Astakhov V. Tribology of Metal Cutting

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Cutting Tool Wear, Tool Life and Cutting Tool Physical Resource 257

0.2

f (mm/rev) f (mm/rev)

0.03

0.3 0.4 0.5

0.05

0.07

0.09

600

700

800

900

q(°)

0.07

0.05

0.03

0.01

0.1 0.2 0.3

=1.00 mm

=2.00 mm

=0.50 mm

(µm/10

)

Fig. 4.24. Inﬂuence of cutting feed in turning stainless steel AISI 303: (a) on cutting temperature,

tool wear rate and h

criterion, tool material – carbide M20 (92% WC, 8% Co), ν = 72 m/min,

= 5 mm, γ = 15

◦

, α = α

= 10

◦

, κ

= 45

◦

, κ

= 10

◦

, r

= 1 mm, (b) on h

criterion under

the optimal cutting temperature θ

opt

= 850

◦

C, tool material – carbide P10 (14% TiC, 8% Co),

γ = 10

◦

, α = α

= 10

◦

, κ

= 45

◦

, κ

= 10

◦

, r

= 1 mm.

t (min)

(b)

t = 0s t = 10s t = 30s t = 60s

468100

40 80

Tool breakage

(a)

, h

(µm)

Fig. 4.25. Plastic lowering of the cutting edge. Turning, work material – steel AISI 1045, tool

material – carbide P20 (77% WC, 15% TiC, 8% Co), ν = 180 m/min, d

= 2 mm, f = 0.3 mm/rev:

(a) dynamics of the lowering and (b) perimeters of lowering as functions of time.

As explained [16], the stress relaxation has temperature-diffusion nature. Therefore, it

is reasonable to assume that plastic deformation in high-temperature creep takes place

along the isotherms of a quasi-state temperature ﬁeld.

Talantov presented another experimental evidence to support the suggestion that the plas-

tic lowering of the cutting edge is related to high-temperature creep [24]. He constructed

258 Tribology of Metal Cutting

325 655 640 525 460

410°C

585

300 610 405 430

300°

575

615°C

530

485°C

(a) (b)

= 0.77mm

= 1.95mm

= 1.10mm

= 0.50mm

Fig. 4.26. Temperature ﬁeld in the cutting wedge in turning. Work material – steel AISI 1045,

ν = 1 m/s, f = 0.3 mm/rev, d

= 2 mm, for two groups of tool materials: (a) carbide M20

(92% WC, 8% Co and (b) carbide P20 (77% WC, 15% TiC, 8% Co).

temperature ﬁelds for two basic groups of carbide tool materials experimentally as shown

in Figs. 4.26(a) and (b). The length of the plastic part l

c−p

of the tool–chip contact length

was measured for each tool and new tools were made with the restricted rake face

equal to l

c−p

. The experimental comparison of the plastic lowering of the cutting edge

of the tool with the natural and restricted contact lengths showed that this lowering is

much smaller for the latter tools. This is because much less heat is generated due to

the friction at the tool–chip interface so that a smaller thermal energy enters the tool.

By excluding this elastic part, the amount of the thermal energy entered into the cutting

tool substantially reduces the plastic lowering of the cutting edge. This explains the major

advantages of the tool with the restricted contact length known in practice.

It is also known [19,24] that signiﬁcant intensiﬁcation of high-temperature creep takes

place if the amplitude of cyclic load increases. As discussed in Chapter 1, the amplitude

of the cyclic cutting force grows when seizure occurs at the tool–chip interface (Fig. 1.25)

that results in the saw-toothed continuous fragmentary chip (Figs. 1.30–1.33). Therefore,

a noticeable increase in the plastic lowering of the cutting edge should be observed under

this condition. Figure 4.27 shows parameters characterizing the plastic lowering of the

cutting edge under the high-amplitude cutting force during the formation of this chip

type. Signiﬁcant intensiﬁcation of the plastic lowering of the cutting edge in this case

is readily observed if one compares Figs. 4.27 and 4.25(b). It explains the well-known

fact: tool life reduces signiﬁcantly when the conditions at the tool–chip interface result

in the occurrence of the saw-toothed continuous fragmentary chip.

