Force and energy during the cold hooker extrusion of sintered iron preforms

Journal o/Mechanical Working Technology, 15 (1987) 357-374

357

Elsevier Science Publishers B.V., Amsterdam - - Printed in The Netherlands

FORCE A N D E N E R G Y D U R I N G THE COLD H O O K E R E X T R U S I O N OF S I N T E R E D I R O N P R E F O R M S

P. VENUGOPAL*, S. VENKATRAMAN**, R. VASUDEVAN* and K.A. PADMANABHAN*

*Metallurgical Engineering Department, HT, Madras (India) **General Technical Manager, M/S Heat Tech Engineers, 112 Kamaraj Avenue, Adyar, Madras (India) (Received January 27, 1986; accepted in revised form March 17, 1987)

Industrial summary The advantages gained by the forgingof a powder metallurgicalpreform are well known. Cold extrusion of a P / M preform combines the advantages inherent in the cold extrusionprocess with the forgingof a preform, but for successfuldeformation of a P / M preform, the inherent brittleness has to be overcome. In thispaper various cold extrusionmethods are examined, based on the state of stress,from which the Hooker extrusion process is chosen as the deformation method due to itsfavourable (compressive) stateof stress.The paper then goes on to deal with the experimental toolingevolved forthe Hooker extrusionof sinteredironpreforms. Compacts made under different compacting pressures (loads) and therefore of differentpreform densitieshave been Hooker extruded at differentextrusion ratios.The forcesand energies have been evaluated and the results compared with values calculatedfor the equivalent wrought steel.An empirical relationshiphas been obtained for the estimation of force during the Hooker extrusion of iron preforms, making use of force equations for the cold Hooker extrusion of wrought steel.

1. Introduction

Cold forging of preforms A process which could combine the advantages inherent in the manufacture of components by sintering and cold extrusion, thereby enhancing the mechanical properties of a powder metallurgical product, opens up a very promising route of manufacture. Although work in this area is not as exhaustive as that in other areas of metal fabrication, the process has generated sufficient interest among researchers. Biehl et al. [ 1 ] have patented a method for the cold extrusion of cans from P / M billets prepared from sponge iron powder. The mean density of the sintered billet increased from 6.2 to 7.4 g/cm a at an extrusion ratio of 3.6. Hoeness et al. [ 2 ] have forward extruded rods at an extrusion ratio of up to 4.0 and backward extruded cans at ratios of up to 5.3 from billets made from mixtures of reduced and atomised iron powders. All the backward-extruded cans were 0378-3804/-87/$03.50

© 198"LEt~ccmr Sciasae~Puhll.qh~r~ RN.

358 i Forward extrusion FORMING S H A P E ~

_~

Reduction % ~

88"9

R-~

6" 57 g/cm 3 R-S 675 g/¢m 3

38-5

o

0 0

x

o

x

E-4 6'04 g/cm 3

x

x

E-6 6.98 g / c m 3

o

x

g-8 7"22 g / c m 3

0

X

g/crn 3

88.3

0

R-8 7'06 g/cm3

A 6"85

Backward extrude

X

85"8

63.6 126-4 o 0

~x

I

i

I h

+

- 1+0

i-x

0

:

xl --o

'-

Pipe [Combined extrusion extrude

"I+1

A

Forw 80.3 Back

64.5 64"5 .beclcwa i d

•

BO-3

o

0

0

0

0

i

--

--

O

i -

-

-

A

-

-

A

-

-

o

--

a

•

[

L. . . .

+ [

0

-

-

Fig. 1. Resultsin the cold forgingof P/M iron preforms,after Obara et al. [5 ]. Results are classified into four groups: o perfect; • no crack but rough surface; A micro-crack; × crack. A dash indicates no experiments;R -- mill scale reduced;E -- electrolytic;A -- atomised. sound, but in the forward extrusion of rods, cracking occurred at extrusion ratios of less t h a n 2:1, when too m u c h of the billet was extruded, whilst at a ratio of 1.4:1 cracking always occurred. Nakagawa et al. [ 3 ] have explored various combinations of backward- and forward-extrusion and combined extrusion-upsetting; cracking tended to occur in products extruded at lower pressures and the lower extrusion ratios. Dower and Miles [ 4 ] have extruded billets compacted from iron powder into rods, tubes and cans, where the correct combination of billet density and extrusion ratio has been shown to result in crack-free extrudes. The results of various forms of cold extrusion of iron preforms investigated by Obara et al. [ 5 ] are shown in Fig. 1. References in literature to the Hooker extrusion [6] of P / M preforms are very limited. The work t h a t has been carried out to date is preliminary and exploratory in nature. T h e work of Ref. [ 4 ], concerned with establishing the viability of cold extruding a P / M preform by forward, backward and tubular extrusion, exploring a limited n u m b e r of variables, has shown that this process is technically viable. No investigative and detailed analysis appears to have been carried out in extruding a P / M preform by the Hooker process in the following areas: (i) force requirements; (ii) energy requirements; (iii) interacting influence of various extrusion parameters such as the die angle a (half included angle ), the logarithmic deformation ratio ~, the friction coefficient #, the initial preform density Pore, and the shape of the preform, towards the obtaining of defect-free Hooker extrudes; (iv) definite guidelines on tool design and the properties of

359 the extrudes; (v) economics of the process as compared to existing conventional practice. These aspects are covered in this paper and in papers to follow [ 7,8 ], based on the work of Ref. [ 9 ].

State of stress in various modes of cold forging Based on Ref. [10 ], the stress conditions in the following modes of cold working have been analysed: (i) drawing of a solid body; (ii) drawing of a hollow body; (iii) forward extrusion of a solid body; and (iv) forward extrusion of a hollow body (Hooker extrusion). ( i) Stress condition in the drawing of a solid body Figure 2 (a) [ 10] shows the stress condition for a work-hardening material. The boundary of the deformation zone is determined by the contact area of the tool with the workpiece. The force is applied by drawing (pulling) the work metal through the die. For axi-symmetrical processes, three principal stresses are present viz. axial stress ~z, radial stress at, and tangential stress at. The axial stress a~ is limited to the value of the tensile strength at the exit zone. a~ is tensile in nature, having its maximum value at the exit of the tool and decreasing to zero at the die entry plane 0. ( ii ) Stress condition in drawing of hollow bodies The stress condition is presented in Fig. 2 (b) [ 10]. The primary forming force is tensile in nature and the tensile stresses generated are maximum at the die exit. ( iii ) Stress condition in solid forward extrusion (Refer to Fig. 2(c) [10]) Deformation force Ftot~ is applied to the punch. The material becomes plastic at the die entry plane (0); axial stress az is a maximum here and is compressive. The plastic zone extends up to plane 1 (the exit plane), where the axial stress is zero. Both the tangential and radial stresses are compressive and there are no tensile stresses in the deformation zone. (iv) Stress condition in hollow forward extrusion (Hooker extrusion) The deformation zone lies between the planes 0 and 1 shown in Fig. 2 (d) [10 ]. The stress condition in the deformation zone is similar to that in the forward extrusion of a solid body. The maximum axial stress az occurs at the die entry plane 0 and decreases to zero at exit plane 1. All the three principal stresses are compressive. (v ) Isostatic and hydrostatic state of stress The cold extrusion of brittle materials can be successfully carried out by hydrostatic extrusion [ 11 ]. Hydrostatic extrusion imposes a totally compres-

360 (o)

(b)

i 1

z

r.

~-Y~ [~r

>.~o1

/stress

~z

-~---~ /

/deformalion zone

w._c

--, . ~e " ~ ' J

z r:

y/,}

"%,~J

J Tens,on

Comp,ess,oo

Axial stress o-z condition ~Tz-~Tr=Y(yietd stress)

Flow (c)

I (Tz

Flow

i

Ten sion

cr Compression

crz

condition

~z-O"r = Y

(a) z

~

I---~r nch

.,

III

Z

cr Compression

___ ~

Tension O'Compression

1

Flow

condition

o~- O~r=Y(yield

stress)

Flow condition

o~z-~r = Y

Fig. 2. Stress condition in (a) the drawing of solid bodies; (b) the drawing of hollow bodies; (c) solid forward extrusion; (d) hollow forward extrusion.

sive state of stress and lends itself very favourably to the deformation of materials; the isostatic pressing of P / M compacts derives its main advantages from this compressive stress state, where the hydrostatic stress contributes to the apparent ductility of the material during deformation. The process is, however, capital intensive and is not yet adaptable for mass manufacture. Hooker extrusion Based on a review of the literature on the cold forging of sintered P / M preforms, investigative and detailed analysis on the cold Hooker extrusion in terms

