Journal of Materials Processing Technology, 33 ( 1992 ) 95-108 95 Elsevier

Evaluation of lubrication and friction in cold forging using a double backward-extrusion process

Armin Buschhausen IBF, Technical University of Aachen, Aachen, Germany

Klaus Weinmann Michigan Technological University, Houghton, MI, USA

Joon Y. Lee and Taylan Altan Engineering Research Center[or Net Shape Manufacturing, Columbus, OH, USA

Industrial Summary

In cold-forging operations, the material flow and the quality of forged parts are functions mainly of the tool geometry, the amount of deformation or strain, and the conditions at the workpiece/ tool interface (e.g., friction, relative surface velocity, surface finish and heat transfer). While many of these parameters are well known and controllable, the influence of friction is often diffi- cult to predict and depends on a variety of factors. To reduce the amount of friction in cold forging, billets are generally coated (e.g., phosphate coating) and lubricated. In this paper a friction test, based on a double backward-extrusion process, is proposed and examined in order to obtain in- formation on lubrication quality. In this test, the upper punch moves downwards, while the lower punch and the die are stationary. An FEM analysis, using the program DEFORM, has been con- ducted for different area reduction ratios and billet heights. The simulation results show reason- able agreement with the results of experimental studies performed in Germany some years ago. The reduction ratio that gives the greatest differences in extruded cup heights was selected for the test design and the influence of friction shear factors between m = 0.08 and m = 0.20 was investi- gated. Based on the FEM simulations, calibration curves were established. Using these calibration curves and measuring only the heights of the extruded cups and the punch stroke in experiments enables the quantification of the friction factor and the evaluation of the lubrication conditions under production conditions.

1. Introduction

In most metal-forming processes the trend is towards near-net-shape and net-shape manufacturing, resulting in savings in material, energy and finish- ing steps. For example, the importance of cold forging for the production of parts in the automotive and transportation industry has increased signifi-

Correspondence to: A. Buschhausen, IBF, Technical University of Aachen, Aachen, Germany.

0924-0136/92/$05.00 © 1992 Elsevier Science Publishers B.V. All rights reserved.

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cantly during the last few years. Most of the parameters defining this metal- forming process are well known and their influence on material flow and part quality can be controlled. One parameter, however, namely friction, is difficult to predict and depends on a variety of factors.

The cost of lubricants is small compared to the costs of items such as raw material, equipment and labor. As a result the economic incentive to evaluate or change lubricants is not always very significant. However, one of the largest factors in lost production during cold forging is excess die wear and die failure, due to a break-down of the lubrication with subsequent product rejection. Therefore, it is essential to evaluate the lubricants in use and to compare them to alternative types of lubricants. Such an evaluation is necessary in order to utilize effectively the large investment required for installing a coating and lubrication line for cold forging.

The objective of this study is to develop a lubrication test that reflects the conditions of industrial production (i.e., high surface pressure, severe material flow, substantial surface enlargement, etc. ). The test design is based on a dou- ble backward-extrusion process (see Fig. 1 ) and should be able to perform the following functions:

initial stage daring deformation

upper

avel

o ± ~ ~S~L±OHaLy) ioweL pUL~CH ~Lationary)

Fig. I. Doub|e backward extrusion.

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(i) determine a friction factor using experiments on the one hand and ap- propriate calibration curves (determined by computer-aided metal-flow simulations) on the other hand;

(ii) give information on the quality of a lubrication line in a cold forging plant; and

(iii) compare various lubricants as well as lubricant application conditions simultaneously.

