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NASA Technical Memorandum 100796

Relationship Between Fatigue Life in the Creep-Fatigue Region and Stress-Strain Response

(liASA-TB-1003$6) BELBTICESGXF BETPEELS N88- 1 60 40 kALTIGDE L I F E Xi T E E C O E X P - f B ! i f G U E R E G I O N A N D 5!IBESS-STBAII BESFCNSE ( E A S A ) 2 1 pCSCL 20s

Unclas 63/39 0125889

A. Berkovits Lewis Research Center Cleveland, Ohio

and

S. Nadiv Technion-Israel Institute of Technology HaifQ, Israel

Prepared for the VI International Congress on Experimental Mechanics sponsored by the Society of Experimental Mechanics Portland, Oregon, June 5-10, 1988

https://ntrs.nasa.gov/search.jsp?R=19880008656 2020-02-29T20:36:13+00:00Z

RELATIONSHIP BETWEEN FATIGUE LIFE IN THE CREEP-FATIGUE

REGION AND STRESS-STRAIN RESPONSE

A. Berkovi ts* National Aeronautics and Space Administration

Lewis Research Center Cleveland, Ohio 44135

and

S. Nadivt

Department of Materials Engineering Haifa, Israel

Technion - Israel Institute of Technology

ABSTRACT

On the basis of mechanical tests and metallographic studies, strainrange

partitioned lives were predicted by introducing stress-strain materials

parameters into the Universal Slopes Equation. This was the result of - I cj I-. correlating fatigue damage mechanisms and deformation mechanisms operating at

W elevated temperatures on the basis of observed mechanical and microstructural rn 0 I

behavior. Correlation between high-temperature fatigue and stress-strain

properties for nickel-base superalloys and stainless steel substantiated the

method. Parameters which must be evaluated for PP- and CC- life are the

maximum stress achievable under entirely plastic and creep conditions

respectively and corresponding inelastic strains, and the elastic modulus.

For plasticity/creep interaction conditions (PC and CP) two more pairs of

stress-strain parameters must be ascertained

INTRODUCTION

Material response in high-temperature turbomachinery i s controlled by

deformation processes such as time-dependent creep and time-independent

*A. Berkovits National Research Council - NASA Research Associate on

+Deceased.

leave from Technion - Israel Institute of Technology, Department of Aeron utical Engineering, Haifa, Israel.

plasticity during, either tension or compression straining in the fatigue

cycle. Success in estimating low-cycle fatigue life under such strongly

rate-dependent deformation has varied considerably (Refs. 1 to 4 ) .

One of the more successful phenomenological methods for predicting

high-temperature, low-cycle fatigue life is the strainrange partitioning

approach, which was developed at NASA by Manson, Halford, and Hirschberg (Ref.

5 > , and has become a viable engineering design tool (Saltsman and Halford

(Ref. 6)). The procedure involves the experimental determination of the four

basic life relationships, resulting from the four possible combinations of

plastic or creep strainrange in the tensile or compressive halves of the

fatigue cycle* for a given material, and their use in conjunction with an

interaction damage rule to predict cyclic lives.

The present work was undertaken in order to enhance understanding of the

mechanisms which are responsible for observed material response, and thereby

achieve a prediction model. The research required analysis of strain-cyclic

mechanical response in the light of monotonic creep and plasticity response,

and metallographic determination of the plasticity and creep damage mechanisms

operating during SRP fatigue loading. Finally a method for the estimation of

the strain-life relations could be developed, based on the mechanical and

microstructural response observed.

*Throughout the discussion strainrange partitioning notation is used, namely:

Subscripts Tensile loading Compressive loading

PP Plasticity Plasticity cc Creep Creep PC Plasticity Creep CP Creep Plasticity

2

In the present program, Refs. 7 to 9, both cyclic and supporting monotonic

tests were conducted on MAR-M200+2%Hf under strain control at constant

strainrates. Cyclic data on the other materials studied were found in the

literature, so that only monotonic tests and supporting microscopy were

required. The emphasis in this paper is on observed mechanical behavior. A

more extensive discussion of microstructural response can be found in Nadiv

et al. (Ref. 10).

CYCLIC AND MONOTONIC DAMAGE MECHANISMS IN MAR-M2OO+Hf

The initial phase of this research included an extensive mechanical and

metallographic study of directionally solidified MAR-M200+2%Hf, tested at

975 "C.

of strainrates, yielded strainrange-partitioned life data and stress-strain

data in both the plasticity and creep regimes.

examination of test specimens revealed the damage mechanisms operating under

the various strain conditions imposed.

