Astakhov V. Tribology of Metal Cutting

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Improvements of Tribological Conditions 337

Other suggested ways of cutting tool penetration are even more illusive. Access through

voids connected with built-up edge is not very likely because the built-up edge does

not exist at the range of cutting speeds used today. Access into the gap created by tool

vibration is not likely because the amplitude of vibration is extremely small so the chip

does not disengage with the tool [31]. Propagation from the chip blackface through

distorted lattice structure cannot happen physically due to extremely short time of chip

deformation [32].

In the author’s opinion, the most conclusive results were obtained by a research team

headed by Talantov [33]. This study can be summarized as follows. Low-, medium- and

high-carbon steels were selected as work materials. A special setup where the dropping

cutter with separable cutting inserts was installed in a specially designed pendulum-type

device. The cutting insert was allowed to disengage “to freeze” the cutting insert –

partially formed chip at different cutting conditions. After the test, a “frozen” part con-

taining the partially formed chip in its contact with the cutting insert was cutoff from the

rest of the workpiece for further studies. A scanning electron microscope was used to

investigate the geometry of the tool–chip contact at 6000–10 000× magniﬁcation. Using

an X-ray microanalyzer, the nature of the tool–chip contact was studied.

The test results proved that there are two distinctive regions within tool–chip contact,

as discussed in Chapter 3. The material of the chip was found to ﬁll out even the

microscopic valleys and hills on the cutting insert contact surface at a level below the

surface roughness, and so no network of ﬁne capillaries was found. Moreover, the zone

of plastic contact was observed at the tool–chip and tool–workpiece interfaces. The

evidence of intensive diffusion was found within this zone that completely excludes

even the ﬁnest capillaries.

To study the possibility of cutting ﬂuid penetration at contact surfaces objectively, a series

of tests were carried out in which the various cutting ﬂuids with added radioisotope T-3

(tritium H3) were used. Test conditions were as follows: work material – AISI steel

1045; tool material – standard carbide M10; tool geometry – normal rake angle 0

◦

normal ﬂank angle 10

◦

, cutting edge angle 45

◦

, cutting edge inclination angle 0

◦

; cutting

speed range 30–140 m/min; feed range 0.12–0.45 mm/rev; depth of cut 2 mm. The tests

included 5 min of cutting at intensive cutting regimes and quick process interruption with

a specially designed quick-stop device. Then, the tool contact surfaces were examined

using radiography.

The results of this experiment showed no traces of radioisotopes at the plastic part of the

tool–chip and tool–workpiece interfaces. Small traces of radioisotopes within the elastic

part of the tool–chip and tool–workpiece interfaces were found only on very few test

specimens.

The major result of this research program conclusively proves that the cutting ﬂuid does

not penetrate the contact surfaces at the discussed interfaces. All the known “look-nice

stories” about EP additives (chlorine, sulfur and fatty acids for example) that, according

to literature sources, are added to the cutting ﬂuid to prevent severity of the chip and

workpiece contacts where the chip has no ground. On the other hand, their use in metal

338 Tribology of Metal Cutting

cutting has proven to be successful in terms of improving the tool life, reducing the

cutting forces and improving the integrity of machined surfaces. Therefore, another

mechanism of their action should be suggested and veriﬁed if one tries to improve the

tribological conditions in metal cutting by using the cutting ﬂuids.

Cooling action. In terms of metal cutting tribology, the commonly referred actions

of the cutting ﬂuid are the lubrication and cooling. Being self-obvious, the former is

most widely tested using various standards (above discussed) and tailored tests while

the latter did not attract much attention besides the general perception that water-soluble

cutting ﬂuids possess higher cooling ability than waterless ones. Unfortunately, there

are no characteristics of the cooling action of cutting ﬂuid is readily available at tool

and process designers’ disposal although a number of studies are known (for example

[34,35]). Moreover, in the author’s opinion, the most important action of cutting ﬂuid is

the embrittlement action which did not attract much attention in the literature on metal

cutting and cutting ﬂuids.