361 of forming variables appears to be unexplored. Cold extrusion as a deforming process for P / M preforms has techno-economic advantages when compared to hot forging, for generating rotational and axi-symmetric components. However, lack of ductility of a P / M preform imposes drastic limitations during cold working; even moderate tensile stresses lead to fracture. Researchers who have worked on cold extruding P / M preforms have had to contend with constraints arising from the inherent brittleness of a P/M compact. The analysis of the state of stress in the deformation zone during cold working, discussed above, leads to the following observations: (1) Drawing or ironing processes impose primary tensile stresses and hence their scope of application for P/M compacts is remote; ( 2 ) Forward extrusion either of a solid body or a hollow body imposes primary compressive stresses and in the deformation zone all the three principal stresses are compressive. Hence, this mode of deformation can be considered for P/M preforms; (3) Even though the forward extrusion of a solid body imposes compressive stresses, secondary tensile stresses can arise, due to the large differential velocities of the core and the skin, leading to chevron failures; ( 4 ) Forward extrusion of a hollow body ( Hooker extrusion) presents, probably, an ideal stress-state since the material is under compression, as with solid forward extrusion, but the differential velocities between the core and the skin are of a much lower order. Hooker extrusion, therefore, appears to offer the best state of stress for the cold deformation of P/M preforms when compared with other modes of cold working. The cold Hooker extrusion of sintered iron preforms

The foregoing dealt with the identification of cold extruding a P / M preform by the Hooker extrusion technique and suggests the need to carry out investigations in terms of the included angle of the extrusion die and the reduction, for Hooker extrusion. A knowledge of forming forces is essential for: (a) machine selection; (b) motor-capacity selection of the machine; (c) tool selection; (d) tool-material selection, (e) estimation of the manufacturing tolerances of formed parts. While deformation is generally assured in cold extrusion (due to all of the three principal stresses being compressive), the process seems to be limited by tool working-stresses and thus a need arises for the prediction of forces. During the cold extrusion of sintered preforms, densification and work hardening of the compacts can be anticipated. The friction associated with P / M cold extrusion can be different to that with equivalent wrought parts. Compacts with varying preform densities may demand varying forces and a partic-

362 ular minimum density of the preform is essential to inhibit workability during cold extrusion. It is to be expected that greater densification can be ensured by greater extrusion reductions, so that sound extrudes are possible; however, greater reductions will result in high tool stresses. The influence of die included angle has been shown to minimise the extrusion load during the cold extrusion of wrought parts, [ 10 ] and this concept can also be extended to the cold extrusion of P/M preforms. Objectives of the present work

The primary objective is to investigate the interacting influences of initial preform density, Pore, extrusion reduction E, die included angle 2~ and the beneficial effect of lubricant in the cold Hooker extrusion of sintered iron preforms. Studies into the force and energy required for P/M Hooker extrusion and their comparison with an equivalent wrought composition are also attempted. Derivation of the force for the cold extruding a wrought material by the Hooker extrusion process has been presented by Lange [ 10 ]. An attempt is made here to establish an empirical relationship for the force required for cold extruding a P / M preform by the Hooker process. On Hooker extrusion

General The forward extrusion of a hollow body is commonly referred to as Hooker extrusion; C.I.R.P. 5126 [ 6 ] defines the term as "extruding a recessed billet". If, instead of pulling a tube through a die to reduce its wall thickness (drawing of a hollow body), pushing is resorted to (i.e. Hooker extrusion), increased deformation can be attained due to the compressive state of stresses in the deformation zone: this method seems to have an attractive potential compared with ironing, yet another cold-finishing process. Forces in the Hooker extrusion of wrought parts Numerous investigators have studied the Hooker extrusion process [ 10,12 ], amongst the work of whom that of Lange is probably the most comprehensive. The forces involved in Hooker extrusion have been presented by Lange [ 10 ], refer to Fig. 3. The total force f t o t in the case of Hooker extrusion without bending is given by Ftot = Fid -t-Fsh + Ffr

(1)

363

i

TOT

?

