Fatigue failure of cold forging tooling: causes and possible solutions through fatigue analysis

Journal of

ELSEVIER

Journal of Materials Processing Technology 46 (1994) 57-71

Materials Processing Technology

Fatigue failure of cold forging tooling: causes and possible solutions through fatigue analysis Markus Knoerr a'*, Kurt L a n g e b, Taylan A l t a n a "Engineering Research Center for Net Shape Manufacturing, The Ohio State University, 339 Baker Systems, 1971 Neil Avenue, Columbus, OH, USA blnstitut far Umformtechnik, Universitdt Stuttgart, Stuttgart, Germany

Industrial Summary

Fatigue failure is a major failure mode in cold forging of complex parts with net shaped surfaces. Tooling costs contribute significantly to the overall cost of a cold forging. It is, therefore, desirable to have the ability to estimate possible tool life during the process and tool design phase. Thus, necessary measures to increase insufficient tool life can be taken early on before actual tooling components are manufactured and pre-production tryouts are conducted. This paper discusses the causes of fatigue failure in cold forging tooling and presents a fatigue analysis concept that can be applied during process and tool design to estimate the tool life of a layout. The concept has been evaluated using the tool life experiments conducted at the Institute for Metal Forming at the University of Stuttgart and an industrial case. The tool life predictions compare favorably with the tool life experienced in the experiments and in industrial production. Material data for two standard cold forging tool steels, which is necessary to perform the fatigue analysis, are also provided. Keywords: Cold forging; Fatigue failure of tooling; Tool life estimation

1. I n t r o d u c ~ o n In t he p r o d u c t i o n of h i g h - v o l u m e cold forged p a r t s w i t h n e t - s h a p e d complex surf a c es , f a t i g u e c r a c k i n g of t h e active tool e l e m e n t s is t h e l e a d i n g c a u s e of fa i l ur e [1]. Th e tool life is g e n e r a l l y m u c h s h o r t e r t h a n in t h e cases of f ai l u r e by w e a r . A n u m b e r of e x a m p l e s h a v e been p r e s e n t e d in t h e l i t e r a t u r e . F i g u r e 1 shows some r e p r e s e n t a t i v e cases. In f o r w a r d e x t r u s i o n , as s h o w n in F i g u r e l a , f a t i g u e cracks i n i t i a t e a t t h e t r a n s i t i o n r a d i u s to t h e e x t r u s i o n s h o u l d e r a n d p r o p a g a t e in t h e r a d i a l d i r e c t i o n . L o n g i t u d i n a l c r a c k s u s u a l l y r e s u l t f r o m o v e r l o a d i n g a n d occur d u r i n g t h e first few loading cycles of a tool. This type of f ai l u r e can be easily avoided by proper design of t h e s h r i n k fit. F i g u r e l b s h o ws a f a t i g u e crack e x p e r i e n c e d in t h e u p p e r p u n c h of a tool a s s e m b l y u s e d in cold f o r g i n g of d i f f e r e n t i a l side g e a r s . T h e f a t i g u e c r a c k i n i t i a t e s at t h e t r a n s i t i o n r a d i u s to t h e p u n c h l a n d an d p r o p a g a t e s from t h e radius. More detailed i n f o r m a t i o n can be found in [2].

*Corresponding author. Elsevier Science S.A. SSDI 0924-0136(94)01566-J

M. Knoerr et aL / Journal of Materials Processing Technology 46 (1994) 57-71

58

Overload Crack Ring

Fatigt

a)

b)

OANFOSS

- - 5 0

~

A. Corner in d i e - i n s e r t

D. STRECON® Prestressed Container

// 1m m

{:)epth=6.Smm

B. Crock in d]e-inser~ c o r n e r

c)

d)

Figure 1. Ca) Schematic presentation of failures in forward extrusion [4}; (b) Fatigue failure of u p p e r p u n c h in cold forging of differential side gears; (c) F a t i g u e failure of tool i n s e r t for radial e x t r u s i o n of helical g e a r s [2]; (d) F a t i g u e failure in n o n - a x i s y m m e t r i c b a c k w a r d extrusion die {3].

