Tool wear and tool life in intermittent cutting of hardened steel using conventional hardmetal inserts

Int. J. Mach. Tool Des, Res. Vo|. 18. pp. 19 28. ~? Pergamon Press Ltd. 1978. Printed in Great Brilain

TOOL

WEAR

AND

CUTTING

OF

CONVENTIONAL

0020-7357;78,0301-0019,?02.(~ 0

TOOL

LIFE

HARDENED

IN

INTERMITTENT

STEEL

HARDMETAL

USING

INSERTS

P. K. PHILIP* (Received 15 July 1977; in final form 13 October 1977)

Abstract--Intermittent cutting of hardened steel is characterised by impact stresses during tool entry, cyclical temperature fluctuation at tool contact zones and severe mechanical loading of cutting edge, which lead under most conditions to premature tool failure by fracture. An oil-hardened Ni-Cr steel of martensitic structure and hardness in the range of HV 400-700 is machined by face milling using conventional sintered carbide inserts. The suitability of various carbides for the task in terms of tool life is evaluated. The nature of initial tool contact for radiused inserts is discussed and the effect of tool entry conditions on tool life is determined. INTRODUCTION A WIDELY accepted feature in the production of metal forming tools and machinery parts in the past has been an adherence to the cycle machining-hardening-grinding. Currently a rationalisation of the process to one of hardening and machining promises to gain acceptance in practice. Even continuous cutting of hard, high strength metals presents problems of accelerated tool wear due to high specific cutting energy and temperature. In intermittent cutting as during face milling tool life is substantially shorter owing to alternating thermal and mechanical loading of the tool. It has been observed that a BUE does not occur in milling hardened steel under conditions corresponding to the use of carbides [-1], [2]. Hence finished surface quality in this case is determined by feed per cutting edge and tool geometry. As these can be controlled suitably without adversely affecting tool life, realisation of required surface quality is a minor problem and face milling hardened steel becomes practically viable, if reasonable tool life under competitive metal removal rate can be achieved. The present investigation deals with this problem. EXPERIMENTAL CONDITIONS The tests were conducted on a vertical milling machine with a spindle drive rated at 11 kW and a feed drive of 3 kW. Spindle speed range is from 63-1000 rev/min with a step factor of 1.4. The table speed has a stepless working range of 25-1000mm/min. Cemented carbide inserts 12.7 x 12.7 x 4.76 m m with a corner radius of 0.8 m m were used for the tests. Each one of these has eight possible cutting edges and are the most economical considering cost per cutting edge. Clamped on to a 2 5 0 m m dia milling cutter the tool geometry conforms to normal practice for face milling high strength steels. Rectangular blocks of hardened low alloy steel 42CrMo4 of following chemical composition and mechanical properties were used for the tests. C 0'o 0.42

Si ~o 0.25

Mn ~o 0.65

Cr % 1.05

Mo ~o 0.20

Ni ~o 0.60

HV B kp/mm 2 N/mm 2 400-700 up to 1400

In addition to high hardness attained this steel can be hardened to considerable depths without too drastic a fall in hardness from surface to interior. In the hardened state it has a martensitic structure. * Assistant Professor in Mechanical Engineering Indian Institute of Technology, Madras-600036, India. 19

20

P . K . PHILIP

Based on preliminary tests a feed per cutting edge, sz of 0.1 mm/rev was chosen bearing in mind the vulnerability of hardmetal to tensile stresses and the disproportionate increase in these at cutting edge with increasing feed. Cutting speed, v was 100m/rain. Depth of cut was kept at 2 m m to reduce variation in hardness across cut. Angle of engagement of tool was 45 °, which corresponded to almost symmetrical milling. EXPERIMENTAL