! Fig. 3. Force components in Hooker extrusion. For symbols, see eqns. (1) - (8).

where Fid is the ideal deformation force, Fsh is shear force, and Ffr is the total friction force. The various components of the above forces are: Fid = Ao Ymean ~max

(2)

Fsh = l OgYmeanAo

(3)

Ffr=Frw+Frs+Frd, +Frd~

(4)

where ( refer to Fig. 3 ) : the frictional force along the container wall

Frw =7~dolYo/~

(5)

the frictional force on the shoulder of the container in the deformation zone

Frs =2YmeanemaxflA1/sin 2a

(6)

the frictional force on the mandrel in the deformation zone F~d~ = A 1 Ymeanem~xfl/tan o~ and the frictional force at the exit

(7)

364 Frd2 -----7~d2/r/21)r

(8 )

The values of/~/~r is 10 to 12 N / r a m 2 for Ma K6 and Cql0 material. Emax is the logarithmic deformation ratio of the extrusion reduction, Cmax= I n Ao/A1

(9)

and Ymeanis the average yield stress for a strain of em~x, ~max

Ymean-'l/~max |

K~"d~=K[E~,x]/(n+l)

(i0)

0

obtained from the flow-curvedata. I is the length of billet within the container, Ir is the length of the die land,/I is the value of boundary-friction coefficient as obtained from the ring-compression test and ~ir is the mean normal stress between the extrude and the die over the die land. Results and discussion

The Hooker extrusions were carried out on a fully i n s t r u m e n t e d Becker und Von Hiillen, 1000 kN, hydraulic deep drawing press. T h e experimental tooling evolved for the Hooker extrusion of sintered iron preforms is shown in Fig. 4. Extrusion loads were monitored by means of a laboratory-made load cell, ( making use of electrical resistance wire strain gauges manufactured by Hottinger Baldwin Messtechnick) and a Linear Variable Differential Transducer ( made by Hottinger Baldwin Messtechnick) of + 50 m m linearity range was used for monitoring the stroke. Results were registered on a Hewlett-Packard X - Y Recorder. The lubricant employed for the Hooker extrusion of all sintered iron preforms was molybdenum disulphide paste applied by hand, as used for the ringcompression test. The preforms tested were compacted at: (a) 150 k N to an initial preform density POre"- 5.9 g/cm3; (b) 250 kN to an initial preform density Pom= 6.35 g/cm 3 and ( c ) 350 kN to an initial preform density P0m= 6.6 g/cm 3.

Preliminary trials T h e preliminary trials conducted covered a range of initial preform densities, Pore= 5.7 to 6.8 g/cm 3 and the extrusion ratios attempted were e = 0.4, 0.6, 0.7, 0.9 and 1.2. T h e expression for the total force required in the case of forward extrusion (see eqn. (1)) was evaluated for dF/do~=O (i.e. for m i n i m u m force requirem e n t ) . Optimal extrusion angles ~ were obtained for the least-force requirem e n t for each value of ~: the values of ~opt obtained were 23 ° , 27 ° , 29 ° and 36 ° for e=0.6, 0.7, 0.9 and 1.2, respectively. Initial experiments were conducted with these optimum angles for E, the

365

~5

P res~ be~

Fig. 4. Experimental tooling for the Hooker extrusion on sintered iron preforms: 1,2 - - hardened plates; 3 - - punch shoe; 4 - - punch holder; 5 - - clamp ring for punch; 6 - - clamp ring for die unit; 7 - - container shrink ring; 8 - - container; 9 - - die ring; 10 - - die shrink ring; 11 - - die bolster; 12 cast iron block; 13 - - preform/billet; 14 ~ dummy hardened ring; 15 - - hooker punch; 16 punch body. -

-

results o f w h i c h revealed t h a t t h e e x t r u s i o n s carried out with OLopt for p r e f o r m densities pore=5.9, 6.35 a n d 6.6 g / c m ~, e n c o u n t e r e d severe c i r c u m f e r e n t i a l c r a c k i n g o f t h e e x t r u d e s , e x c e p t in the case o f e-- 1.2 for Pore-- 6.35 g / c m 3 a n d 6.6 g / c m ~. T h e ~opt c o r r e s p o n d i n g to e-- 1.2 was 36 °. T h e r e a s o n s for t h e cracking are to be dealt with in Ref. [13] .) U n d e r t h e s e c o n d i t i o n s of e=l.2 a n d ~opt, t h e m a x i m u m e x t r u s i o n load was f o u n d to e x c e e d 1000 kN, i.e. e x c e e d i n g t h e available press capacity. B a s e d on this c o n s t r a i n t , it was decided to r e s t r i c t t h e s t u d y to e = 0 . 9 . E n c o u r a g i n g results ( c r a c k free) were o b t a i n e d d u r i n g p r e l i m i n a r y trials with ~ - - 45 °, 60 °, e = 0.4 to 0.9 a n d initial p r e f o r m densities g r e a t e r t h a n 6.2 g / c m a.