2. Evaluation of friction in metal forming

Lubricity in cold-forging processes is most commonly measured by the ring- compression test developed by Male and Cockcroft [1 ]. A flat ring-shaped specimen is compressed to a defined reduction. The change in internal and external diameters of the forged ring depends on the friction at the tool/work- piece interface. When the ring is compressed with zero friction, it deforms in

AL FORGING(M=O. I), • I + 0 0 ! 0 o . l o 3 C A L l SO .O01 [SZ I 0 . 6 0

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i i i v i v 4 . 0 0 0 . 0 0 0 . 5 0 l . O 0 l . S O 2 . 0 0 2 . 5 0 | . 0 0 3 . 5 0

I A D I U l ( Z l )

• ] . 0 0

| . 0 0 S C A L E

5 0 . O O E S Z

1 . 0 0

O,O0

- 1 . t 0 I I I I I I

- 2 . 5 0 - z . s o - e . s e e . s e 1 . 5 0 2 . S 0

Z - A Z I • I | 1 1 )

Fig. 2. Forging pressures for the ring test and a "bucket" forging [3 ].

i

S . S •

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the same way a solid disk would. Each element of the workpiece flows radially outward at a rate proportional to its distance from the center [2]. With in- creasing deformation, the internal diameter of the ring is reduced if friction is high and is increased if friction is low.

High interface pressure and severity of deformation in industrial cold-forg- ing operations, however, cannot be reproduced realistically in the ring test, as discussed in an earlier investigation [3] and seen in Fig. 2. Therefore, there is a need for a new test method, which reflects the actual process conditions in cold forging more accurately. In such a test the following variables must be comparable to those observed under real cold forging conditions: (i) the contact time between the workpiece and the tool under pressure; (ii) the ratio of newly generated deformed surface area to the original surface

area of the billet before deformation; and (iii) the relative velocity between the workpiece and the die.

Recently, a new test method for cold-forging lubrication method was pro- posed by Sanchez et al. [4] at Michigan Technological University (MTU). This test is based on a combined forward rod and backward cup extrusion process. A global friction factor is estimated quantitatively by relating it to the length of the forward extruded bar for a specified displacement of the punch. Calibration curves, based on upper-bound calculations, give m-values between m-- 0 and m-- 1. This test was developed for the cold extrusion of aluminum and should have been adapted for extrusion of steel. Corresponding experi- ments conducted by Popilek [5] at the Saginaw Division of General Motors showed, however, that the test is not very sensitive to small changes in the actual lubrication conditions.

3 . A n e w f r i c t i o n t e s t

A new test design, using a double backward-extrusion process, see Fig. 1, was developed. In order to determine the reliability of the FEM simulation, it was necessary to compare simulation results immediately to experimental data. Therefore, the test geometry for the FEM analysis has been designed after experiments conducted by Geiger [6].

In Geiger's experiments (Fig. 1 ) the workpiece has a diameter of 28 mm; the height varies between 28 mm and 56 mm. The upper punch moves with a con- stant velocity of 10 mm/s; the lower punch and the die are stationary. The punch design is based on recommendations by the German Society of Engi- neers [7]. The same information regarding tool design for cold extrusion of steel is included in ICFG data sheet 5/71. This punch design has been found to be most useful for the backward extrusion of steel. Experiments with differ- ent punch shapes, conducted at Battelle Memorial Institute [8 ], show that the punch design chosen for this project ensures an optimum spread of lubricant along the die/workpiece contact area during deformation. The lubricant is re-

99

tained under the punch nose and meters out a small amount at a time, just enough for continuous uniform flow. It enables the punch to penetrate deep into the workpiece without galling or increase in punch load.

Geiger used a slightly different punch design for the majority of his experi- ments. He also examined the punch shape that is used in the D E F O R M simu- lations, and reported that the difference between the conical faced punch, used for the majority of his experiments, and a conical faced/flat centered punch is almost negligible. Small variations of the punch nose do not influence signifi- cantly the material flow or the load requirement. Consequently, the compari- son between FEM simulation results and Geiger's experimental data is as- sumed to be acceptable.

The task for the present project has been to find a combination of process parameters that would best indicate a small change in the lubrication condi- tions. The friction conditions might change towards the end of the deformation due to thinning of the lubrication layer, but this effect can be neglected if the punch does not move very deep into the workpiece. The test method for lubri- cant evaluation, proposed in this study, can be described as follows. The heights of the extruded cups hi and h2 as well as the punch stroke are measured in the experiments (see Fig. 1 ). These data are inserted in calibration curves, which are generated by FEM simulations using different m-values. Thus it is possible to determine the quality of the lubrication. In the near future, experimental studies will be performed in close cooperation with Saginaw Steering Gear Division of General Motors to verify the simulation results in detail and to prove the practical usefulness of the proposed method.