M. Knoerr et aL / Journal of Materials Processing Technology 46 (1994) 57-71

Figure lc shows a fatigue crack initiating in the root of a helical gear tool insert as reported by Schrnieder [3]. Figure ld shows a fatigue crack that occurs in the transition radius of an octagonally shaped backward extrusion die as reported by the companies Danfoss and Presta and presented by Lange et al. [4]. In this geometry the fatigue crack propagates longitudinally from the transition radius. A fifthcase is discussed by Nagao et al. [6] in this publication. 2. F a t i g u e Analysis Fatigue Analysis is a common design technique used in the evaluation of components for automotive applications [7]. It consists of the following steps: • • •

Determination of loading conditions, Stress-strain analysis to determine stress-strain response in the zone of highest loading in the component, Damage analysis to estimate cycles to failure.

This general approach has been modified for the analysis of forging tooling. Figure 2 shows the Fatigue Analysis Concept for forging tooling developed at the ERC/NSM. Process and Tooling Design

Material Data

[

I

I

Tool Geometry

Process

Conditions

Flow Curve of

FEM Process Simulation

Workpiece Material

+

Thermal and Mechanical Material Data of Workpiece and Tooling

Tool Loads

~-

Cyclic (or static) Stress-Strain Curve of Tooling Material

Elastic-plastic Stress-Strain Analysis to determine the maximum Strain Amplitude

I

Strain - Life Data from Fatigue Tests with Smooth Specimens

(or estimated from static or cyclic Stress-Strain Curve )

Damage Analysis

Service Life to ~1

Crack Initiation

Figure 2. FatigueAnalysisConcept for ForgingTooling.

N=

I

59

60

M. Knoerr et al. / Journal of Materials Processing Technology 46 (1994) 57-71

It is essential for the fatigue analysis to have detailed loading d a t a for the tooling. This d a t a can only be determined with reasonable accuracy by means of a F i n i t e E l e m e n t (FE) based process simulation. The input d a t a for the process analysis is determined from a preliminary process and tooling design. The process analysis delivers the tool load in the form of the contact stress distribution at the die-workpiece interface. This data is subsequently used as load input for the stress-strain analysis. An elastic-plastic stress-strain analysis is performed for two loading cycles in order to determine the cyclic response of the tooling at the highest loaded zone. The total strain amplitude is calculated. The stress-strain curve of the material is needed for the analysis. The local s t r a i n a p p r o a c h is then used d u r i n g the d a m a g e a n a l y s i s to correlate the total strain amplitude with the number of cycles to failure. Strainlife d a t a for the material of the active toot component must be available. The fatigue analysis concept has been implemented at the ERC/NSM using commercial software, when available. The 2-D forging process s i m u l a t i o n package DEFORM and the general purpose FE code ABAQUS are being used. A customized d a t a exchange program TRANSFER was developed to automate the d a t a exchange from the process simulation to the s t r e s s - s t r a i n analysis. TRANSFER has been described in [2]. The fatigue analysis concept was e v a l u a t e d with help of a n u m b e r of tool failures reported in the literature and provided by industry. The results of two of these evaluations will be presented later in this paper. The process s i m u l a t i o n and stress analysis t a s k s of the fatigue a n a l y s i s concept were used to investigate the s t r e s s - s t r a i n state and the cyclic stresss t r a i n response of the tool material in order to explain the causes of failure. The necessary material d a t a was found in the literature.

3. Analysis of two fatigue failures E l a s t i c - p l a s t i c s t r e s s - s t r a i n a n a l y s i s was performed for the two f a t i g u e failures shown in F i g u r e l a and lb. Crack initiation and propagation were investigated for the forward extrusion die insert and for the upper punch used in forging of differential side gears. The two cases are used to explain the causes of crack initiation and propagation in cold forging dies.