RESULTS AND DISCUSSION

Normal steel cutting grade of carbide, namely TiC-containing P grade is widely used for machining hardened steel in practice. However it was decided to try out various hardmetals before choosing a smaller section for more intensive study. Tool entry and exit conditions were chosen considering industrial practice. Results of preliminary tests are given in Fig. 1. Representative grades showing minimum flank and crater wear, namely P10, P20, M10 and K10 were chosen for more exhaustive tool life tests. The larger representation from P-group is due to its wide use in industry for similar tasks. Practical tool life criteria for machining of high strength materials have been proposed based on limitation of flank wear to a specific magnitude [3]. This is readily acceptable for turning and precision milling due to the accent on dimensional accuracy and surface integrity, but appears unsuitable in case of face milling hardened steel where the accent is on material removal rate and total machined volume per cutting edge. Considering the high rate of wear and probability of premature tool failure the cutting edge in this case should be used as long as it can effectively cut. Hence tool life corresponds to tool failure and this is the criterion employed in the present tests. Tool failure reveals itself through easily identifiable side effects such as sparking at cutting edge, change in form of chips and a tendency for instability in the tool-work system. Flank and crater wear variation for the four carbide grades plotted against cutting time and machined volume are shown in Fig. 2. It is seen that both types of wear are less for higher TiC contents of tool. This is expected, as TiC is known to be more wear resistant due to its higher hot strength, resistance to adhesive reactions with Fe and better stability against diffusion as compared to WC. The trend is confirmed by results obtained at a higher material hardness, which are presented in Fig. 3.

25

t = 12.Smin

>==!

t =12.5min

~m

//

zo

~o~m

500 ~

0o!

1s

//

3oo 6 200

-

¢

1oo P05 P10 PZ0

P30 P&O M10 M20

K01 K10 KZ0

P05

P10

Carbide grade

MIO M20 K01 K10

Carbide grade

Operation : Face milling Work materiat : Steel 1.2CrMo{, HV {,00

Depth of cut Feed per tooth

Toot

Machinedvotume

: Cemented carbide

P20 P30 P40

: a =2mm i Sz= 0.1mm

= 40cm 3

Cutting speed: v = 100mlmin. Cutting time : t =12-5min

Width of ftankwear l a n d VB and crater depth KT for various carbide grades. FrG. 1. Width of flank wear land V8 and crater depth

K T

for various hardmetals.

K20

Intermittent Cutting of Hardened Steel lO00

pm

Operation: Face milling

KT '

~brkmateriaCarbi l:St,2CrMO,Tool deHV ,00 :

V.

&00 Tool geometry:

i

~m

/

/

/.

75o10.8mm

63

/

/,///

o=,mm.

g0o_~o

/

/./"

63C

>I 250

21

4O

/

..-"

l0 .= o

160

o 10C ~. o

6.3 &.0

~.,,--~'~.~..~ 63 • K I0 /' o MIO p./" ./" j.1*" , P20 _ _ ~ 11 • P10 &G2.3 6.3 10 16 Cutting time t

='0

,~s

3¢5

5'o

W- Contact

25

8b

v~ 100m/rain 2.5 40 rain 63

260 cm3 3,'s

1~s

Machined volume

FIG. 2. Variation of flank and crater wear with cutting time for the hardmetals investigated.

It was observed that P10 and P20 tools failed after a short cutting time, although flank and crater wear magnitudes prior to failure were small. Profiles of Fig. 4 taken with a Talysurf shows wear development at rake face with cutting time. It is seen that in case of P-group of carbides wear begins nearer to the cutting edge as compared to M10 and K10. This leads to a weakening of the immediate cutting edge against impact with work at tool entry. Added to this is the inherent brittleness of TiC, of which P10 has around 45% by volume. Examination of cutting edges revealed microchip-

/

1600

ym 40-

°oT"22:£':":,0,

ym lOOC

25-

- Tool: HV500 Carbide a= 2turn , SI =0.05mm 63O Tool geometry:

16"

&O0

lO.