As pointed out in the literature on metal cutting, most of the mechanical energy associated

with chip formation converts into the thermal energy [30,36]. For example, Epifanov and

Rebinder [38] conducted a series of tests using a specially designed setup to study the

energy balance in metal cutting. The setup included a calorimeter and a dynamometer

so the thermal energy generated in cutting was measured directly while the mechani-

cal energy spent in cutting was calculated as a product of the measured cutting force

and the cutting speed. Wide ranges of work materials and cutting regimes were used

in the study. The results of this study conclusively proved that 99% of the energy

is spent in cutting converts into the thermal energy while in standard metrical test-

ing (tensile, compression, impact) such a conversion was found to be in the range

of 0.75–0.93%.

Because practically all the mechanical energy associated with chip formation converts

into the thermal energy, the heat balance equation is of prime concern in metal cut-

ting studies [32]. Accounting for the energy ﬂows in the cutting system, discussed in

Section 6.2.2, this equation can be written as

= F

ν = Q

= Q

, (6.1)

where P

is the cutting power, F

is the power component of the cutting force, ν is the

cutting speed, Q

is the total thermal energy generated in the cutting process, Q

the thermal energy transported by the chip, Q

is the thermal energy conducted into the

workpiece and Q

is the thermal energy conducted into the tool.

Table 6.1 presents the result of energy balance study in metal cutting [32] in machining

steel 1045 using experimentally obtained chip compression ratios. As seen in these exper-

iments, most of the thermal energy generated in the cutting process is carried away by the

moving chip. At cutting speeds used today, this portion has reached 80–85%. Approx-

imately, equal portions of the total thermal energy generated in cutting are conducted

into the workpiece and into the tool. As follows from the data presented in Table 6.1, the

ratio of these portions varies signiﬁcantly with the cutting speed. As the speed increases,

greater portion of the thermal energy is conducted into the tool.

Improvements of Tribological Conditions 339

Table 6.1. Energy balance in machining (steel 1045).

(

m/s

)

Pe Q

(

J/s

)

(

J/s

)

(

J/s

)

(

J/s

)

(

)(

)

0.10 1.76 47.9 50.2 38.4 40.2 9.2 9.695.5

0.20 3.52 93.7 55.7 63.7 37.8 11.0 66.6 168.4

0.50 8.80 272.3 70.3 100.3 25.9 14.7 3.8 287.3

1.00 17.60 501.6 76.2 136.9 20.8 19.7 3.0 658.3

2.00 35.20 1177.1 82.8 217.5 15.3 27.0 1.0 1421.6

4.00 70.40 2306.2 86.3 336.7 12.6 29.4 1.1 2572.3

Although most of the thermal energy generated in the cutting process is carried away by

the moving chip, it must, however, not be concluded that the chip temperature is higher

than the tool temperature. In reality, it is much lower. This is simply because the chip

moves and thus mass transfer takes place so that any elementary volume of the moving

chip is not exposed to high contact temperatures for sufﬁcient time to increase the chip

temperature signiﬁcantly. The same can be said about the workpiece as the tool moves

over its surface, “spreading” the generated thermal energy over all machined surface.

On the contrary, the tool contact surfaces at the tool–chip and tool–workpiece interfaces

do not move. As a result, the tool contact temperatures are normally much higher than

those of the chip and the workpiece. The matter worsens because new superhard tool

materials (CBN, cermets, ceramics, etc.) have very low thermal conductivity, so the tool

cannot dissipate the thermal energy generated at the interfaces fast enough, which results

in extremely high temperatures at the tool contact surfaces.

The observations of the cooling action of a cutting ﬂuid allow the following conclusions:

• Slightly reduces the cutting temperature at cutting speeds up to 150 m/min. At higher

cutting speed, it just stabilizes the temperature of the workpiece. As discussed in

Chapter 4, there is an optimal cutting temperature that correlates the minimum tool

wear. Therefore, the cooling action of a cutting ﬂuid is helpful in terms of tool life

if it brings the cutting temperature closer to the optimal cutting temperature. Other-

wise, tool life is not improved but rather reduced in many conventional machining

operations as found by Seah et al. [18]. Unfortunately, this important issue has

never been explained and thus addressed in the published materials on cutting ﬂuids.