A comparison of the model of plastic lowering (shown in Figs. 4.23(a) and 4.25(a))

and the data shown in Fig. 4.27 shows that the model of plastic lowering is more com-

plicated than described so far. As shown, for example, h

ﬁrst increases with increasing

Cutting Tool Wear, Tool Life and Cutting Tool Physical Resource 259

120

160

46810

, h

(µm)

t (min)

Fig. 4.27. Intensiﬁcation of the parameters of the plastic lowering of the cutting edge under the

condition of the increased cyclic loading (seizure). Turning of modiﬁed AK Steel (AISI Austenitic

Stainless Steel 305), ν = 1 m/s, f = 0.3 mm/rev, d

= 2 mm and carbide M20 (92% WC, 8% Co).

and then, reaching a certain value, does not change any further. Talantov [24] sug-

gested that this is because the so-called wear land (of width h

in Fig. 4.23(a)) increases

not only due to the regular wear of the tool ﬂank but also due to bulk deformation of the

cutting wedge.

To verify this suggestion, a special turning cutting test was carried out where the

actual width h

f −a

of the wear land was measured and the calculated wear land h

f −c

as determined through the radial tool wear h

and normal ﬂank angle is,

f −c

= h

/tan α

. (4.28)

The following machining regime was used: workpiece material – AISI steel 1045, cutting

feed, f = 0.3mm/rev, depth of cut d

= 2 mm; cutting tool: tool materials – carbide

M20 (92% WC, 8% Co) and carbide P20 (77% WC, 15% TiC, 8% Co), tool geometry:

rake angle γ

= 0

◦

, ﬂank angles α of the major and minor cutting edges α

= α

= 10

◦

tool cutting edge angle of the major cutting edge κ = 45

◦

, tool cutting edge angle of the

minor cutting edge κ

= 15

◦

, tool nose radius r

= 0.25 mm.

The test results shown in Fig. 4.28 fully support the Talantov’s suggestion. As shown,

the actual width of the wear land h

f −a

is greater than the calculated (Eq. (4.28)) width of

this land, h

f −c

. This is because the actual width of the wear land increases not only due

to tool wear but also due to the creep of the cutting wedge. Analysis of the temperature

ﬁeld shown in Fig. 4.26(a) and the contact conditions at the tool–chip interface (discussed

in Chapter 3) allows offering the following explanations to the described phenomena.

The cutting wedge is subjected to complicated plastic deformation in the cutting process.

A model of this deformation is shown in Fig. 4.29. A part of the cutting wedge (Zone 2 in

Fig. 4.29) deformed and moved toward the tool ﬂank more intensively than its other parts

260 Tribology of Metal Cutting

812

f−a

, h

f−m

(mm)

t (min)

0.09

0.06

0.03

f−a

f−m

f−a

1045-M20

n=1.67m/s

1045-M20

n=1.33m/s

1045-P20

n=1.67m/s

Fig. 4.28. Comparison of the actual (experimentally obtained), h

f −a

and calculated, h

f −c

widths

of the wear land.

f−a

Zone 1 Zone 2 Zone 3

Fig. 4.29. A model of the plastic lowering of the cutting edge.

shown as Zones 1 and 2. Such a behavior is explained by the temperature-deformation

stress relaxation where the temperature is a decisive factor. Starting with a certain tem-

perature, the intensity of stress relaxation increases signiﬁcantly. The process of plastic

deformation where the intensity of stress growth is equal to that of stress relaxation due to

plastic deformation relates to creep. Analysis of the temperature ﬁelds of the cutting edge

(for example shown in Fig. 4.26) shows that there are some zones heated to signiﬁcant

temperature over the whole volume from the rake face to the tool ﬂank, as shown by the

corresponding isotherms. High-temperature creep takes place along these isotherms that

leads to the pushing out of some volumes of the tool material below the formed wear

Cutting Tool Wear, Tool Life and Cutting Tool Physical Resource 261

land on the tool ﬂank. As tool wears, high-temperature creep of the material in Zone 2

becomes a continuous process as the temperatures in this zone increase proportionally

to tool wear.

At the same time, any motion of the tool material within Zone 1 is impossible as this

zone is loaded by the compressive stresses from the rake and ﬂank contact surfaces.

Moreover, the temperatures in Zone 1 are lower than those in Zone 2. As a result, the

tool material in Zone 1 is not subjected to creep. As for Zone 3, high-temperature creep

does not occur here because of a number of reasons: lower temperatures, greater cross-

sectional areas and much lower contact stresses. As such, these temperatures in Zone 3

cause stress relaxation in this zone without its plastic deformation.

The description above allows us to conclude that some zones in the cutting wedge can be

subjected to plastic deformation under certain combination contact stresses/temperatures.

As such, this plastic deformation causes the growth of the wear land on the tool ﬂank

without actual increase in tool wear. Therefore, the notion of the wear land on the tool

ﬂank surface is conditional as this land forms not only due to tool wear but also due

to high-temperature creep (and thus plastic deformation) of the cutting wedge. This

conclusion is of particular importance for cutting tools with coatings.