Punch force-stroke plot P u n c h f o r c e - s t r o k e plots were o b t a i n e d d u r i n g t h e h o o k e r e x t r u s i o n o f P / M p r e f o r m s of: ( a ) initial densities p - - 5 . 9 , 6.35 a n d 6.6 g / c m 3 c o r r e s p o n d i n g to

366 75t

350 k I~ compact

(a)

= O. 916

/•,•-F.

60(

~.~,~£

Z

; 0.693

P450: 0

E = 0.442 JE u

c 300

:3 0-

0 L

0

5

20ram 10

15 Punch

"1 20

25

30

s t r o k e (ram)

~o__~o~. ~o~o~, ~_ ~

!

600

450 .c a.

300!

15[

0

5

10 Punch

15 20 stroke(mm)

25

30

Fig. 5, Hooker extrusion punch force-stroke curve for: (a) o~=45°; (b) or---60°. compacting loads of 150, 250 and 350 kN; (b) logarithmic deformation ratios of e=0.442, 0.579, 0.693 and 0.913; (c) die extrusion angles a = 4 5 ° and 60 °. Typical curves are shown in Figs. 5 (a) and (b). The characteristics of the punch force-stroke plots obtained are very similar to w h a t is obtained for a

367 TABLE 1 Variation of force and energy values with e for compacts compacted at various pressures during the Hooker extrusion of P/M iron preforms 3[ ~J~. [ ] ~ ~ - ~ - F i g u r e to indicate how Fmaxond F were measured from exhusion punch force stroke experimentol / f ~ " ' ~ F (of ZOmm curves ~~

stroke)

0. Punch stroke, mm

Die angle o~ ( o)

Compacting Initial load preform (kN) density, P0m (g/cm3)

e

45

150

5.9

0.442 0.579 0.693 0.916

45

250

45

350

Force (kN)

Energy (kNm)

F (at 20 mnl punch stroke)

Fn~x

Fm~x/F

315.0 394.0 435.0 600.0

315.0 420.0 480.0 630.0

1.00 1.07 1.10 1.05

3.66 4.97 5.74 6.77

6.35

0.442 309.0 0.579 397.5 0.693 457.5 0.916 574.0

330.0 457.5 487.5 607.5

1.07 1.15 1.07 1.06

4.29 5.89 6.56 7.56

6.60

0.442 0.579 0.693 0.916

367.5 442.5 540.0 682.5

1.11 1.09 1.16 1.14

5.53 6.28 7.89 9.87

330.0 405.0 465.0 600.0

wrought steel material, but the fairly sharp load peak observed for wrought material is not observed in the case of P / M preforms. T h e load peak, which is an indication of the break-through pressures required to initiate the extrusion process, is a fairly broad region and is seen distinctly for greater deformations. T h e ratio o f the peak load Fmax to the extrusion load F is indicated in Table 1. Since th e extrusion load after reaching the peak load shows a gradual decrease with p u n c h penetration, the value of F is t a k e n as a mean value for a p u n c h p e n e t r a t i o n of 20 m m for all cases so t h a t comparisons under identical conditions can be made. ( T h e m e t h o d is indicated in the top figure of Table

1). T h e Fm~ value is found to be up to 16% greater t h a n the extrusion load F for the range of extrusion variables studied. T h e ratio increases with increasing deformation and greater values are obtained for higher-density preforms. T h e F~a,IF ratio is, however, far less t h a n what is usually obtained for wrought

368

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6

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--A

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.

,

.

.

.

.

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.

.

.

.

.

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400

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300 0.3

/ X

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.

.

/ /

!/ ,./

~

.

X / =

~ .

/ ~ W,o~ght0.0s'/. Carbo,, Stee,

.

X X experimental o o experimental

I 0.L, 0.6 0"8 1.0 Extrusion reduc'(ion expressed in true stroin e.(In AO/A ~)

Fig. 6 (a). Showing the variation of maximum punch-force with extrusion reduction for different iron preforms.

materials; this appears to be due to the conditions of greater friction under which the extrusion of P / M sintered preforms takes place.