Strain-controlled cyclic and monotonic tests, conducted over a range

Subsequent microscopic

Monotonic Behavior

The tensile stress-strain response of MAR-M200+Hf material at 975 " C is

represented by the curves in Fig. 1 . Examination o f the failed test specimens

revealed three distinct damage mechanisms, depending on the region o f strain

and strainrate.

Beyond the points of maximum true stress the stress-strain curves formed a

set of more or less parallel, gradually decreasing lines. This region was

governed by dynamic recovery of the material.

the maximum stress developed under a given strainrate.

region was bounded by the fracture strain, which was constant a t the lower

strainrates which characterize creep, and decreased somewhat in the

high-strainrate, plasticity-affected regime.

It was bounded on the left by

On the right the

3 -

Plasticity was the dominant process at strainrates above 0.001 sec-1 in

the region before the maximum stress was attained for the given strainrate.

Here the stress-strain relationship was independent of the strainrate at low

inelastic strain. At these relatively high strainrates, dislocation locking

akin to strain-aging did not begin to be felt until considerable strain had

taken place under strain-hardening conditions. In fact, by the time

strain-aging-like effects could take place in this range they were

over-shadowed by the recovery process which eventually led to failure.

lower boundary of the plasticity region was defined by a time-constant,

controlled by the material diffusion rate.

The

Dislocation generation, pile-ups and interactions, causing a hardening

effect, dominated the low strainrate, low strain region of the stress-strain

curves. At low strainrates, low to moderate stresses developed due to

dislocation climb and diffusion processes in the material.

the low energy levels involved here were over-shadowed until substantial

strain had been achieved.

by continuing dislocation locking, with the result that as strainrate decreased

the strain to achieve a given stress increased, until sufficient strain had

accumulated so that recovery became the dominant cause of straining.

Recovery rates at

The diffusion processes were increasingly inhibited

Thus an area where time-independent strain occurs at high-strainrates

prior to initiation of the recovery process, and a region of time-dependent

deformation at more moderate strainrates, were defined. Time-dependent

deformation is caused by dynamic recovery in the material, evidenced as

straining initiated at moderate strainrate and again when failure is

approached. Between these extremes dislocation locking mechanisms retard

deformation and lead to transitory hardening of the material. These

characteristics of monotonic straining have been emphasized because of their

significance relative to cyclic response.

4

Cyclic Behavior

The cyclic response curves obtained under symmetrical strain-control

( R = -1) were unique functions of loading frequency, which was constant

throughout each cyclic test.

hysteresis loops resulted, bounded by common tensile and compressive inelastic

stress-strain curves, regardless of strain amplitude, for both PP and CC

tests (Fig. 2 ) . Furthermore, the plasticity half of CP and PC hysteresis

loops coincided with PP results, while the creep half matched results

obtained from CC tests. The fact that the cyclic stress-strain

relationships did not change as a result of altering the strain amplitude (and

accordingly the strainrate) at a given test frequency, indicates that little

or no hardening occurred under cyclic loading, the material remaining almost

fully recovered and soft. This suggests that radically different

characteristics were active in the microstructure under cyclic and monotonic

conditions, before the onset of the cracking and failure phase.

response under cyclic conditions was noticed previously by Berkovits (Ref. 1 1 )

for Udimet-700, as well as for 316 stainless steel.

E

for a given frequency a single family of

Similar

The phenomenon of preserved "softness" of the material under cyclic

conditions has also been noted by others (Kear and Oblack (Ref. 12)). It i s

apparently the result of relaxation of dislocation back-stress and pile-up

during reverse loading, so that, following each load reversal, the material

responds as if it were virgin material. In this respect, the cyclic response

i s quite different from monotonic behavior, a fact which will be important

when fatigue life prediction is discussed in the next section.

STRAINRANGE PARTITIONED LIFE PREDICTION

The crux of the study of constitutive behavior of materials is its use in

predicting failure. This objective has seldom been achieved (cf. Ross (Ref.

1 3 ) ) . Fatigue-life data obtained from MAR-M200+Hf material at 975 " C will be

5 -

reviewed as the basis for the proposed method for predicting fatigue life from

monotonic stress-strain behavior.