Although the reduction of the cutting temperature due to the cooling action of the

cutting ﬂuid appears to be small [39] and thus it might appear that the cutting ﬂuid

has little effect on the tool–chip interface tribological process. In reality, however,

it is not quite true. Experiments showed that if the cutting temperature exceeds the

optimal temperature at the tool–chip interface and diffusion wear takes place there,

a 10–15% decrease in the contact temperature leads to a 2–3-fold reduction in the

coefﬁcient of diffusion between the tool and the work materials.

• Improves accuracy of machining. By reducing the temperature of the tool and

workpiece, better accuracy of machining can be achieved as a result of reduced

temperature deformation of the components of the machining system.

340 Tribology of Metal Cutting

• Reduces the radius of curvature of the chip. The lower the thermal conductivity of

the work materials, the greater the reduction. This aspect should be accounted for

in the design of obstruction-type chip breakers, where this radius is an important

design consideration.

• Reduces the tool–chip contact length. This is a negative aspect because the cutting

forces are not normally reduced. As a result, higher contact stresses act at the tool–

chip interface that may result in a lower tool life. This is particularly important for

brittle tool materials, which may experience the chipping of the cutting edge.

• Increases thermal shock in interrupted cuts. For example, such as those occuring

in milling can cause thermal cracking on the cutting wedge. Once started, the crack

will grow causing either a catastrophic failure of the cutting insert or a loss of

dimensional accuracy and surface ﬁnish on the parts.

Direct cooling action. The cooling action of the cutting ﬂuid is due to forced

convection. As such the heat transfer is given by

Q = A



−θ



, (6.2)

where A

is the heat transfer area of surface, h

is the convection heat transfer coef-

ﬁcient of the process, θ

and θ

are the temperatures of the surface and cutting ﬂuid,

respectively.

Although convective heat transfer problems can seem incredibly confusing given the

multitude of different equations available for different systems and ﬂow regimes, one

should keep in mind that the whole goal of the problem is to ﬁnd the overall heat transfer

coefﬁcient, h

so the problem of heat transfer from the object in the ﬂuid medium can

be solved using Eq. (6.2). Because this coefﬁcient depends on many physical and design

parameters, its determination requires introduction of some similarity numbers as:

The Reynolds Number (Re)

Re =

, (6.3)

where ν

is the velocity of the cutting ﬂuid, b

is the equivalent length (diameter or

other length-scale) and ν

is the kinematic viscosity of the cutting ﬂuid.

For forced convection problems Re determines whether or not the ﬂow is turbulent or

laminar. For ﬂow past ﬂat plates, the transition region from laminar to turbulent is about

< Re < 10

The Nusselt Number (Nu)

Nu =

, (6.4)

where k

is the thermal conductivity of the cutting ﬂuid.

Improvements of Tribological Conditions 341

The Nusselt Number is deﬁned as the ratio of convection heat transfer to ﬂuid conduction

heat transfer under the same conditions.

The Prandtl number (Pr) characterizes thermal properties of the cutting ﬂuid

Pr =

, (6.5)

where α

is the thermal diffusivity of the cutting ﬂuid. It is related to the following

characteristics of the cutting ﬂuid: thermal conductivity k

, speciﬁc heat C

P−cf

, density

P−cf

(6.6)

To describe heat transfer, the following generic relationship is used

Nu = c

, (6.7)

where c

is a constant.

The particular values of c

, n and m depend on the relative location, the cutting ﬂuid

ﬂow and the cooled surface [40,41].

The most general case is the interaction of the cutting ﬂuid ﬂow with a ﬂat surface set

perpendicular to this ﬂow. In this particular case

Nu = 0.20Re

0.65

0.33

(6.8)

or substituting Eqs. (6.3)–(6.6) into Eq. (6.8), recalling that ρ

= γ

/g, and expressing

, one can obtain

0.20

0.35

0.33

0.65

0.67

P−cf

)

0.33

0.32

, (6.9)

where γ

is the speciﬁc weight of the cutting ﬂuid and g is the acceleration due to

gravity.

