. Unfortunately, in the application

of equipment grants for metal cutting studies, metallographic equipment is not

considered as needed by many grant committees.

• Difﬁculties in altering the metallurgical and mechanical properties of the work

material.

The objective of this section is to present two simple examples showing that the machin-

ability of the work material and thus the tribological conditions at the interfaces and, as a

result, tool wear (tool life) can be signiﬁcantly improved via controlling the metallurgical

state of the work material.

6.3.1 Heat treatment

Heat treatment regime of many work materials alters their microstructural characteristics

and physical properties so it affects tool wear and thus the tribological process at the

tribological interfaces. Among many pre-process heat treatment used today, isothermal

annealing is the most widely used process. The practice of annealing which results in

transformation at a constant temperature is referred to as isothermal annealing in con-

trast to continuous cooling annealing, where transformation to ferrite and pearlite takes

place over wide range of temperatures. This also produces full annealing structure for

the purpose of: (a) inducing softness, (b) producing deﬁnite microstructure, (c) alter-

ing mechanical properties and (d) removing stresses. Although it is also believed that

382 Tribology of Metal Cutting

isothermal annealing improves the machinability of materials [86], the known literature

sources did not provide any speciﬁc quantitative information on improved machinability

due to annealing.

To clarify the issue and to show the inﬂuence of the regime of isothermal annealing

on tool wear, Talantov conducted a series of machining tests [87]. Forgings made of

AISI steel 4320 were subjected to 10 different industry regimes of isothermal annealing.

The forgings selected for the control batch were normalized at a standard temperature

of 935

◦

C. The tests were carried out using accelerated cutting regime: cutting speed

ν = 1.5 m/s, cutting feed f = 0.815 mm/rev. Tool material was P20 (75% WC, 5% TiC,

10% Co), tool geometry was: rake angle γ

= 0

◦

, major ﬂank angle α

= 10

◦

, minor

ﬂank angle α

= 10

◦

, tool cutting edge angle κ

= 45

◦

, tool minor cutting edge angle

= 45

◦

and inclination angle λ

= 0

◦

. The tool life criteria was the average width of

the ﬂank wear VB

= 0.4 mm.

The experimental results showed that tool life differs signiﬁcantly depending upon a

particular regime of isothermal annealing. Two regimes of this annealing were selected

for further detailed study: AN1 which results in the poorest tool life and AN2 which

results in the greatest tool life. The results obtained for the normalized work material (N)

were used as a reference as this heat treatment process is widely used in practice as a pre-

machining heat treatment. Figure 6.18 shows wear curves for the said heat treatments.

As seen, the difference is so signiﬁcant that it cannot be just ignored in the research and

practice of metal cutting.

The next step was to determine the root cause of the discussed difference in tool wear.

It was suggested that this cause lays in the differences in the deformation properties

of the two solid phases found in the work material used, namely, ferrite and pearlite.

To verify this suggestion, the partially formed chips obtained using a quick-stop device

were subjected to a metallographic study. These chips were sectioned from the rest of the

workpiece, mounted, ground, polished and etched to obtain specimens that can then be

(min)

0.1

0.2

0.3

0.4

AN1 N AN2

Fig. 6.18. Wear curves.

Improvements of Tribological Conditions 383

0.1

0.2

0.3

0.4

0.5

HV−p

HV−f

N AN2 AN1

Fig. 6.19. Hardening coefﬁcients for ferrite (k

HV−f

) and pearlite (k

HV−p

) solid phases.

viewed microscopically. The difference in the deformation properties of the phases was

assessed by the level of strain hardening of the phases. To do this, the microhardness

(load 0.05 N) of the ferrite and pearlite phases was measured in the above-mentioned

specimens. Using the average microhardness data, the hardening coefﬁcient k

was

determined for ferrite (k

HV−f

) and pearlite (k

HV−p

) phases as

= HV

−

, (6.48)

where HV and HV

are the microhardnesses of a phase before and after machining,

respectively.

The results shown in Fig. 6.19 suggest that the reason for good machinability achieved

in the annealing AN2 is approximately equal strain hardening of the two discussed solid

phases in the workpiece structure. A further metallographic study showed that the reason

for reduced wear is the increase in the actual contact area at the tool–chip and tool–

workpiece interfaces as in the case of AN2 as both the solid phases bear the contact load

so the contact stresses become much lower.

6.3.2 Inﬂuence of grain size

Another way to reduce the severity of the contact conditions at the tool–chip and tool–

workpiece interfaces is to create a ﬁne-grain structure with uniform distribution of small

ﬂakes of hard cementite (Fe

C). A perception was that smaller and evenly distributed

hard inclusions should reduce the severity of the contact conditions at the tool–chip and

tool–workpiece interfaces. To verify it experimentally, Talantov [87] carried out a series

384 Tribology of Metal Cutting

of cutting tests. The following conditions were used in the tests: work material – three

steel grades, namely, AISI 1045, 1020 and 1070. Tool material – P20 (85% WC, 5% TiC,

10% Co). Cutting feed f = 0.4 mm/rev, depth of cut d

= 1.5 mm. Tool geometry: rake

angle γ

= 0

◦

, major ﬂank angle α

= 10

◦

, minor ﬂank angle α

= 10

◦

, tool cutting

edge angle κ

= 45

◦

, tool minor cutting edge angle κ

= 45

◦

and inclination angle

= 0

◦

. The tool life criteria was the average width of the ﬂank wear VB

= 0.4 mm.