256

6.3,

160

~1/1/ k'ln

~~K10

g°l-9 J9o°l-5 l?5°lO.emm

/

o o

/

eVB lO0

6.3 I

10

V2Contact

/'

o KT

ll6

/ 10 I

25

16

v = 100m/min. = 63 rain t00

/-0

Cutting time

25 t

/o'O

65

100

160'cm,~ 2,~0

Machined volume

FIG. 3. Variation of flank and crater wear with cutting time for M I 0 and K I 0 at work hardness of HV500.

22

P.K.

P20

P10

t

PHILIP

MlO

TiC*TaG - 45%Vol. TIC-, TaG - 30%Vo

K10

TiC ÷ ToC - 20%Vol. "tiC* ToG <5% Vol.

4.0 rain

10

16

1

25 i

40

"luU,~m

Tool failure

Operation : Face milling

Tool : Carbide

Tool geometry :

Work material: ~2Cr Mo/, HV 400

Cutting conditions: a=2mm s Sz =0.1ram

= go

_9~o g~O

/~ I 7:o r -50 0.8ram

FIG. 4. T a l y s u r f profiles of r a k e faces s h o w i n g p r o g r e s s i o n of crater wear.

HV 500

s z = 0.05mm

HV 600

s z = 0.05mm

HV 650

s z = O.05mm

HV 700

100" rain80-

100 min, 80

min80-

.-~ 6o-

60-

60

60-

&0-

&O-

20.

20.

Ioo-

min~- 80:~ ,,.

-

s z = O.OSmm

_ &0-

,o

20-

V1 P 10 P20 HV400

M 10

KIO

sz =o.lmm

PIO

HV S00

P20 M10 K10 s z =o.lmm

P10 HV 550

s z = 0.1ram

HV 650

80m~ 60-

80-

80

60"

60-

80rain60-

40-

t.O-

1,0-

&0-

20

H

P10 P20

M10

K10

PIO

P20

B-

M10 K10

Operation : Face milling Work material: 42Cr Mo/~ Toot : Carbide a =2ram v = 100m/min.

20"

20-

20"

s z =0.1mm

eta

PTI

I;721 gZi

m

M

P10 P20

M10 K10

Pl0

1°20 M10

,,,,

Ir21 K10

FIG. 5. T o o l life v a r i a t i o n w i t h w o r k m a t e r i a l h a r d n e s s for feeds sz = 0.05 m m and sz = 0.1 mm.

Intermittent Cutting of Hardened Steel . i

63

23 315

i

Operation : Face m i l l i n g

cm 3

mm W o r k material:/*2 Cr Mo 4

6C

Tool

a=2.0mm, b

Tool

2OO

: Carbide s z = 0-1mm

geometry:

125

:o o'J r g z-, Ig0;l- l' O.'mm 80

16

,

\

o

50

e ~r 31.5

~ 6.3

20

4.0 V *-Contact v = 100rn/min 2.5

IZ5 3

/,~]0

450

500

560

630

710 k p

800

Work material hardness HV

FIG, 6. Tool life lines for the hardmetals with increasing work material hardness. ping in case of P10, P20 and M10. This is not critical as it has only a bevelling or rounding effect of cutting edge. With K10 wear is more conventional and tool failure occurred only at large values of wear. Carbide M10 has wear resistance and toughness between P l 0 and K10 and gave the highest tool life at work material hardness of HV400, which is about the maximum commonly encountered in heat treated steels. Tool life obtained at different hardnesses for the carbides tested is presented in Fig. 5. The vulnerability of P10, P20 and to a lesser extent M10 to premature failure through fracture is evident. This is all the more so at higher hardnesses, where all three forms of stresses, namely impact, mechanical and thermal are higher. Overall suitability of carbides from a tool life point of view for the task is seen from Fig: 6. It can be seen that reasonable tool life and material removal rate can be attained at hardness around HV400, which corresponds to most heat treated steel parts. In spite of its poor wearing qualities WC hardmetals give better performance at higher hardness compared to TiC/TaC tools. The ideal tool material for this purpose is one which combines high toughness with wear resistance. WC tools coated with TiC or TiN or both hold forth excellent promise on this consideration and investigation of their suitability form the next stage of the planned work. INFLUENCE OF TOOL ENTRY CONDITIONS AND RESULTANT IMPACT EFFECT ON TOOL LIFE Past investigations on face milling with sintered carbides have identified the following modes of tool failure [4], [5], [6], [7], [8]. (1) Conventional flank and crater wear leading to ultimate tool breakage. This form of wear is ideal but is achieved in practice in hardly 30% of cases. (2) Failure governed by brittle fracture in one form or another. The mechanisms singly or collectively active in this case are microchipping at cutting edge, local fractures due to initial impact and its propagation as well as fatigue cracks at contact zones and resultant fractures. This type of failure is typical of face milling and is influenced by the following factors. (a) Location of initial impact on tool and its magnitude. (b) Exit conditions in terms of chip weldments to tool. (c) Thermal fatiguing and cracking at rake and flank governed by milling width, chip section and cutting speed.

24

P . K . PHILIP

Work

FIG. 7. Type of initial contact in face milling with a sharp-cornered tool.

Thermal fatiguing leading to comb cracks was not significant at the low feeds that had to be employed in the present case. A remainder chip sticking to the tool on its exit from work piece is bound to cause fracture on tool entry, but this was eliminated by choosing easy exit conditions in which remainder chip stays attached to the exit edge of work. Microchipping of the sharp edge in the initial stages of cutting was observed to have only a rounding effect and did not contribute to subsequent fracture. Thus the predominant mechanism of tool failure for all carbides was fracture along transverse cracks due to mechanical fatigue and this phenomenon is primarily governed by entry conditions of tool in to work and resulting impact effect. It was decided to investigate this aspect in depth. Various possible initial contact points for a tool with a straight cutting edge and sharp corner are shown in Fig. 7, where parallellogram S T U V represents the region of tool contact on entry, if all points on rake face were to contact the work simultaneously. However this is seldom so and initial tool-work contact can be either at points S, T, U, V or along lines VU, UT, TS, SV in addition to area S T U V involving the whole section of the undeformed chip. The type of contact depends on the relative STUV

Idea I isad section of undeformed chip

S'TUV'

Actual section of undeformed chip

V'

Point of initial contact for the tests

u ~ T

YlI i!

! I r

I .-----4------4

,

I I

/

!

FIG. 8. Actual cross section of undeformed chip for the carbide inserts used.

Intermittent Cutting of Hardened Steel

25

position of tool and work in terms of angle of engagement as well as tool geometry [9]. A cutter having negative axial and radial rake as in the present case does not allow S, S T and S V contact and hence is a U-contact milling cutter 1-10]. The effect of tool entry conditions on tool life has been investigated in the past for both U- and S-contact cutters I-I1], [9]. In the former case a rapid reduction in tool life took place during the V-contact interval and in the latter during the T-contact interval. In all reported investigations initial contact conditions were found to be critical to tool life. The idealised chip section S T U V with a sharp corner S was used by Kronenberg and others to theoretically determine the location of initial contact from known geometrical parameters. This model chip section is a good approximation for HSS tools but involves a substantial deviation from actual geometry in case of hard metal inserts with generous nose radii. The chip section obtained in the present case S T U V ' s is given in Fig. 8 beside the idealised form S T U V . A particularly vulnerable contact point such as S in the idealised scheme does not exist in this case. Locations S' and V' are both weak but to a lesser degree compared to S. It is also to be noted that unlike in the case of V for the pointed tool V' is determined by tool geometry in terms of cutting edge angle and nose radius and not by feed per cutting edge s=. From Fig. 8 the distance of point V' from the main cutting edge sx = r(1--Sin(90--X)). Except in case of unusually low values of r and Z and large feeds it is impossible that s~ will be less than undeformed chip thickness and V' will be determined by the feed per tooth. On these considerations it was decided to determine the type of contact directly instead of using calculations or nomograms, which would have been tedious and incorrect anyway. This was done by using an easily identifiable coating on the vertical work surface which marked the point, line or area of initial contact on tool face, when the cutter was manually brought into contact with the leading edge of work for different values of angle of engagement. This proved to be a simple and speedy process. The slope angle of work surface at tool entry was changed for each value of axial displacement to correspond with the slope of transient surface produced by the tool at that angle of engagement. The results are presented in Fig. 9. The notation of points in Fig. 9 is to correspond with those of Fig. 8. Milling tests at various values of EE were conducted to determine the effect of different tool entry conditions on tool life. As the main effect of initial impact is on tool fracture, a hard metal particularly prone to this type of failure namely P10 was used for these