4. S imulat ion results and discuss ion

4.1. Numerical simulations by DEFORM Numerical modelling of metal-forming processes is based on the plasticity

theory. Since the plastic deformation in most metal-forming operations is much greater than the elastic deformation, the influence of the elastic component can be neglected. This so-called rigid-plastic model assumes that the material does not show any deformation until the stresses have reached the yield point. The fundamentals of the rigid-plastic FEM and the mathematical formula- tions used to develop the computer program DEFORM, used in this study, are well established and given elsewhere [9,10].

The objectives of the D E F O R M simulations of the ERC test design were spec- ified as follows: (i) to find a reduction ratio that best indicates changes in lubrication

conditions; (ii) to examine the influence of the billet height; and

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(iii) to simulate metal flow for various friction conditions in order to establish calibration curves for different friction conditions.

In the combined backward-extrusion process chosen as the test geometry, the reduction ratios are the same for the top and bot tom areas of the billet. Both punches have the same nose shape. While the lower punch is stationary, the upper punch is moving with a constant velocity of 10 mm/s . The higher cup is extruded against the direction of punch movement due to different rel- ative velocities between the container, the punch, and the workpiece: this re- sults in different friction conditions [6].

In this study, the friction shear factor "m" was chosen for characterization of friction because it represents the friction conditions in forging, where the interface stresses are large, bet ter than the Coulomb law. Furthermore, the m- value is very common in the USA. Thereby any results of the proposed test method could easily be compared with other methods or data.

Based on Geiger's experiments, the following reduction ratios

2 2 e = dpunch/dbil let

have been analyzed first: e---0.25; e--0.51; and e=0.71 for punch diameter: d = 14 mm; 20 mm; and 23.6 ram, respectively.

The material selected for the simulations was the same as that used by Geiger, i.e., AISI 1006 with the following flow stress behavior

d = K ( n

with K = 700 N / m m 2 and n = 0.245.

E

r -

40 n~ 0

ooOOOoooOO° 0 acP °on oOO'1 + "I-

1 .°" ° [ ] ° °

1 ° ° ° ° 1°t ° 0 i ~ - i

0 5 1 15 20 25

hl, e=0.20

h2, e=0.20

hl, e=0.25

hl, e=0.25

hl, e=0.51

h2, o=0.51

hl, o=0.71

h2, o=0.71

punch stroke [mm] Fig. 3. Cup heights hi and h2 versus punch stroke for reduction ratios e = 0.71; 0.51; 0.25; and 0.20. (friction factor m =0.1; and billet geometry ho/do = 1 )

s = 12 mm

5 2 . 0 5 2 . 0

s = 24 mm

3 9 . 0 3 9 . 0 c

u

o h p 2 6 . 0

i 13 .0

g

h

t 0.0

(

• - 1 3 . 0 ]

- 2 6 . 0

- 3 9 . 0

0 .0 10.

radius ( m n ]

s = Omm

26 . 0

1 3 . 0

0 . 0

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3 9 . 0

2 6 . 0

13.0

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- 1 3 . 0

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0 . 0 1 0 . 1

Fig. 4. DEFORM simulat ion of E R C tes t design. ( reduct ion ratio e = 0.71; and frict ion factor m = 0.10)

s = 0 mm s = 12 mm s = 24 mm

3 0 . 0

c

u

p 1 5 . 0

h

e

i

g

h 0 . 0

t

(

] - 1 5 . 0

- 3 0 . 0

0 . 0 10 . 0

0.0

-15.0

- 3 0 . 0

i

. 0 1 0 . 0

I

0 . 0 1 0 . 0

101

30.0

1 5 . 0

0 . 0

- 1 5 . 0

r i ~ l U ! [ 1 1 ]

Fig. 5. DEFORM simulat ion of E R C tes t design. ( reduct ion ratio e = 0.51; and fr ict ion factor m = 0.10 )

15.0

c

u

p

h

e

i

g

h

t

[

m 0.0

m

I

s = 0 mrn

102

0.0 g.o

radius [mm]