Although Eq. (6.9) has a clear structure and thus can be analyzed directly, much more

general results can be obtained if one realizes that its ﬁrst term is an application-speciﬁc

constant. Therefore, we can introduce a new parameter called the cooling intensity K

which includes only the second cooling-process-speciﬁc term of Eq. (6.9),

= ν

0.65

0.67

0.33

P−cf

0.33

−0.32

(6.10)

As seen, the velocity of the cutting ﬂuid ν

affects its cooling ability almost as much as

its thermal conductivity k

and much more than its speciﬁc heat C

P−cf

. Table 6.2 shows

342 Tribology of Metal Cutting

Table 6.2. Cooling intensity of different media.

Cutting ﬂuid Thermal

conductivity

)



kcal

m·h·

◦



Speciﬁc heat

P−cf

)



kcal

kg·

◦



Kinematic

viscosity

(ν

×10

)





Speciﬁc

weight

(γ

)



kg/m



Cooling intensity

In%of

for water

Water – normal application (ﬂooding from the

chip free surface at ν

= 0.2m/s and

temperature θ

= 20

◦

0.515 0.999 3.60 998.20 2715 100.0

Soluble oil of 10% concentration – normal

application

0.297 0.884 5.16 990.80 1646 60.6

Kerosene – normal application 0.126 0.448 7.92 800.00 589 21.7

Carbon tetrachloride (CCl

) – normal application 0.087 0.127 2.19 1594.90 531 19.6

Pure gear box oil – normal application 0.109 0.436 162.00 875.60 211 7.8

Cutting strait oil with chlorine and sulfur –

normal application

0.103 0.240 337.00 900.10 155 5.7

Compressed air at velocity 100 m/s and

temperature θ

= 20

◦

0.0223 0.240 54.31.205 479 17.6

Compressed air at velocity 100 m/s and

temperature θ

=−40

◦

0.0182 0.242 36.01.515 515 19.0

Compressed air at velocity 300 m/s and

temperature θ

= 20

◦

0.0223 0.240 54.31.205 970 36.7

Compressed air at velocity 300 m/s and

temperature θ

=−40

◦

0.0182 0.242 36.01.515 1046 38.6

Compressed air at velocity 500 m/s and

temperature θ

= 20

◦

0.0223 0.240 54.31.205 1360 50.1

Compressed air at velocity 500 m/s and

temperature θ

=−40

◦

0.0182 0.242 36.01.515 1465 54.0

Improvements of Tribological Conditions 343

the results of calculations of the cooling intensity K

for some common cases in metal

cutting.

The summary of the results on the cooling action of the cutting ﬂuid allows pointing out

the following:

• The cooling action of the cutting ﬂuid assists the cutting process as far as it brings the

cutting temperature closer to the optimal cutting temperature. Otherwise, the use of

cutting ﬂuid reduces tool life [42]. When one considers the whole machining system,

the cooling action often stabilizes the temperature of the machining system compo-

nents that results in better machining accuracy due to reduced thermal deformations

of these components. Another down-to-earth but important aspect is the tempera-

ture of the part after machining. Often, when no coolant is used, this temperature

is high so the operator ﬁnds it difﬁcult to handle the part (including unloading,

measurements and/or gaging).

• Signiﬁcant intensiﬁcation of the cooling action of the cutting ﬂuid is achieved by

increasing its velocity, which has a dual result. First, it increases the convection

heat transfer coefﬁcient and second, high-velocity jets of the cutting ﬂuid blow a

boundary layer formed on high-temperature surfaces. This explains the efﬁciency

of the high-pressure cutting ﬂuid supply which increases the velocity of the cut-

ting ﬂuid. Although it is often claimed that the high-pressure supply of the cutting

ﬂuid increases its penetration ability into the tool–chip and tool–workpiece inter-

faces [43,44], in reality it is not so because, as it is shown in Chapter 3, the contact

stresses are a way higher than the maximum pressure of the cutting ﬂuid. It was

conclusively proven that tool life (tool wear) does not signiﬁcantly increase in cut-

ting under very high static pressure [45]. In reality, high pressure of the cutting ﬂuid

increases its velocity, which, in turn, signiﬁcantly improves the cooling action of

this ﬂuid. Moreover, it does not affect the cutting forces [46].