II-

J

D = 250mm

'

b = 180ram

"~Ve ~

0

~-

Ve

Ev

mm

120

95

80

60

50

40

30

20

10

10

20

30

40

50

60

BO

95

120

EE

o

75

50

40

30

25

20

15

10

5

5

10

15

20

25

30

/,0

50

75

V'

V'

V'

V'

V'

V'U

V'U

U

U

U

UT

T

T

T

T

T

T

T

Type of

contact Ev (~E -

Axial

displacement

U

Angle of (mgagement

FIG. 9. Type of initial contact of tool for various values of angle of engagement.

26

P.K. PHILIP luuu v rain

Operation: Face millin 9 Work material :42Cr Mo4 Tool

/

Carbide PT0

a = 2.0mm

v=t00mhnin

p Sz =0.2mm , D=250mm

d°

Tool geometry: g°

T

875 "7

HV 500 K p l m m2

-9 °

z

90 °

-5 °

~ o"

/

• 625-5

75° (~8mrr

500 -4

T

.1.

+

120

100

80

mill

•

60 ve

20

40

0

20

Axial displacement

40

6O

Ev

8O

~

i

i

,0

20

-100

-ve i

3b

Angle of engagement E E

mm

400 -re

FIG. 10. Tool life in terms of number of impacts plotted against angle of engagement.

tests. A moderate work material hardness of HV500 enabled the use of a feed of 0.2 mm at v = 100m/min. Proper choice of milling width ensured rather easy exit conditions for the tool. The results of the tests are given in Fig. 10 and Fig. 11. It is seen from Fig. 10 that tool life expressed in terms of number of impact cycles is lowest in the range of transition from V' to V'U contact. This corresponds to beginning of contact along a line almost parallel to the cutting edge at the radiused corner and marks out this location of initial impact as most vulnerable from tool life point of view. In fact transverse cracks resulting from mechanical fatigue were evident most frequently at this location for all carbide grades tested for tool life. Local fractures preferentially occurred in this region by propagation of these transverse cracks.

2O

\

T

T

,~ L6 U ~f

\

U o-

Operation : Face mitring Work material: 62 Cr Mo 4 HV 500 K p / m m 2

o

Tool

: Carbide P10

2.0mm , Sz=0.2mm v ~ loomlmin. , D=250mm

a=

Tool geometry : =E

9" I-9" I e0* I-s* I 75*latBm~ + 120

100 mm

80

~0

60 ÷ Ve

.eVe

9

2O

Axial

displacement

2O Ev

t.0 Im

-ve

mm

Angle of engagement El=

FIG. 11. Tool of life variation with angle of engagement in terms of machined volume.