Variation of maximum punch force with extrusion reduction The variation of Fro= with e for the Hooker extrusion of P / M preforms for initial preform densities of 5.9, 6.35 and 6.6 g/cm 3 is shown in Fig. 6(a). For comparison purposes, a calculated plot for 0.05% C steel has also been included: the values for this wrought equivalent have been based on values calculated from Lange's expression for force, eqn. (1). The decision to use calculated values instead of experimentally derived values for the wrought equivalent was taken after obtaining extremely erratic results when the wrought material was extruded with graphite dispersion as the lubricant: this lubricant was chosen on the basis of ring-compression test results for obtaining identical frictional conditions to P / M preform extrusions using MoS2 lubricant. The erratic results were due to severe die pick up and the Hooker extrusions using graphite lubricant for the wrought equivalent could not be continued. The value of the coefficient of friction and the flow-curve data for the wrought equivalent

369 (b)

For die angle°(-=/,5 ° Porn 6"6glcm3 compocting IoQd 350kN Porn 6"35 g/crn3 ~m 5"9 g/cm 3

i

do d o -

250kN 150kN

I ¢c~,~,- k g/cm 3

~7 LU

;/ 3 0.3

0 ''~ 0.5 0"6 0.7 0"8 0"9 1"0 Extrusion reduction expressed in true strain E(InAo/A I)

Fig. 6 (b). Showingthe variation of energywith extrusionreductionfor differentiron preforms. were selected on the basis of ring-compression tests. Table 2 details the computations made. As can be seen from Fig. 6(a), the forces required for extruding P / M preforms are greater than those required for wrought materials, presumably due to two main reasons: (i) apart from the forces required to deform the material, densification occurs in the case of sintered P / M compacts. This is also evident from the greater experimentally-observed work-hardening exponent na for a sintered iron preform when compared to the wrought equivalent; (ii) far greater internal friction is likely in the case of the deformation of a porous compact. This is also supported by the lower ratio of Fmax/Fobserved in the case of P / M preform extrusions as against the wrought equivalent.

Energy The energy required as a function of deformation ratio e for initial preform densities of 5.9, 6.35 and 6.6 g/cm 3 is shown in Fig. 6 (b). The energy values were computed, using a planimeter, from the area under the force-stroke plots for an identical punch penetration of 20 mm. The greater-density preforms

370 TABLE 2 Calculated values of force Fto t during t h e Hooker extrusion of wrought 0.05% carbon annealed steel at the s t a r t of extrusion* Properties: Ao-- 530.14 ram2; Yo=251 N/ram2; K = 5 3 9 N/ram2; n = 0 . 2 6 ; ~u=0.125; o~=45 °.

L=&L& ~

e = in AoIA1 1Z--A1

E

Ymea. = K E " / ( n + 1 )

Forces ( k N )

( N / r a m 2)

0.60 0.70 0.80 0.90

374.6 389.9 403.7 416.2

Fid

F~h

F~'*

Frs

Srd 1

Srd z

Fto t

119.1 144.7 171.2 198.6

78 81.2 84.0 86.6

133 133 133 133

29.8 36.2 42.8 49.6

9.6 10.1 i0.7 9.94

2.8 2.8 2.8 2.8

372.3 408 444.5 480.5

*Values of Yo, K, n a n d / ~ were d e t e r m i n e d experimentally by t h e ring-compression test on t h i s material, using graphite dispersion in oil as a lubricant, **for l = 45 ram. TABLE 3 Force values ( kN ) during the Hooker extrusion of P / M preforms with Pore= 5.9 g/cm 3 ( compacted with 150 kN) and 0~=45 ° Experimentally derived values: Ao = 530.14 mine; Yo = 112 N/ram2; Ka = 485 N/ram2; na = 0.38; # = 0.14.

2~= 1 2 0 ° ~A ~ I " [~=In A°IA1

e

Y. . . . Fid (N/ram 2 )

F,h

F~

Fr~

F~d~

Frd~

Ftot (calc.)