3.1. F o r w a r d Extrusion Two i n v e s t i g a t i o n s performed at the I n s t i t u t e for Metal F o r m i n g at the University of S t u t t g a r t by Reiss [5] and Hettig [6] provide detailed information on tooling m a t e r i a l s for cold forging and on fatigue failure. H e t t i g monitored the crack initiation and propagation in forward extrusion dies using two different die materials, two transition radii and two different die opening angles. Thus, a n u m b e r of different s t r e s s - s t r a i n responses were induced in the critical area of the tool inserts at the transition radius. For each geometry/die m a t e r i a l combination five die inserts were tested and monitored to obtain the number of cycles to initiation and the crack propagation path. F i g u r e 3 shows the tool layout of the forward extrusion tooling and indicates the p a r a m e t e r s t h a t were varied during the investigation. Table 1 gives an overview of the p a r a m e t e r combinations that were investigated with the fatigue analysis concept. The table also lists the experimentally determined insert life to crack initiation.

M. Knoerr et aLI Journal of Materials Processing Technology 46 (1994) 57-71 17.7

.~rt Materials: AISI M2 AISI D2 rference: 4.8 %0 /

in m m 8.75 120

=

Figure 3. Layout of tooling and calculated material flow used for die life investigations at Institute for Metal Forming, University of Stuttgart [8].

Table 1. Parameters of analyzed test cases for the forward extrusion investigation. Case

Insert Material

Hardness [HRC]

Die opening angle 2c( [°]

Transition radius R [rnm]

Insert life to crack initiation [8]

I

AISI M2

61

120

1

50 - 400

II

AISI D2

60

120

1

65 - 200

III

AISI D2

60

90

1

9 0 0 - 1000

IV

AISI D2

60

90

2,5

10.000 - 11.000

3.1.1 Process S i m u l a t i o n Process s i m u l a t i o n s w e r e p e r f o r m e d for th e t h r e e g e o m e t r i c a l v a r i a t i o n s (2cc = 1 2 0 ° / R = i ram, 2cc = 9 0 ° / R = 1 m m , 2 a = 9 0 ° / R = 2.5 m m ) listed in Tab l e 1 to d e t e r m i n e t h e contact s t r e s s distributions. T h e predicted m a t e r i a l flow for t h e v a r i a t i o n 2 a = 1 2 0 ° / R = 1 m m is shown in F i g u r e 3. F i g u r e 4 shows t h e n o r m a l c onta c t s t r e s s d i s t r i b u t i o n a t t h e die-workpiece i n t e r f a c e for t h e two d i f f e r e n t die o p e n i n g angles and a t r a n s i t i o n r a d i u s R = 1 ram. Th e s h a l l o w e r o p e n i n g angle 2 a = 120 ° produces a h i g h e r tool load. 3.1.2.

Die S t r e s s an d D a m a g e A n a l y s i s

Elastic-plastic die stress analysis was performed for three points over two loading cycles: •

• •

initial m a x i m u m load point, unloading after the first loading cycle and m a x i m u m load point for subsequent loading cycles.

61

62

M. Knoerr et aL / Journal of Materials Processing Technology 46 (1994) 57-71

2500

P

|Workpiece Material: AIS14320 (BS970)

~_

I

l,

i

~:

D~e shoulder

....................................................... Tran"on" 15oo

.............

............................... i............................... i............................... i..........................

1000

.................

0

z 0 0

5

10

15

20

25

Length of contact from top of billet in m m

Figure 4. Normal contact stress distribution for the geometric variations 2a = 120°/R = 1 mm and 2a = 90°/R = I mm. It w a s found t h a t t h e s e t h r e e load points are sufficient to e s t i m a t e t h e cyclic b e h a v i o r of t h e i n s e r t m a t e r i a l at t h e h i g h e s t loaded zone. D u e to t h e elastic d e f o r m a t i o n of t h e bulk cross section of t h e inserts, t h e m a t e r i a l b e h a v i o r in t h e t r a n s i t i o n r a d i u s is s t r a i n c o n t r o l l e d d u r i n g t h e l o a d i n g cycles. T h u s , t h e h y s t e r e s i s plot of t h e s t r e s s - s t r a i n response will a l w a y s cycle w i t h i n t h e s a m e s t r a i n b o u n d a r i e s . F i g u r e 6 shows t h e c a l c u l a t e d h y s t e r e s i s plot for t h e i n v e s t i g a t e d Case I, w h e r e t h e s t r a i n r a n g e is predicted to be 0.60 %. T h e elastic-plastic m a t e r i a l d a t a used for t h e s i m u l a t i o n s w e r e t a k e n from t h e l i t e r a t u r e a n d are shown in F i g u r e 5.