The key to cyclic-creep life prediction is to be

of the cyclic and monotonic stress-strain response i

found by interpretation

terms of the controlli

deformation process.

deformation is defined by the negative slope of the stress-strain curves in

the high-strain region, and by a line which separates the plasticity and

It was stated previously that recovery-controlled

dynamic recovery regimes at low strains.

diffusion energy relation.) The hardening process which occurs at low strains

can be described as a family of curves emanating from the recovery initiation

line and intersecting the recovery curves at strains corresponding to the

maximum stress developed at each strainrate.

two families of curves which represent the controlling processes, i s

significant when attention is turned to the cyclic loading case.

(These curves can be related to the

The line of intersection of the

Under cyclic CC loading (Ref. 91, stress ranges above 600 MPa indicated

some cycl c hardening, while lower stress amplitudes exhibited slight cyclic

softening

tests at commensurate strainrates. It is significant that the observed

hardening and softening rates were such that the cyclic response converged in

all cases on a state-point defined by a stress-range equal to 600 MPa and a

corresponding creep strain amplitude of 0.007.

on the line of maximum stresses for given strainrates (Fig. 31, and on the

recovery curve for a stress of 600 MPa, may not be coincidental. Failure of

MAR-M200+Hf at 975 " C i s due to the dynamic recovery processes and occurs in

the creep region at an upper stress limit of approximately 600 MPa.

recovery mechanism appears to be coarsening of y ' precipitate particles

(Ref. 10). Although at higher stresses failure occurs after plastic

deformation, it is still due to recovery. However, here the recovery

6

Very low frequency tests underwent hardening similar to monotonic

The fact that this point falls

The

mechanism is dissolution of y ' , and ductility is reduced as a result of the

earlier plastic deformation.

developments is given in Ref. (101.)

(An extensive discussion of the microstructural

Thus the monotonic stress and inelastic strain coordinates, 600 MPa and

0.007 respectively, indicated in Fig. 1 , are sufficient to define conditions

leading to failure under the completely relaxed creep cycling, which controls

CC, CP and PC failure in the directionally solidified MAR-M200+Hf material

under discussion. (In most Raterials CP and PC damage mechanisms are not

identical to CC mechanisms, as will be discussed presently.) On this basis,

a relation of the Universal-Slopes type can be suggested as:

U -- A& - UCC N-0.12 0.6 -0.6 2 - E + &UCC (1)

where uuCc

is the ultimate time-dependent stress (equal to 600 MPa in

MAR-M200+Hf at 975 "C), and is the corresponding time-dependent strain

(equal to 0.007 in MAR-M200+Hf at 975 "0, see Fig. 3 .

The present contention, that fatigue life is determined by the material

response at its maximum monotonic strength, before the onset of dynamic

recovery leading to failure, can be applied to PP data as well. Maximum

stress under plasticity conditions occurred in the material investigation at

~ u p p equal to 900 MPa. The corresponding plastic strain EUPP was

approximately 0.007, coincidentally the same value as for

of the constitutive parameters were substituted in a PP relation of the form:

EUCC. These values

A& U UPP N-0.12 0.6 -0.6 + %PP -- -

2 - E ( 2 )

Lifetime of strainrange parti tioned tests conducted under constant

strainrates are shown as symbols in Fig. 4. The PP-life results show

significantly longer lives than CC, PC or CP results. The latter all have

similar fatigue lives, because there are almost no transverse grain boundary

7 -

segments t o a c t as c r a c k i n i t i a t i o n s i t e s under t e n s i l e creep c o n d i t i o n s i n

t h e d i r e c t i o n a l l y s o l i d i f i e d m a t e r i a l (Manson and H a l f o r d ( R e f . 1 4 ) ) . I n a l l

cases d e t e c t a b l e f a t i g u e c racks developed o n l y d u r i n g t h e l a s t 20 p e r c e n t o f

the c y c l i c t e s t s .

d u r i n g which t h e t e n s i l e phase was a t low s t r a i n r a t e (CC and CP) t han i n

those w i t h h i g h t e n s i l e s t r a i n r a t e (PP and PC).

Somewhat more o x i d a t i o n occu r red on t h e c rack f a c e i n t e s t s

C o r r e l a t i o n o f c a l c u l a t e d r e s u l t s o b t a i n e d from E q s . (1 ) and (21, and

shown as curves i n F i g . 4, w i t h CC, CP, PC as w e l l as PP d a t a i s e x c e l l e n t .