For years Talantov [24] asked a simple question “How come a coating is able to reduce

tool wear of the tool ﬂank for a long time period? It is known that the coating thick-

ness (the thickness of the coated layer) typically ranges from 3 to 10 µm. If the cutting

wedge would be ideally rigid having the ﬂank angle α

= 10

◦

and the coating thick-

ness is 8 µm, then this coating should be gone when the width of the ﬂank wear land

reaches 40 µm. Therefore, the coating under these conditions should be wiped out dur-

ing the ﬁrst minute of machining. In practice, however, such a coating reduces the tool

wear rate during several dozens of minutes as the width of the ﬂank wear land reaches

0.4–0.7 mm. The known explanations of this fact rely on the existence of “an overhung”

of the coating material over the tool ﬂank although not one study can ﬁnd this overhung

in reality [24].

Talantov proposed and proved a much simpler and real mechanism to explain the dis-

cussed fact. Figure 4.30 shows the cutting wedge after 10 min of cutting. As shown,

the coating exists over the whole ﬂank wear land having a width of 0.4 mm. Therefore,

this is not just the wear land but rather it is the contact land formed due to the above-

described plastic lowering of the cutting wedge. To demonstrate the validity of such an

explanation, Talantov studied [24] the formation of the contact length in cutting tools

with TiN coating. Figure 4.31 shows the cutting wedge after 3 min of cutting. As shown,

the coating is still intact on the tool ﬂank contact surface (except for some random chip-

ping) while there is no coating on the rake face in the zone, where certain volumes of

the tool material were pushed out towards the tool ﬂank.

According to Talantov [24], the mechanism of high-temperature creep (the model shown

in Fig. 4.29) can be described as follows. In Zone 2 (Fig. 4.29), the ﬁst phase of the

carbide tool material to deform is the cobalt matrix. Due to the action of the contact

stresses, this phase deforms along the direction shown by arrows in Fig. 4.29. This leads

to the destruction of the cobalt matrix which holds carbide particles as a frame. To prove

262 Tribology of Metal Cutting

Fig. 4.30. Micrograph of the cutting wedge. Cutting conditions: work material – steel AISI 1070,

tool material – M10, coating – TiN, cutting speed ν = 90 m/min, feed f = 0.1 mm/rev, depth of

cut d

= 1.5 mm, cutting time τ = 10 min, magniﬁcation ×120.

Fig. 4.31. Micrograph of the cutting wedge. Cutting conditions: work material – steel AISI 1070,

tool material – P20, coating TiC, cutting speed ν = 100 m/min, feed f = 0.3 mm/rev, depth of

cut d

= 1.5 mm, cutting time τ = 3 min, magniﬁcation ×100.

Cutting Tool Wear, Tool Life and Cutting Tool Physical Resource 263

that this is the case, the worn tool wedge was mounted, polished and etched in a ferric

chloride solution in hydrochloric acid. The results of the described destruction can be

shown as a black region in Fig 4.31. As shown, the conﬁguration of this black region

resembles the model shown in Fig. 4.29. The destruction of the cobalt matrix as a frame

can clearly be shown by comparing the initial carbide structure shown in Fig. 4.32(a) and

that after 7 min of machining shown in Fig. 4.32(b). This explains the fractography of

the broken cutting wedge shown in Fig. 4.33. As shown, the surface of fracture includes

two distinctive zones: one closer to the rake face represents the black region shown in

Fig 4.31 while the other resembles the fractography of normal brittle fracture of tool

carbide.

To evaluate the described destruction quantitatively, a microhardness scanning test was

carried out along the lines shown in Fig. 4.34(a). Figure 4.34(b) shows the distribution of

microhardness along these lines. As shown, the maximum microhardness was observed

in the regions of the cutting wedge adjacent to Zone 1 as this zone is always subjected

Fig. 4.32. Micrographs of the cutting wedge in Zone 2: (a) new tool and (b) tool after 7 min

of machining. Work material – tool steel HVG (1% C, 1% Cr, 1.5% W, 1% Mn), tool material

M30 (82% WC, 7% TaC, 3% TiC, 8% Co), cutting regime: cutting speed ν = 3.3 m/s, feed

f = 0.3 mm/rev, depth of cut d

= 1.5 mm, cutting time τ = 7 min, magniﬁcation ×2000.