35oo ~.

!

3000

i

i

.......................................................................

~E

._

2500 2000 15oo

1000 w

500

o I 0.00

~ 0.05

s

l

D2, 62HRC, R~,, IS]

0.10

0.15

0.20

0.25

0.30

Engineering Strain

Figure 5. Stress-strain curves for insert materials AISI M2 and AISI D2.

M. Knoerr et a L I Journal of Materials Processing Technology 46 (1994) 57-71

3.1.2.1 Case I: I n s e r t M a t e r i a l AISI M2 The s t r e s s - s t r a i n h y s t e r e s i s plot calculated for Case I is shown i n F i g u r e 6. T h e a n a l y s i s of two l o a d i n g cycles predicts a total s t r a i n a m p l i t u d e of eat = 0.0060 (0.60 %), which is equal to h a l f of the s t r a i n r a n g e A£t. The plastic zone in the t r a n s i t i o n r a d i u s is shown in Figure 7.

2000 J// // . Material : AISIM2 ................................................ ::Z.),/7./~ ................ i ...... Hardness: 62 HRC o i 7'/ ]/ i Geometry: 2a = 120 ,

1500 1000 (D

.=__ -o rr ._~

500

{D ~D

q~l

-500

I~-

Ae = 0.0016

/

..........~! .........!-~...........P.....................................1......... Calculated ;iAet = 0.0119i

Total Strain Amplitude: Eat11 = 0.0060 (0,60%)

~

-1000 0

0.005

0.01

0.015

S t r a i n in R a d i u s

0.02

0.025

ell

Figure 6. Calculatedstress-strainresponse in the transitionradius for insert material AISI M2 and geometry 2a = 120°/R = I ram.

T h e total strain range value is then used for the d a m a g e analysis o estimate the insert life to crack initiation. Kocanda [i0] provides strain-life data for AISI M 2 high hpeed hteel at room temperature and elevated temperatures. The room temperature strain-life data for a heat treatment with a hardening temperature of 9 H = 1150°C and tempering of t~T = 2x 560°C was used for the d a m a g e analysis, since this heat treatment compares best with the heat treatment applied by Hettig [8] (gH = 1165°C, ~ T = 555/570/560°C) in his experiments. Figure 8 shows the results of the d a m a g e analysis. Insert life for Case I was predicted to be approximately 280 parts to crack initiation. Hettig [8] reports a monitored insert life of 50 - 400 for 5 inserts. The predicted insert life compares well with the actual life experienced in the experiments.

3.1.2.2 Cases II - IV: I n s e r t M a t e r i a l AISI D2 These t h r e e cases, listed i n Table 1, were a n a l y z e d to i n v e s t i g a t e the effect of g e o m e t r i c v a r i a t i o n s on t h e s t a t e of stress i n the t r a n s i t i o n r a d i u s . P l a s t i c d e f o r m a t i o n i n t h e t r a n s i t i o n r a d i u s was f o u n d for C a s e II (2a = 120°/

63

64

M. Knoerr et aL / Journal of Materials Processing Technology 46 (1994) 57-71

Insert Material: AISI M2, 62 HRC Geometry: 2 a = 120 °, R = l m m

Plastic Zone ~vmax = 3005 N/mm 2

Ov = 2700 N/mm 2 11

22.2

Z

Figure 7. Plastic zone in the transition radius for insert material AISI M 2 26 = 120°/R = I m m shown by the equivalent (von Mises) stress distribution.

and geometry

0.1 ~

i

Insert Material : Heat Treatment:

~J

S 6-5-2 (AISI M2) ~H = 1150°C

---<>--- ~T = 2 x 580°C, TTesl = 350°C ~T = 2 x 560°C, TTest = 20°C •

Q_

E < e-

~

c

e

OT = 2 x 580°C, TTest = 20°C :

Kocanda [10]

0.01 Ca,cu,ate

Insert Life Range in Experiments " ~ i by Hettig [8] "~

b-

0.001

i

1

i

i ' l l

I

10

I

I

Predicted Insert Life: 280 Parts II i

100

k

[

I

1000

IIIII;

. . . . . . .