0.0

-15.0 - 1 5 . 0

.0 9 . 0

~. mm

]

15 .0

I

9 . 0 0 . 0

2 4 mm

Fig. 6. DEFORM simulat ion of E R C tes t design. ( reduct ion ratio e = 0.25; and friction factor m = 0.10)

30

20 _t~O IJ'- 0 hl, m=O.08 ,(~1oOO O ~ O h2, m=0.08

,O • hl, Geiger's test

10 _ ~ ~ • h2, Geiger's test

0 5 10 15 20 25

punch stroke [mm] Fig. 7. Compar i son of expe r imen ta l [6 ] and p red ic t ed cup he ights hi and h2 versus p u n c h s t roke for the reduct ion rat io e = 0.25 and bil let geomet ry h o / d o = 1.

E E

c -

x :

Q_

O

4.2. Results and discussion Figure 3 shows the heights of the extruded cups hi and h2 versus punch

stroke for the selected reduction ratios and "m" =0.10. The cup heights have been measured from the plots of distorted mesh, obtained with D E F O R M and

103

E E

o t - -

f 3 L

0

30

20

10

0

J

u u m n

0 5 10 15 20 25

O hl, m=0.20

O h2, m=0.20

• hl, Geiger's test

• h2, Geiger's test

punch stroke [mm]

Fig. 8. Comparison of experimental [6 ] and predicted cup heights hi and h2 versus punch stroke for the reduction ratio e = 0.51 and billet geometry ho/do = 1.

3

od ¢ -

0 000000000 o °_aaaaa a °Oooo _ .

O DD Q DE~3 D O0000000uo" [] °Oaaaaoaaaaaac

o n

[ ]

~1.+ -r -,--r:F-r,e=~

i ! ! !

5 10 15 20 25

0 e=0.20

[ ] e=0.25

A e=0.51 + e=0.71

punch stroke [ram] Fig. 9. Predicted ratio of cup heights (h i~h2) versus punch stroke for different reduction ratios e; for the billet geometry ho/do = 1; and for friction factor m = 0.10.

given in Figs. 4, 5 and 6. The data for punch stroke have been extracted from the results of DEFORM simulations. The height of the upper cup hi is bigger than the height of the lower cup h2 due to different relative velocities resulting in different friction conditions as described before. Simulation results and ex- perimental data are compared in Figs. 7 and 8 for reduction ratios of e = 0.25 and e=0.51 with m--0.08 and m=0.20, respectively. Figure 7 shows good agreement between the D E F O R M simulation and Geiger's experimental data [6]. However, Fig. 8 indicates tha t the simulation data do not follow the t rend

104

4

oJ c -

¢ -

0 U- 0 - " ~ - - ~ 0 0 - - 0 0 ^ 0 0 0 0 O O w ~ o,-,,-,o ° 9~ ~==¥,==Wm=~======

O ~Vm • • • m 0 mm •

• m m + + + + + + + _ _ _

m+ + + + r r + + + + + + + + + - I ' + + + + + + +

' ' ' 2 ' 0 5 10 15 25

0 ho/do=2.0

• ho/do=l.5

4- ho/do=l.0

punch stroke [mm]

Fig. 10. Predicted ratio of cup heights (hi~h2) versus punch stroke for various billet geometries, ho/do; for friction factor m=0.10; and for the reduction ratio e=0.25.

of Geiger's experimental data. This may be caused by the measurement errors and the change of friction during extrusion.

In Fig. 9 the ratio of cup heights hi~h2 is plotted versus punch stroke for e = 0.71, e-- 0.51, e-- 0.25, and e-- 0.20. The ratio hi~h2 gives a non-dimensional parameter, which allows better comparison between results. The higher the value hi~h2, the greater is the difference between cup heights. For the reduction ratio of e = 0.20 the cup heights differ most. Consequently, any changes in the friction conditions would be best detected by using this reduction ratio. How- ever, the reduction ratio of e=0.25 was chosen for this study because the re- sults from D E F O R M simulations could be compared to Geiger's experimental data [6 ]. Furthermore, punches with small diameters are difficult to manufac- ture and may fail relatively easily. The thinner is a punch compared to its height, the higher is the possibility of vertical bending of the punch. This re- sults in eccentric parts, which would impair the quality of the experiments. In the worst case bending even could lead to failure of the punch.