F, ct

A1 Fact (ram 2) ~ =~

0.442 0.579 0.693 0.916

257.7 285.5 305.7 339.9

53.6 59.4 63.6 70.8

14.8 14.8 14.8 14.8

16.9 24.5 31.4 46.2

5.4 6.9 7.9 9.2

3.1 3.1 3.1 3.1

154.2 196.3 233.1 309.2

315.0 394.0 435.0 600.0

340.87 297:07 265.21 212.13

60.4 87.6 112.3 165.1

2.0 2.0 1.87 1.94 1.95 Average

371 TABLE 4 Force values (kN) during the Hooker extrusion of P / M preforms with po~--6.35 g/cm 3 (compacted with 150 kN) and a---45 ° Experimentally derived values: Ao = 530.14 mm2; Yo -- 121 N/mm2; K a = 510 N/ram2; n~ -- 0.30;/~--=0.14. E

Y~ea. Fid ( N / m m 2)

0.442 0.579 0.693 0.916

307.1 333.0 351.4 382.1

72 102.2 129.1 185.6

F~h

F~

F~

Frd I

63.9 69.3 73.2 79.5

15.97 15.97 15.97 15.97

20.15 6.5 28.6 8.0 36.15 9.0 51.95 10.4

Frd2

Ftot

Fact

A1 Fact (mm ~) F-~-l~= o

3.1 3.1 3.1 3.1

181.6 227.2 266.5 346.5

309.0 397.5 457.5 574.0

340.87 297.07 265.21 212.13

1.70 1.75 1.72 1.66 1.71 Average

TABLE 5 Force values during the Hooker extrusion of P / M preforms with po~=6.6 g/cm 3 (compacted with 350 kN) and ¢z=45 ° Experimentally derived values: Ao -- 530.14 mm2; Yo-- 186 N/ram2; Ka = 554 N/mm2; n~ = 0.28;/z = 0.125. e

Y,~ean Fid ( N / m m 2)

F,h

F,~

F,~

Frdl

Frd~

Ftot (calc.)

Fact

A1 Fact (mm 2) ~ = ¢ ~

0.442 0.579 0.693 0.916

344.4 371.4 390.6 422.3

71.7 77.3 81.3 87.9

21.9 21.9 21.9 21.9

20.2 28.5 35.9 51.3

6.5 8.0 9.0 10.3

2.8 2.8 2.8 2.8

203.8 252.5 294.4 379.3

330.0 405.0 465.0 600.0

340.87 297.07 265.21 212.13

80.7 114.0 143.5 205.1

1.62 1.60 1.58 1.58 1.595 Average

TABLE 6 Force values during the Hooker extrusion of P / M preforms with pore=6.6 g/cm 3 (compacted with 350 kN) and or=60 ° Experimentally derived values of: A o = 530.14 ram2; Yo= 186 N/ram2; Ka = 554 N/ram2; na = 0.28;/L = 0.125.

e

Ymean Fid (N/ram e)

0.442 0.579 0.693 0.916

344.4 371.4 390.6 422.3

F~h

F~

80.7 95.6 21.9 114.0 103.1 21.9 143.5 108.3 21.9 205.1 117.2 21.9

F,~

F,~l

Fr~

Ftot (calc.)

Fact

A1 Fact (mm 2) ~ = ~

23.3 32.9 41.45 59.2

3.75 4.6 5.2 5.9

2.8 2.8 2.8 2.8

228.0 279.3 323.2 412.1

440.0 585.0 640.0 750.0

340.87 297.07 265.21 212.13

1.93 2.09 1.98 1.82 1.96 Average

372

2.0 l

LL

LL° 1.5;

r a m - -

\\

\ \~\\\

1.05.0

6.0

Density(g/cm3)

70

I \ 1 7.86 \,

Fig. 7. Variationof the ratio ¢ = Fact/Fcalcwith preformdensityPorein the Hookerextrusionof iron preforms. require more energy, despite the~ncreased energy required for densification in the case of the lower-density preforms.

Force required for Hooker extrusion of P / M preforms - - An empirical relationship The expression for the total force required in the case of the forward extrusion of hollow bodies has been presented by Lange and is dealt with in eqn. (1). Values for the force for P / M sintered iron preforms extruded under various extrusion parameters, calculated using Lange's expression [ 10 ], have been tabulated in Tables 3-6. The values of the calculated forces are seen to be far less than the actual values obtained experimentally. The ratio of Fac t t o Fcalc , termed correction factor 0, has also been indicated in the tables. It is seen that the values of decrease from 1.95 for an initial preform density pore--5.9 g/cm 3 to 1.71 for Pore-- 6.35 g/cm 3 and to 1.595 for Pom----6.6 g/cm 3. For the same initial preform density Pom= 6.6 g/cm 3, the values of ~ increase from 1.595 to 1.96 when the extrusion die angle ~ is increased from 45 ° to 60 °. The variation of ~ with initial preform density Poreis seen in Fig. 7. Extrapolation of the straight-line relationship obtained to the theoretical density ( 7.86 g/cm 3) gives a value of ~=1. Conclusions