104

105

C y c l e s / F o r g e d Parts to C r a c k Initiation

Figure 8. Damage Analysis to predict insert life for Case I.

R = 1 ram) an d C a s e I I I (2a = 1 2 0 ° / R l o w e s t tool load of t h e e x p e r i m e n t s , t r a n s i t i o n radius. T a b l e 2 s u m m a r i z e s r e s p o n s e for t h e a n a l y z e d cases. T h e

= 1 ram). Case IV, w h i c h r e p r e s e n t s t h e did n o t show a n y p l a s t i c zone in t h e t h e a n a l y s i s r e s u l t s a n d cyclic m a t e r i a l n o r m a l c on t act s t r e s s p e a k •nmax, t h e

65

M. Knoerr et a L I Journal of Materials Processing Technology 46 (1994) 57-71 Table 2. S u m m a r y of analyzed results for insert material AISI D2.

Geometry

(~nm~ [N/mm 2]

[N/mm 2]

Eat11 [%]

eap11 [%]

II

2a=120°/R=1mm

2150

2810

0,66

0,135

446

III

2a=90°/R=lmm

2014

2791

0,54

0,023

408

IV

2o~=90°/R=2,5mm

1973

2068

0,33

Case

Ovmax

max. Load measured [kN] Load [kN] [8]

390

450

377

e~,l~) m a x i m u m e q u i v a l e n t s t r e s s C ~ v m aaxn d t h e total (e~,11) . . a .n d . plastic . . - - s t r a.i n a m p l i t u d e a r e listed. T a b l e 2 also compares t h e p r e d i c t e d m a x i m u m load wxth m e a s u r e d v a l u e s p r o v i d e d by H e t t i g [8]. These v a l u e s also c o m p a r e well. I n o r d e r to e v a l u a t e the s t r a i n - l i f e b e h a v i o r of t h e A I S I D2 tool steel, t h e p l a s t i c a n d t o t a l s t r a i n a m p l i t u d e s for t h e four c a s e s w e r e p l o t t e d over t h e i n s e r t life d e t e r m i n e d e x p e r i m e n t a l l y by H e t t i g [8]. F i g u r e 9 shows t h e s t r a i n life b e h a v i o r for A I S I D2. F o r c o m p a r i s o n p u r p o s e s F i g u r e 9 also shows t h e s t r a i n - l i f e b e h a v i o r of t h e A I S I M2 h i g h s p e e d steel. The e s t i m a t e d s t r a i n - l i f e b e h a v i o r of t h e A I S I D2 cold forging steel is s i m i l a r to t h e b e h a v i o r of t h e A I S I M2 h i g h s p e e d steel. T h e c a l c u l a t e d d a t a c a n be u s e d for f u t u r e d a m a g e analysis. 0.1 \

AISI M2, T-/est. = 350°C

\

~ .

AISIM2,

TTest = 20°C m =

"O

'¢~ ¢"~

0.01

.......................... :~ .............---------..........-~ ........:L.........i..................................................... \ ap ",.

AISI D2: Strain Ampl.

"~ 0.001

~...

- A i ' S i - M 2 : I , ...............................T ' ~ ~ = 1150 C H ' •3 A = 2x580°C

i i i

Source: Kocanda [10] 10 -4

T i iiiii

1

i

10

i

from FE-Simulations ....... Ni-Valuesfrom

\ ~ experiments by Hettig [8] -. . " \ ~ i

i

i ,1111:

100

"~, i

i

i Ilrll

!

i i

1000

I

i

tlITir

!

i

104

C y c l e s / F o r g e d P a r t s to C r a c k Initiation

105 N.

I

Figure 9. Strain-life data for AISI D2 as calculated from experiments by Hettig and AISI M2 as reported by Kocanda [10]. T h e s t r e s s s t a t e a t t h e t r a n s i t i o n r a d i u s for C a s e I I is s h o w n in F i g u r e s 10 a n d 11. F i g u r e 10 shows t h e location a n d direction of t h e c r a c k in a die i n s e r t as r e p o r t e d by H e t t i g [8]. The location of t h e crack i n i t i a t i o n c o r r e s p o n d s w i t h t h e l o c a t i o n of t h e h i g h e s t e q u i v a l e n t s t r e s s in t h e i n s e r t u n d e r m a x i m u m load. I t is a l s o t h e l o c a t i o n of t h e h i g h e s t m a x i m u m p r i n c i p l e s t r e s s , as s h o w n i n F i g u r e 10. T h e c r a c k t h e n p r o p a g a t e s p e r p e n d i c u l a r to t h e p l a n e s of t h e maximum principle stress at the transition radius.