Intermittent Cutting of Hardened Steel

27

Contrary to what one would normally expect, a T-contact is found to give better tool life than a U-contact. Under U-contacts the rake face of tools were observed to be prone to spalling. The magnitude of initial impact and its overall effect on tool life for a given feed and cutting velocity was sought to be explained by Kronenberg [-9] on the basis of 'partial time of penetration' meaning thereby the time elapsed between first tool contact and entry of weakest spot on tool, point S, in to work. The inadequacy of this parameter led Opitz and Beckhaus [-10] to formulate 'partial area of penetration' giving the portion of tool area S T U V that enters work before the weakest point S makes contact and to use it as a criterion for the impact effect on tool life. According to this when initial contact occurs at point U, all impact is born by the chip cross sectional area S T U V of tool rake face before point S enters the work and consequently tool life is high. In the case of T-contact a lesser proportion of S T U V will be loaded prior to point S and tool life is consequently lower. However the results of the present investigations are in contradiction to this and the lack of agreement can be explained on the basis of validity of the above criterion limited to the idealised sharp-cornered tool and not applicable to tools with generous nose radii. Figure l0 also shows that the initial range of V' contact, where large values of angle of engagement are involved, represent relatively mild conditions as far as impact effect is concerned. Examination of cutting edges under these conditions revealed local chipping at the point of initial contact, but this did not grow in size with increasing cutting time perhaps due to the very small undeformed chip section at the spot. The advantage of working in this region is further brought out by Fig. 11, which shows tool life in terms of machined volume per cutting edge for various types of initial contact. Tool life expressed in this fashion is more significant to practice than number of cutting cycles withstood prior to tool failure. The figure shows that machined volume per edge is maximum at large values of EE corresponding to V'- contact. The T-contact region is not far inferior in this respect and shows a maximum tool life corresponding to ee = - 2 5 °. This means that work pieces having large width can be advantageously milled with a large positive angle of engagement and narrow work pieces milled with a moderate negative angle of engagement. CONCLUSIONS Cutting tests show that conventional hardmetal inserts can be successfully applied in face milling hardened steel. The characteristic premature failure of tool by fracture can be avoided through proper selection of tool material, cutting conditions and toolwork position. Impact resistance of tool merit prime consideration in obtaining reasonable tool life. It was found that for steel with hardness around HV400 a TiC-TaC-WC hard metal, namely M10 is most suitable and that for higher hardness a straight WC grade K10 gives the best performance. The actual geometry of chip section for hardmetal inserts differ from the ideal model used in previous analysis. Initial contact conditions for the tool are determined directly for various angles of engagement. The tests reveal a significant effect of type of initial contact on tool life. It is found that machined volume per cutting edge is highest at large positive values of angle of engagement. T-contact at moderate negative values of eE gives comparable performance. Acknowledgements--The author thanks Prof. K. Tuffentsammer, director of the Machine Tool Institute at

Stuttgart University for providing facilities for doing the reported work, which was carried out under a research fellowship of the Alexander von Humboldt Foundation, Bonn. Thanks are also due to Dr. Ing. H. Schlayer for his encouragement. REFERENCES [l] U. KLICEPERA,Werkzeugmaschine Int. 6, 31 (1974). [2] YA. I. ADAMand S. Z. LOZEVA,Machines & Tooling 43, 45 (1972). [3] G. BARROW,Fertigung 5, 129 (1975). [4] W. LEHWALD,lndustrie Anzeiger 46, (1963).

2g

P . K . PHILIP

[5] N. N. ZOREV and H. A. SAW1ASKIN,Annls C.I.R.P. 18, 555 (1970). [6] B. L. DSA/VlOJEV,Maschinenmarkt 76, 988 (1970). [7] K. OKUSHIMA and T. HOSHI, Annls C.I.R.P 15, 309 (1967). [8] H. JONSSON, Tz f. praktische Metallbearbeitung 68, 139 (1974). E9] M. KRONENBERG, Trans. A.S.M.E 68, 217 0946). [10] H. OPITZ and H. BECKHAUS, Annls C.I.R.P 18, 257 (1970). I l l ] N. N. ZOREV, Annls C.I.R.P 12, 159 (1964).