Cooling action due to evaporation. To comprehend the cooling action of the cutting

ﬂuid when it boils due to the contact with hot surfaces of the tool, one should consider the

process of heat exchange between a solid and a boiling liquid [47]. The boiling process

includes formation, growth and separation of bubbles.

When the surface temperature is slightly hotter than the saturation temperature of the

liquid, the excess vapor pressure is unlikely to produce bubbles. The locally warmed

liquid expands and convection currents carry it to the liquid–vapor interface where evap-

oration takes place and thermal equilibrium is restored. Thus, in this mode, evaporation

takes place at small temperature differences and with no bubble formation. As the sur-

face becomes hotter, nucleate boiling takes place. As such, the excess of vapor pressure

over local liquid pressure increases and eventually bubbles are formed. These occur at

nucleating points on the hot surface where miniature gas pockets, existing in surface

defects, form the nucleus for the formation of a bubble.

As soon as a bubble is formed, it expands rapidly as the warmed liquid evaporates into

it. The temperature of vapor in the bubbles is equal to the saturation temperature (θ

344 Tribology of Metal Cutting

Table 6.3. Saturation temperature (θ

), heat of vaporization (r

), density of vapor

(ρ

), equivalent length (diameter or other length-scale) (b

) for water vapor as

functions of pressure (p

(

MPa

)

(

◦

)

×10

−3



J/kg





kg/m



×10

(

)

0.01 45.8 2392 0.07 3836

0.02 60.1 2357 0.13 954.6

0.04 75.9 2318 0.25 289.6

0.06 85.9 2293 0.36 137.4

0.08 93.5 2273 0.48 79.4

0.10 99.6 2257 0.59 50.6

0.20 120.2 2202 1.13 14.2

This temperature depends on the properties of the liquid and pressure (Table 6.3), which

remains invariable during the boiling process. For water vapor under the standard atmo-

spheric pressure, this temperature is θ

= 100

◦

C. The buoyancy detaches the bubble

from the surface and another starts to form. Nucleate boiling is characterized by vigorous

bubble formation and turbulence. Exceptionally high heat transfer rates and heat trans-

fer coefﬁcients with moderate temperature differences occur in nucleate boiling, and in

practical applications. Boiling is nearly always in this mode.

The temperature of the boiling liquid in contact with the hot surface is equal to the

temperature of this surface, θ

and then reduces with distance from this surface within

the limits of the so-called boundary layer to a temperature slightly higher than θ

. The

pressure of the vapor within a bubble is higher than that of the surrounding. Due to this

fact, this bubble grows to a certain size and then detaches from the surface. The diameter

of the bubble, d

at the instant of detaching is

= 0.2ψ



/(ρ

−ρ

), (6.11)

where σ

is the surface tension, ρ

and ρ

are the densities of the liquid and vapor,

respectively, ψ

is the contact angle (Fig. 6.4), for kerosene ψ

= 26

◦

, for water

= 50

◦

and for mercury ψ

= 137

◦

Fig. 6.4. Contact angle.

Improvements of Tribological Conditions 345

The contact angle (ψ

) is a quantitative measure of the wetting of a solid by a liquid. It is

deﬁned geometrically as the angle formed by a liquid at the three phase boundary where

a liquid, gas and a solid intersect, as shown in Fig. 6.4. Low values of ψ

indicate that

the liquid spreads, or wets well, while high values indicate poor wetting. If the contact

angle is less than 90

◦

the liquid is said to wet the solid. If it is greater than 90

◦

it is

said to be non-wetting. A 0

◦

contact angle represents complete wetting. Two different

approaches are commonly used to measure the contact angles of non-porous solids, i.e.

goniometry and tensiometry. Goniometry involves the observation of a sessile drop of

test liquid on a solid substrate. Tensiometry involves measuring the forces of interaction

as a solid is in contact with a test liquid.