66

M. Knoerr et a L I Journal of Materials Processing Technology 46 (1994) 57-71

2000

Material: AISID2 Hardness: 60 HRC • Geometry: 2ct=120°, R=lmm

N/mm 2 13.. t-

1000 500

i

"10

n-

0

09

Calculated i StrainAmplitudes: ......if Cat11=0,0066 (0,66%)

-500 AEtl I =0~0132

i ¢apll

-1000 0

a)

0.005

=0,0016 (0,16%)

0.01 0.015 Strain in Radius Ell

0.02

0.025

I

I I

I

I

= 1969 N / m m 2

/

Crack growth in insert as reported by Hettig [8] Material: AISI D2 Geometry: 1

~ ./P ~ "

t

~o~//

b)

~

'"

/

,I

I

/ /

/

/

II

I /

r

/

Figure 10. (a) Calculated s~ess-st~am response in the transition radius for insert material AIS[ D 2 and geometry 2a = 120" / R = 1 rnm; Co) Crack growth and m a x i m u m principle stress distribution in the transition radius.

M. Knoerr et aL / Journal of Materials Processing Technology 46 (1994) 57-71

I J

f

I I

i

J f f J

"1t

I-----

r

i t L

!

J ! tll

",t

t

Figure 11. Three-dimensional presentation of the maximum principle stress distribution near the transition radius.

Figure 11 shows a three-dimensional view of the maximum principle stresses. In the greater part of the insert the maximum principle stress are compressive and their direction is tangential. This is due to the effect of the shrink ring on the insert. However, during the forging process the maximum principle stresses near the transition radius change their direction. They become tangential to the radius and reach high tensile values.

3.2. Upper Punch Failure The upper punch failure shown in Figure lb was investigated in a previous study. The full details of the analysis of the fatigue failure are published in [2]. The analysis showed a similar material behavior in the transition radius, where the crack initiates, as seen in the forward extrusion inserts. Figure 12 shows the plastic strain distribution in radial direction. The highest plastic strain is experienced at the crack initiation site in the radius. Figure 13 shows the maximum principle stress distribution at the maximum load point during a forging cycle. The principle stresses are mainly compressive and in the tangential direction, as can be seen in the three-

67

68

M. Knoerr et aL / Journal o f Materials Processing Technology 46 (1994) 57-71

J

0.88%

Fig. 12. Plastic strain in radial direction.

d i m e n s i o n a l p r e s e n t a t i o n shown in F i g u r e 13a. I n t h e a r e a of t h e t r a n s i t i o n r a d i u s t h e principle s t r e s s e s c h a n g e direction into t h e r-z p l a n e a n d r e a c h h i g h t e n s i l e values. A d a m a g e a n a l y s i s w a s p e r f o r m e d for t h e punch, which is m a d e from A I S I M2 h i g h s p e e d steel h e a t t r e a t e d to a h a r d n e s s of 65 - 66 HRC. The d a m a g e a n a l y s i s , w h i c h u s e d t h e s t r a i n - l i f e d a t a for a n A I S I M2 steel h a r d e n e d to 61 - 62 HRC p r o v i d e d by K o c a n d a [10], p r e d i c t e d a tool life to i n i t i a t i o n of a p p r o x i m a t e l y 12,000 parts. The life in production was close to 10,000 parts. It is a s s u m e d t h a t t h e d a m a g e a n a l y s i s o v e r e s t i m a t e d t h e tool life to i n i t i a t i o n b e c a u s e m a t e r i a l d a t a for a softer, more ductile h e a t t r e a t m e n t was used.