CP AND PC LIFE PREDICTION FOR OTHER MATERIALS

I n d i r e c t i o n a l l y s o l i d i f i e d MAR-M200+Hf m a t e r i a l , CP and PC f a t i g u e

l i f e was dominated by t h e c reep l o a d i n g . More i m p o r t a n t l y , t h e p l a s t i c i t y

h a l f o f the s t r a i n c y c l e d i d n o t s i g n i f i c a n t l y a f f e c t t h e response o f t h e

c reep h a l f o f t h e h y s t e r e s i s l oop . A s a r e s u l t , t h e c reep parameters remained

t h e c o n t r o l l i n g f a c t o r s i n CP and PC l i f e . However, i n o t h e r m a t e r i a l s

t h e e f f e c t o f compressive p l a s t i c i t y on t h e subsequent t e n s i l e c reep

h a l f - c y c l e , i n a CP t e s t , and o f compressive c reep on subsequent t e n s i l e

p l a s t i c i t y i n a PC t e s t , must be q u a n t i f i e d i n o r d e r t o p r e d i c t CP and

PC l i f e t i m e s from t h e Un iversa l -S lopes Equat ion . The a p p r o p r i a t e va lues o f

uu and EU xmay be determined from t h e t e n s i l e ;ide o f o n e - c y c l e - t o - f a i l u r e

t e s t s , i n which compressive s t r e s s and s t r a i n e q u i v a l e n t t o t h e n e g a t i v e o f

the maximum ( P or C) s t r e s s and s t r a i n a r e a p p l i e d p r i o r t o t e n s i l e (C or

P r e s p e c t i v e

and PC cond

y) l o a d i n g t o f a i l u r e . The l i f e - p r e d i c t i o n r e l a t i o n s f o r CP

t i o n s then become:

U A C UCP N-0.12 0.6 -0.6 2= E + %CP (3 )

and

CY A C UPC ,,,-0.12 0.6 -0.6 2= E + €UPC ( 4 )

8

respectively.

other elevated temperature materials.

The validity o f this approach was investigated for a number of

COMPARISON BETWEEN PREDICTED AND EXPERIMENTAL SRP LIFE

The key to cyclic life prediction for MAR-M200+Hf at elevated temperature

was perceived by interpretation of the cyclic and monotonic mechanical

response in terms of the controlling mircostructural processes. When the

interaction o f dynamic recovery and strain-hardening processes was understood

in terms of the microstructure, a method of life prediction evolved.

Clearly, in other materials different chemomechanical processes may be

responsible for strain hardening. However, the significant point is that

cyclic strength persists as long as the hardening process, whatever its cause,

dominates the deformation. When the hardening mechanism is superseded by

recovery processes, both monotonic and cyclic resistance are essentially

exhausted. Thus the stress and strain parameters obtaining at the point of

mechanism-change in the monotonic test are also the controlling parameters for

estimating cyclic life. These parameters are the true stress and the

corresponding true stra n in the plasticity and creep-range strength tests.

Evaluating these pa ameters in the plasticity stress-strain curve is

straight-forward. However, they must also be ascertained for the (unknown)

strainrate at which the full creep capability of the material is realized. In

principle the required stress-strain curve is that for which the ultimate

stress corresponds to the proportional limit stress o f the plasticity curve.

In practice, the correct curve may be better determined by taking some small

plastic strain such as the 0.1 percent offset as the limit, instead of the

proportional limit (Fig. 1).

In order to test the val

critical damage point on the

life, a comparison was perfo

dity of the relationship between the perceived

stress-strain curve and strainrange partitioned

med for three nickel-base superalloys and a

stainless steel. Fatigue life data were taken from the literature for the

following materials: AF2-1DA at 760 "C, Saltsman and Halford (Ref. 6);

Udimet-700 at 760 "C, Berkovits (Ref. 1 5 ) ; S.S. 316 at 705 "C, Hirschberg and

Halford (Ref. 16).

For each material a series of six stress-strain tests were conducted under

strain control at the appropriate temperature. The six tests consisted of:

1 . High-strainrate test to define p1asticit.y parameters;

2. Moderate-strainrate tests to define the creep region upper bound;

3. One-cycle-to-failure tests (one CP, one PC). Strainrates and

compressive strain limit were as defined by the previous tests, so that

appropriate &U-strains were achieved in compression, and creep-phase

strainrates were at the plasticitylcreep boundary.