When the wetting is good, the bubbles formed detach easily, whereas when the wetting

is poor, the bubbles formed spread over the surface forming a vapor ﬁlm. The greater

the wetting, greater the ease with which the bubbles formed leave the surface frequently,

higher is the heat transfer coefﬁcient of this process, h

. The frequency of separation

from the surface and the number of bubbles formed as well as h

depend on the temper-

ature difference ∆θ

= θ

−θ

. Figure 6.5 shows the dependence of the heat transfer

coefﬁcient on the boiling of water on the temperature difference ∆θ

under the atmo-

spheric pressure of 0.1 MPa. When ∆θ

is small, the heat transfer process is the same as

that due to convection (Zone 1 in Fig. 6.5). Then the heat transfer coefﬁcient increases

dramatically further increase in ∆θ

(Zone 2 in Fig. 6.5) because the frequency of bubble

formation and the number of bubbles formed increase. The boiling regime corresponding

to this state is called convective boiling.

As the heating surface temperature rises, the rate of production of vapor bubbles becomes

so high that eventually the surface becomes enveloped in a blanket or ﬁlm of vapor which

prevents the liquid from wetting the surface. When this happens, the insulating effect

0.1 1.0 10 100

∆q

(°C)

(W/(m

°C)

1234

Fig. 6.5. Dependence of the heat transfer coefﬁcient of this process (h

) on the boiling of water

on the temperature difference ∆θ

= θ

−θ

346 Tribology of Metal Cutting

of the ﬁlm greatly reduces the rate of heat transfer. As seen in Fig. 6.5 (Zone 3), after

reaching a certain maximum (for water) h

max

= 4.65 ×10

W/m

2 ◦

C under the critical

temperature range ∆θ

= 20−25

◦

C, the heat transfer coefﬁcient reduces when ∆θ

increases further. Convective boiling becomes ﬁlm boiling. Zone 4 (Fig. 6.5) corresponds

to the stable ﬁlm boiling. Here, the heat transfer coefﬁcient shows practically no depen-

dence on ∆θ

. The foregoing analysis suggests that the highest cooling effect is reached

in Zone 3 so the design of the cooling system should be optimized to assure the operation

in this zone.

The maximum heat transfer occurs when h

= h

max

and ∆θ

= ∆θ

, so the maximum

speciﬁc heat ﬂux can be written as q

b−max

= h

max

∆θ

. Because h

max

and ∆θ

depend

on the physical properties of the cutting ﬂuid so that they differ for different cutting

ﬂuids. For a wide range of water-based cutting ﬂuids, (h

max

= 4.65 × 10

W/m

2 ◦

and ∆θ

= 20−25

◦

C), the maximum speciﬁc heat ﬂux can be written as q

b−max

1.16 × 10

W/m

As in the previous section, heat transfer parameters can be calculated using Eq. (6.7).

The similarity numbers, however, should be considered as follows:

• The thermal characteristic of the ﬂuid in the Nusselt Number (Eq. (6.4)) should be

taken at temperature θ

• Parameters b

and ν

in the Reynolds number (Eq. (6.3)) should be interpreted

as follows: the equivalent length (diameter or other length-scale) b

is selected

proportional to the diameter of the bubble, d

formed on boiling; ν

= ν

is the

conditional velocity of vapor, where ν

= q

ν−f

(

)

, where q

ν−f

is the speciﬁc

heat ﬂux into the liquid phase



W/m



, r

is the heat of vaporization



J/kg



, and

is the density of vapor at θ

Using the parameters introduced in Eq. (6.7), one can obtain

= c



∆θ



(6.12)

Coefﬁcient c

, and powers p

and n

in Eq. (6.12) are shown in Table 6.4.

To calculate heat transfer in the most practical case, where a water-based cutting ﬂuid

is supplied into the machining zone at the atmospheric pressure, Eq. (6.12) can be

signiﬁcantly simpliﬁed using the data presented in Tables 6.3 and 6.5 keeping in mind

Table 6.4. Coefﬁcient c

, and powers p

, m

and n

in Eq. (6.12).

Re c

≤ 0.01 0.00390 2.00 1.00 0.66

> 0.01 0.00263 2.86 1.86 0.95