4. C a u s e s of fatigue f a i l u r e i n c o l d f o r g i n g t o o l i n g T h e a n a l y s i s of t h e two f a i l u r e cases c l e a r l y p o i n t to t h e c a u s e s of f a t i g u e f a i l u r e in cold forging tooling. Two i m p o r t a n t a s p e c t s t h a t l e a d to t h e f a t i g u e cracks can be identified: C r a c k i n i t i a t i o n occurs in t r a n s i t i o n radii, if t h e tool load exceeds t h e yield s t r e n g t h of t h e tool m a t e r i a l a n d a localized p l a s t i c zone forms in t h e r a d i u s . T h i s zone g e n e r a l l y forms d u r i n g t h e f i r s t l o a d i n g cycle a n d u n d e r g o e s p l a s t i c cycling d u r i n g s u b s e q u e n t u n l o a d i n g a n d r e l o a d i n g . T h e p l a s t i c cycling l e a d s to t h e i n i t i a t i o n of m i c r o s c o p i c c r a c k s . T h e location of crack i n i t i a t i o n can be t r a c e d to t h e location of t h e plastic zone.

Fig 13a. Maximum principle stress distribution in three-dimensional presentation. Maximum compressive value at tip: 2634 MPa

y

i

j f

J

~ -

Fig. 13b Maximum principle stress distribution in r-z plane. Maximum tensile value: 1532MPa

0

7

t~

v,

7O

M. Knoerr et al. / Journal o f Materials Processing Technology 46 (1994) 57-71

The maximum principle stresses significantly change their magnitude a n d d i r e c t i o n n e a r t r a n s i t i o n r a d i i u n d e r t h e f o r m i n g load. Before loading, t h e y u s u a l l y a r e compressive a n d in t a n g e n t i a l d i r e c t i o n due to t h e p r e s t r e s s i m p o s e d by the s h r i n k ring. U n d e r t h e f o r m i n g load, t h e maximum principle stresses near the transition radius change their d i r e c t i o n into t h e r-z p l a n e a n d r e a c h h i g h t e n s i l e v a l u e s . This effect c a u s e s t h e m i c r o s c o p i c c r a c k s to grow a n d l e a d s to t h e s u b s e q u e n t p r o p a g a t i o n of the cracks into t h e cross section of the tooling.

5. P o s s i b l e s o l u t i o n s

F a t i g u e a n a l y s i s is a p p l i e d to d e t e r m i n e t h e life of a tooling design. Should t h e p r e d i c t e d tool life be insufficient, c h a n g e s in t h e process a n d tooling d e s i g n m u s t m a d e to r e d u c e the l o a d i n g conditions. A s i g n i f i c a n t i n c r e a s e can be a c h i e v e d by r e d u c i n g t h e s t r e s s e s in t h e h i g h e s t l o a d e d zone below t h e y i e l d s t r e n g t h of the tool m a t e r i a l . This m a y be achieved by a n u m b e r of m e a s u r e s : • • • • • •

C h a n g e m a t e r i a l flow in t h e die to reduce the contact s t r e s s e s on the tool; R e d e s i g n t h e process to avoid d r a s t i c c h a n g e s in t h e d i r e c t i o n of t h e m a t e r i a l flow, which u s u a l l y leads to p e a k s in the contact stress; I n c r e a s e t h e t r a n s i t i o n r a d i i to reduce t h e notch effect; Split t h e tooling at the h i g h e s t loaded zone; I n c r e a s e t h e i n t e r f e r e n c e of t h e s t r e s s ring, if possible; A p p l y a d v a n c e d s t r e s s r i n g techniques, such as s t r i p - w o u n d c o n t a i n e r s or profiled s t r e s s rings.

6. C o n c l u s i o n s

A c o m p u t e r - a i d e d f a t i g u e a n a l y s i s concept for cold forging tooling h a s been introduced. The fatigue a n a l y s i s consists of t h r e e tasks: • •

•

F E - b a s e d process s i m u l a t i o n , F E - b a s e d e l a s t i c - p l a s t i c s t r e s s - s t r a i n analysis, a p p l i c a t i o n of t h e local s t r a i n a p p r o a c h to d a m a g e a n a l y s i s , w h i c h provides a n e s t i m a t e of the tool life to crack initiation.