From the above results, the following can be concluded: (1) The experimentally derived values of forces are greater than the calculated values, although the calculated values of force are based on experimen-

373

tally derived values of/~, na and Ka. This is because the densification which occurs during the Hooker extrusion process is different from that which takes place during the ring-compression test. Densification is of a much greater magnitude and rate during the Hooker extrusion process, where far greater restraint to the movement of material exists as against the free ring-compression test. This beneficial effect of lateral restraint in increasing densification in the case of Hooker extrusion is primarily due to the side walls of the die container and the die angle. When a is increased from 45 ° to 60 °, ratio 0 increases from 1.595 to 1.96 for the same initial preform density (P0m----6.6 g/cm3). Lange's expression for the estimation of the force required for wrought extrusion does not account for densification and hence greater values are found in practice than those calculated. (2) The value of 0 does not vary appreciably with extrusion ratio e for the same preform density (P0m) • Hence for all extrusion ratios e, the same proportional increase in force over the calculated values is observed. (3) The relationship between 0 and Poreis observed to be linear, as seen in Fig. 7 for P / M extrusions carried out with a die angle a = 45 °: the force required for the densification of preforms of lesser initial density is proportionally greater. Extrapolation of the line to the theoretical density gives a value of 0 = f act/Fcalc -~ 1. (4) ~ c a n be r e g a r d e d as a c o r r e c t i o n f a c t o r w h i c h c a n be a p p l i e d to L a n g e ' s e x p r e s s i o n for force. Hence Ftot : (Rid -[-Fsh

+Ffr)O

(11)

where 0 is the correction factor for sintered iron preforms and is primarily a function of the initial preform density and the die angle a, for a particular Hooker extrusion arrangement. The correction factor reduces to unity for an initial preform density equal to the theoretical density of 7.86 g/cm 3, when the expression of Lange can be applied directly.

References 1 H.R. Biehl and D.T. Smith, U.S. Patent 3060560, 1962. 2 H. Hoeness, W. Kramer, P.S. Raghupathi and H. Wilky, Industrie Anzeiger, 93 (101) (1971) 2563-2567. 3 T. Nakagawa, T. Amano, K. Obara, Y. Nishino and Y. Meda, On the cold forging of sintered iron powder preforms, Proc. 13th Int. Mach. Tool Des. Res. Conf., Birmingham, Sept. 1972, Macmillan, London, 1973, pp. 455-461. 4 R.J. Dower and G.H. Miles, The cold extrusion of mild steel billets produced by powder metallurgical techniques, Mod. Dev. Powder Metall., 7 (1974) 175-201. 5 K. Obara, Y. Nishino and Y. Salt, The cold forging of ferrous P/M preforms, Mod. Dev. Powder Metall., 7 {1974) 423-440.

374 6 7 8

9 10 11 12

13

C.I.R.P. Production Engineering Dictionary, Vol. 5, Cold Forging and Cold Extrusion, Verlag W. Giardet, Essen, 1969, Definition Number 5126. P. Venugopal, S. Venkatraman, R. Vasudevan and K.A. Padmanabhan, Ring compression of sintered iron preforms, J. Mech. Working Technol., in press. P. Venugopal, S. Venkatraman, R. Vasudevan and K.A. Padmanabhan, On the properties and economics of sintered iron powder metallurgical extrudes, J. Mech. Working Technol., in press. S. Venkatraman, Some studies on cold extrusion of sintered iron preforms by the Hooker process, M.S. Thesis, IIT, Madras, 1983. K. Lange, Text Book of Metal Forming, Vol. 2: Bulk Forming, Springer-Verlag, Berlin, 1974, pp. 243,244, 246, 247 (in German). H. Ll.D. Pugh, Hydrostatic extrusion of difficult metals, Sheet Metal Ind., 42 ( 1965) 572. K. Lange and E. Steck, Some recent results of investigations in rod Hooker and can extrusion and upsetting, Paper presented at the Third Int. Conf. on Cold Extrusion of Steel, The Winter Gardens, Eastbourne, November 2-4, 1965. P. Venugopal, S. Venkatraman, R. Vasudevan and K.A. Padmanabhan, Some failure studies in the Hooker extrusion of sintered iron powder metallurgical preforms, J. Mech. Working Technol., in press.