Each curve reaches a maximum after the initial increase. With increasing reduction ratio, the magnitudes of initial peaks decrease. It was found that the stroke for the peak ratio (h~/h2) coincides with the punch stroke that provides the lower cup height close to the length of the punch land. As the top surface of the lower cup passes through the space between the punch land and the die insert, the resistance to metal flow in the lower cup increases, compared to that of the upper cup. This transitional behavior of the cup heights is dependent on the length of the punch land. The initial irregularity of predicted cup height ratios, seen in Figs. 9 and 10, is caused by two factors: (a) the volume loss in the billet during DEFORM simulations (due to penetration of the mesh area

4 5 . 0

105

3 5 . 0

25.0

15.0

5.0

-5 .0

-15 .0

0 . 0 1 0 . 0

m21

0

2 0 . 0 3 0 . 0 4 0 . 0

r a d i u s m m

Fig. 11. Normal pressure on boundary nodes at container side. (reduction ratio e = 0.25; friction factor m--0.10; billet geometry ho/do = 1; and material AISI 1006)

into the punch surfaces at the sharp corner of radius); and (b) the cup height measurement errors especially for the smaller punch strokes.

Figure 10 shows the ratio of cup heights hz/h2 for various billet height to diameter ratios, ho/do=l.O, 1.5, and 2.0. The differences in cup heights are largest for ho/do = 2.0. The larger the billet height to diameter ratio, the bigger is the maximum difference in cup heights. As shown in Fig. 10, the magnitude of the cup height ratio does not increase significantly in the range from ho/ do--1.5 to ho/do--2.0. Therefore, the billet height ratio does not have to be bigger than 1.5. Moreover, the height of the extrusion punch should not exceed

106

i 'M t -

.E-

+ , - k + + + + + + +'1"+ + + +

+.I.+:A A A & A ~ A A A . +++++-I--F+++.~

A DDDi3DDD ~aAAAA&AAA~AAA~' + ADo DE]DO,.,_ A~u OoOoOOOo ~[3DDDDDDDDDDDE

+nooO o OOoOooooooOooOOO C O

O

' ' ' 2'0 0 5 10 15 25

punch stroke [ram]

o m=O.08 [ ] m=0.10 & m=0.12 4- m=0.15

Fig. 12. Predicted ratio of cup heights (hl/h2) versus punch stroke for various values of the friction factor m. (reduction ratio e=0.25; and billet geometry ho/do= 1 )

TABLE 1

The hi~h2 ratios for various friction values: m=0.08 , 0.10, 0.12, and 0.15 for the steady state portion of the hi/h2 curves versus punch stroke

Friction values (m) hi~h2 ratios

0.08 1.86 0.10 2.15 0.12 2.37 0.15 2.73

2.5 times its diameter to prevent failure. This causes a limitation to the punch travel for the chosen geometry. The stroke of the upper punch should be less than 35 mm for a reduction ratio of e--0.25. To be thorough, punch strokes up to 30 mm have been investigated. Since for ho/do = 2.0 the workpiece is not deformed sufficiently with this limited punch travel, m-value variations have been examined for ho/do = 1.0 and ho/do --- 1.5 only.

Figure 11 shows the normal pressure on the boundary nodes calculated with DEFORM. For the area between the punches there is an approximately constant normal pressure on the container wall. The pressure decreases at the sides of the punches. Along the relieved part of the punches there is no pressure. No significant influence on the amount of normal pressure could be detected for variation of the m-value between m = 0.10 and m = 0.15.

Figure 12 shows the ratio of cup heights versus punch stroke with different

107

m-values for ho/do = 1.0. The curves have the same overall tendency and reach a constant value after the initial transitional stage, as discussed earlier. As expected, the ratio hi~h2 increases with increasing friction factor, m. This characteristics will enable using these curves for the calibration of the friction values for cold/warm forging. The hi~h2 ratios are listed in Table 1 only for the steady state portion of the curves. No significant influence of various m- values on the load data could be detected.