The ﬁrst series of tests included a comparison of tool lives for the two batches of work-

pieces made of AISI steel 1045 having different grain size. The ﬁrst batch was heat treated

so that it had the ﬂaked pearlite metallurgical structure while the second – tempered sor-

bite. To make a relevant comparison, both the structures were brought to the same

hardness HB 1800 MPa.

The microhardness of the chip contact area was measured at the middle of the tool–chip

contact length (the end of the zone of plastic contact). To do that, the specimens of the

partially formed chip were produced using a quick-stop device during actual machining

at a cutting speed of 1 m/s. These chips were sectioned from the rest of the specimens,

mounted, ground, polished and then subjected to microhardness tests. The test results

revealed that microhardness of the chip contact layer was HV = 5800 MPa for the

ﬂaked pearlite metallurgical structure, while it was HV = 4800 MPa for the micrograin

tempered sorbite structure.

To explain the phenomenon, the chip contact layers were studied using scanning electron

microscopy. It was found that there is almost no ferrite at the chip contact surface for the

ﬂaked pearlite metallurgical structure while both the solid phases (ferrite and pearlite) in

approximately equal proportion were found for the micrograin tempered sorbite structure.

The result obtained is readily explained by the fact that a micrograin structure prevented

the independent deformation of the solid phases and thus squeezing the pearlite phase

into the contact surface. Moreover, cementate was distributed much more evenly in this

structure. As a result, the hardness of the contact layer is smaller for this case.

As anticipated, the revealed difference in the microhardness of the chip contact layer

affect other tribological characteristics of the cutting process:

• For the same cutting speed (ν = 1 m/s), the cutting temperature was measured to be

940

◦

C in machining of the work material having the ﬂaked pearlite metallurgical

structure, while this temperature was only 895

◦

C for the work material having

micrograin tempered sorbite structure.

• Tool life was found to be 1.42–2.10-fold greater for the work material having

micrograin tempered sorbite structure, as shown in Fig. 6.20.

• Cutting force was found to be 25–30% smaller (depending on the cutting speed and

feed) for the work material having micrograin tempered sorbite structure.

The tests with AISI steels 1020 and 1070 did not show the same dependence. It was

explained by the fact that both the steels have only one predominant solid phase, namely

1020 – ferrite, 1070 – pearlite so that the tribological conditions in machining these

steels do not depend on their grain size.

Improvements of Tribological Conditions 385

22001800

HB (MPa)

120

T (min)

n=0.83m/s

n=1.70m/s

Fig. 6.20. Inﬂuence of hardness of AISI steel 1045 on tool life for the workpiece having

(1) tempered sorbite and (2) ﬂaked pearlite structures, at two different speeds.

6.3.3 Inﬂuence of minor elements in work materials

Normally, the inﬂuence of alloy additions on mechanical and physical properties, cor-

rosion and chemical behavior, and processing and manufacturing characteristics are

considered by mechanical and physical metallurgy. Normally, the coverage considers

“alloying” to include any addition of an element or compound that interacts with a base

metal to inﬂuence the properties. The prime consideration addresses the beneﬁcial effects

of major alloy additions, inoculants, dopants, grain reﬁners and other elements that have

been deliberately added to improve the performance of alloys [88]. Although the detri-

mental effects of minor elements or residual (tramp) elements included in the charge

materials or that result from improper melting or reﬁning techniques are discussed, such

discussions are of qualitative nature and primarily relates to the mechanical and phys-

ical properties. It is a common practice in the automotive industry to add some minor

alloying elements and elements traces to improve the castability, formability, heat treat-

ment properties or some minor working properties of the parts concerned, for example

crankshafts.

Unfortunately, possible changes in the machinability of various modiﬁed work materials

(castings, forgings, bar stocks etc.) are not considered as important. This is because,

there are very few studies that relate the inﬂuence of minor alloying components and

components in traces on machinability so many tool manufacturing processes and quality

specialists are not aware about the strength of the correlation discussed. As a result, wide

range of variations of these components is allowed by common speciﬁcation of materials

used in the automotive industry. When the allowed range is changed or minor component

is added to the speciﬁcation of a material by part designers, the manufacturing and process

specialists may not be informed about this change.

386 Tribology of Metal Cutting

A good example of the inﬂuence of minor components on the machinability of gray

cast iron, widely used in the automotive industry, was presented by Grifﬁn and

co-workers [89]. They found that within the range allowed for manganese (0.3–0.8%),

the longest tool life corresponds to 0.3% and it reduces by more than twice when the

content of manganese is 0.8%. The same result was obtained for the allowable range of

tin. Even more pronounced inﬂuence of volume percent of hard inclusion was found in

this study. Unfortunately, the conclusive results obtained did not affect the methodology

used by the automotive industry in specifying composition of gray, ductile and mal-

leable cast irons although signiﬁcant cost saving can be achieved with minimum efforts.

Moreover, quality and reliability of cars can be improved by the reduction in the number

of defective transmissions and engine due to metallic chips and burrs left in weans and

gates of case body, pump cover, upper and lower valve bodies, etc.

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