264 Tribology of Metal Cutting

Fig. 4.33. Cutting wedge breakage as a result of plastic lowering. Fractography of breakage shows

two distinctive surfaces of fracture.

to high contact compressive stresses during machining [25,26]. The normal, for the tool

material used in the test, microhardness was observed in the regions of the cutting wedge

adjacent to Zone 3 as these regions were not subjected to high temperature and stresses

in cutting. The lowest microhardness readings are observed in the regions of the cutting

wedge adjacent to Zone 2 as a result of high temperature enhanced difussional stress

relaxation that led to the softening of the tool material in these regions.

Cutting Tool Wear, Tool Life and Cutting Tool Physical Resource 265

(a) (b)

0 0.2

13000

14000

15000

16000

17000

18000

HV (MPa)

0.4 0.6 0.8 1.0

Distance from

the flank (mm)

III

Fig. 4.34. Cutting wedge structure after 7 min of machining: (a) micrograph ×400, (b) micro-

hardness distribution along lines I, II and III. The cutting conditions are the same as in

Fig. 4.31.

Talantov suggested [24] that the plastic lowering of the cutting edge takes place dif-

ferently for different tool materials. He developed two distinctive modes shown in

Figs. 4.35 (a) and (b) for carbide tool. The ﬁrst one shown in Fig. 4.35(a) is charac-

terized by the plastic lowering of the cutting and by bulging of the tool ﬂank surface.

This mode is common for carbides having low thermal conductivity as carbides of ISO

P group (especially containing titanium carbides) when they are used for the machin-

ing of difﬁcult-to-machine work materials of high strength and low thermoconductivity

(a)

(b)

Fig. 4.35. Two modes of plastic lowering of the cutting edge: (a) for low-thermoconductivity

carbides, ISO P group (b) for high-thermoconductivity carbides, ISO M group.

266 Tribology of Metal Cutting

i.e. when high temperature and high stress region in the cutting wedge are located close

to the cutting edge. The second mode shown in Fig. 4.35(b) is particularly for carbides of

higher thermoconductivity such as carbides of ISO M group, especially when the uncut

chip thickness (and thus the contact length) is great.

Tool breakage takes place when plastic lowering of the cutting edge reaches a certain

limit. This is because great interfacial stresses develop between plastically deformed due

to creep volumes of the tool material in Zone 2 and those in Zones 1 and 3. As a result

of these stresses, fracture takes place over the border between Zones 1 and 2 (more

common) or between 2 and 3 (less common). The former case is shown in Fig. 4.33.

Reading the described model, one may wonder if the plastic lowering of the cutting edge

is not a continuous phenomenon but rather occurs during a short time just before the

actual breakage of the cutting wedge. In other words, it can be assumed that the occur-

rence of the plastic lowering of the cutting edge results in a signiﬁcant increase in the

cutting force that, in turn, leads to breakage. To prove that this is not the case, a special

cutting test was carried out where the components of the cutting force were recorded

using a high sensitive tool dynamometer as well as the level of micro vibrations of the

cutting tool using built-in accelerometers. The results of this test indicated that there is

no increase in the cutting force and tool vibration or any other spikes prior breakage of

the cutting wedge. Moreover, a good repitition of tool breakage life (12 min 30 s ±12

)

conﬁrms that the plastic lowering of the cutting edge is a steady-state physical process.

4.8.3 Reduction

Plastic lowering of the cutting edge can be controlled to a large extent by the parameters

of the tool geometry. Among these parameters, the ﬂank angle, the conﬁguration of the

rake surface and the nose radius are of prime importance. Figure 4.36 shows proﬁles

of the ﬂank surface for different ﬂank angles. As shown, the smaller the ﬂank angle,

smaller is the plastic lowering of the cutting edge. This leads to an important conclusion

known from machining practice: in machining of difﬁcult-to-machine materials, the ﬂank

angle should be chosen as small as is possible. The limitation is the possible interference

of the tool ﬂank face and the machined (transient) surface. Therefore, an additional

module should be included in commercial programs for the veriﬁcation tool path on

CNC machines that can verify the absence of the said interference.

Figure 4.37 shows the inﬂuence of the conﬁguration of the rake surface. As shown, the

maximum cutting edge lowering occurs when “normal conﬁguration” of the tool rake

face is used. A signiﬁcant reduction in this plastic lowering is achieved when the rake

face is reinforced by a chamfer. It follows from Fig. 4.37 that the optimal geometry of

this chamfer exists for a given cutting condition. It should be pointed out that the width

of the discussed chamfer is much lower than the uncut chip thickness hence the tool is

not of the so-called negative geometry.

Figure 4.38 shows the inﬂuence of the tool nose radius. As shown, with the increase

in nose radius, plastic lowering decreases particularly in the region of the tool point.