5. Conclusions and recommendations

In this study a new test method has been proposed for the evaluation of friction and lubrication, using the double backward-extrusion process. The material flow during deformation has been investigated by numerical modell- ing using the 2D-FEM program DEFORM. The influence of various reduction ratios and billet geometries has been examined.

Calibration curves have been determined for the geometry that best indi- cates small changes in the friction conditions. The differences in load on the upper and lower punches are relatively small. The calibration curves, i.e., hi~h2 versus upper punch stroke for various m-values, show distinct differ- ences even for small variations in friction conditions.

The results of the simulations show that the proposed test is able to evaluate friction quantitatively as well as qualitatively. The test conditions are close to those found in industrial production, i.e., they exhibit a high interface pressure and a large amount of deformation. The practical application of this test should allow the evaluation and control of lubrication quality under production conditions.

The assumption of a constant friction factor along the entire interface may not reflect the microscopic conditions in the actual process. Comparison of some of the results of the simulations with experimental studies performed in Germany several years ago, however,show reasonable agreement [6]. The evaluation of process conditions by computer-based simulations has been found to be accurate in many previous cases. Therefore, the results are considered to be realistic.

Future research should be aimed at: (a) conducting experiments under "near-production" conditions to illustrate

the "real" value of the proposed test; and (b) exploring the use of a simple backward-extrusion process, together with

FEM simulation, for evaluating friction.

Acknowledgements

The authors wish to thank various industrial members of the ERC for Net Shape Manufacturing, in particular Mr. Gerald O'Brien of GM-Saginaw, for

108

their advice t h r o u g h o u t the course of the work. The work was conduc ted under the N S F cooperat ive ag reement # ECD-8943165, suppor t ing the E R C / N S M , located at the Ohio Sta te Universi ty .

References

1 A. Male and M. Cockcroft, A method for the determination of the coefficient of friction of metals under condition of bulk plastic deformation, J. Inst. Met., 93 (1964) 38.

2 T. Altan, S. Oh and H. Gegel, Metal Forming: Fundamentals and Applications, American Society for Metals, Metals Park, OH, 1983.

3 E. Kropp, T. Udagawa and T. Altan, Investigation of metal flow and lubrication in isothermal precision forging of aluminum alloys, in: P.W. Lee and B.L. Ferguson (Eds.), Proc. A S M Conf. on Near Net Shape Manufacturing, Columbus, OH, November 1988, ASM Interna- tional, Metals Park, OH, p. 37.

4 L.R. Sanchez, K.J. Weinmann and J.M. Story, A friction test for extrusion based on com- bined forward and backward flow, T. Altan (Ed.), Proc. the 13th North American Manufac~ turing Research Conference, May, 1985, Society of Manufacturing Engineers, Dearborn, MI, pp. 110-117.

5 M. Popilek, A friction test for extrusion using combined flow with emphasis on backward extrusion, M.Sc. Thesis, Department of Mechanical Engineering, Michigan Technological University, Houghton, MI, 1991.

6 R. Geiger, Metal flow in combined cup extrusion, (in German), Reports from the Institute for Forming Technology of the University Stuttgart, Nr. 36, Girardet, Essen, Germany.

7 VDI-Richtlinie 3186, Werkzeuge ffir das Kaltflie£pressen von Stahl, (in German), Blatt 1- 3 (ICFG Data Sheets 4/70, 5/71, 6/72).

8 J. Becker, H. Henning and F. Boulger, Studies on workability by backward cup extrusion, Topical Report No. 9, Battelle Memorial Institute, Columbus, OH, 1970.

9 S. Kobayashi, S. Oh and T. Altan, Metal Forming and the Finite-Element Method, Oxford University Press, New York, Oxford, 1989.

10 S. Oh, W. Wu, J. Tang and A. Vedhanayagam, Capabilities and applications of the FEM code DEFORM: The perspective of the developer, J. Mater. Process. Technol., 27 ( 1991 ) 25-42.