T h e c o m p u t e r - a i d e d a n a l y s i s t e c h n i q u e s were a p p l i e d to a n a l y z e two f a t i g u e failures, r e p o r t e d in t h e l i t e r a t u r e a n d provided by i n d u s t r y . Both cases show i d e n t i c a l c a u s e s l e a d i n g to t h e i n i t i a t i o n a n d p r o p a g a t i o n of t h e fatigue cracks. The causes are: = •

f o r m a t i o n of a p l a s t i c zone in t h e tooling t h a t l e a d s to t h e f o r m a t i o n of microscopic cracks due to t h e cyclic loading, a n d a maximum principle stress state near the transition radius that e n h a n c e s the f o r m a t i o n a n d p r o p a g a t i o n of fatigue cracks.

It w a s shown t h a t fatigue a n a l y s i s is a c a p a b l e tool for t h e p r e d i c t i o n of tool life. H o w e v e r , m o r e d e t a i l e d m a t e r i a l d a t a is n e c e s s a r y for t h e c o m m o n tool steels a n d h e a t t r e a t m e n t s used in cold forging. O v e r a l l , it is s u f f i c i e n t to d e t e r m i n e life to c r a c k i n i t i a t i o n , s i n c e it is i m p o r t a n t to d e s i g n cold forging tooling in a w a y t h a t t h e i n i t i a t i o n p o i n t is

M. Knoerr et al. I Journal of Materials Processing Technology 46 (1994) 57-71 p u s h e d i n t o a life r a n g e w h e r e tool w e a r will l e a d to t h e e n d of t h e tool s e r v i c e life. T h i s a p p r o a c h g u a r a n t e e s r o b u s t p r o d u c t i o n c o n d i t i o n s , s i n c e tool s e r v i c e life c a n n o w b e m o n i t o r e d b y m e a n s of s t a t i s t i c a l p r o c e s s c o n t r o l .

l:~aferenoes [1]

iange, K. , Umformtechnik - Handbuch fiir Industrie u. Wissenschaft, 2. Aufl. Bd. 2: Massivumformung. Berlin Heidelberg New York: Springer 1984.

[2]

Lange, K., Hettig, A. and Knoerr, M., Increasing Tool Life in Cold Forging through Advanced Design and Tool Manufacturing Techniques, J. Mat. Proc. Techn. 35 (1992) 3-4.

[3]

Schmieder, F, Beitrag zur Fertigung yon sehriigverzahnten Stirnriidern durch QuerflieBpressen. Berichte aus dem Institut ftir Umformtechnik der Universitiit Stuttgart Nr. 118. Berlin Heidelberg New York: Springer 1992.

[4]

Lange, K., Cs6r, L., Geiger, M., Kals, J.A.G., Hansel, M., Tool Life and Tool Quality in Bulk Metal Forming, Annals of the CIRP Vol. 4/211992, 667-675.

[5]

Reiss, W., Untersuchung des Werkzeugbruches beim Voll-Vorwiirts-FlieBpressen. Berichte aus dem Institut ftir Umformtechnik, Universittit Stuttgart, Nr. 94. Berlin Heidelberg New York: Springer 1987.

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Nagao, Y., Knoerr, M. and Altan, T., Improvement of Tool Life in Cold Forging of Complex Automotive Parts, J. Mat. Proc. Techn. 46 (1994) 73-85.

[7]

SAE: Fatigue Design Handbook AE-10. Warrendale, P.A.: Society of Automotive Engineers 1988.

Is]

Hettig, A., EinfluBgrSflen a u f den W e r k z e u g b r u e h beim Voll-Vorw~irtsFliel3pressen, Berichte aus dem Institut fiir Umformtechnik, Universit~it Stuttgart, Nr. 106, Berlin Heidelberg New York: Springer 1990.

[9]

Neitzert, T., Auslegung von rotationssymmetrischen FlieBpreBwerkzeugen im Bereich elastiseh-plastischen Werkstoffverhaltens, Berichte aus dem I n s t i t u t ftir Umformtechnik, Universit~it Stuttgart, Nr. 63, Berlin Heidelberg New York: Springer 1982.

[lOl

Kocanda, A., Die Steel for Warm Working - An Evaluation of Resistance to Cyclic Loading, Advanced Technology of Plasticity 1990, Vol. 1, 349-354.

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