Fractography performed on two o f the materials tested (Figs. 5 and 6)

indicated that the failure-surface characteristics of the

one-cycle-to-failure CP and PC tests corresponded to those of the tensile

creep-range and plasticity-range failures respectively. Stress-strain curves

for the materials tested are presented in Fig. 7, and parameters used in order

to define the plasticitylcreep boundary are shown in Fig. 8. Values of stress

and strain parameters obtained from these tests f w use in the life-prediction

Eqs. (1) to ( 4 ) are tabulated in Table 1. Values for MAR-M200+Hf are also

included for completeness. Total strainrange against lifetime curves for the

four materials calculated with the aid of Table 1 are compared with

experimental da a in Figs. 4 and 9. In general the correlation supports the

proposed method Results of the comparison indicate that parameters required

for Universal-S opes prediction of SRP lifetimes can be evaluated from

critical damage points of monotonic stress-strain data.

10

CONCLUSION

A metallographic study of the correlation between SRP fatigue damage

mechanisms, and deformation mechanisms operating in the early stages of

monotonic plastic and creep ductility at elevated temperatures, evolved a

method of evaluating material parameters for Universal-Slopes prediction of

fatigue life. Besides the elastic modulus of the material, the required

parameters are the maximum true stress and corresponding inelastic strain

under the four strainrange partitioning conditions. These conditions include

the strainrates which produce entirely plastic and upper creep-boundary

deformations, and two one-cycle-to-failure cases which define plasticitylcreep

interactions.

few as six stress-strain tests.

Comparison of predicted lifetimes with experimentally determined fatigue

The parameters can be evaluated at a given temperature by as

lives for four elevated temperature materials demonstrated that the principles

on which the prediction method is based are sound.

REFERENCES

1 . Fong, J.T. (ed.), "Fatigue Mechanisms." ASTM STP-675, ASTM, 1979.

2. Burke, J.J. and Weiss, V. (eds.), "Fatigue Environment and Temperature

Effects." Plenum, 1983.

3 . Amzallag, C., Leis, B.N. and Rabbe, P. (eds.1, "Low-Cycle Fatigue and Life

Prediction." ASTM STP-770, ASTM, 1982.

4. Brunetaud, R . , et al. (eds.), "High Temperature Alloys for Gas Turbines

1982." D. Reidel, 1982.

5 . Manson, S.S., Halford, G.R. and Hirschberg, M.H., "Creep-fatigue analysis

by strain-range partitioning," Design for Elevated Temperature

Environment, S . Y . Zamrik, ed., ASME, 1971, pp. 12-24.

6. Saltsman, J.F. and Halford, G . R . , "Life Prediction o f Thermomechanical

Fatigue Using Total-Strain Version of Strainrange Partitioning (SRP),"

NASA TP-2779, 1988.

1 1

7. Berkovits, A. and Nadiv, S. , "Constitutive relationships for creep-fatigue

in high-temperature materials," Mechanical Behavior of Materials - IV,

Vol. 1 , J. Carlsson and N.G. Ohlson, eds., Pergamon, 1984, pp. 149-156.

8. Berkovits, A. and Nadiv, S . , "Creep-fatigue in PWA 1422 material,"

Proceedings of the V International Congress on Experimental Mechanics,

Society for Experimental Stress Analysis, 1984, pp. 664-669.

9. Berkovits, A., "Estimation of high temperature low cycle fatigue on the

basis of inelastic strain and strainrate," NASA TM-88841, 1986.

10. Nadiv, S. , Berkovits, A. and Shalev, G., "Estimation of high-temperature

low cycle fatigue on the basis of inelastic strain and strain rate,"

Fatigue Life:

Park, OH, 1986, pp. 399-404.

Analysis and Prediction, American Society of Metals, Metals

1 1 . Berkovits, A., "Hodographic prediction of cyclic creep behavior," J.

Aircraft, 11, 10-14, 1974. 12. Kear, B.H. and Oblak, J.M., "Deformation modes in y' precipitation

hardened nickel base alloys," Journal de Physique Colloque, 3, 7-35 to 7-45, 1974.

13. Ross, D.A., Chang, T.Y. and Thompson, R.L. (eds.), "Nonlinear Constitutive

Relations for High Temperature Applications", NASA CP-2271, 1983.

14. Manson, S.S. and Halford, G.R., "Complexities of high temperature metal

fatigue: Some steps toward understanding," Israel Journal o f Technology,

- 21, 29-53, 1983.

15. Berkovits, A., "Prediction o f inelastic high temperature materials

behavior by strain-rate approach," J. Eng. Mater. Technol., 96, 106-108, 1974.

16. Hirschberg, M . H . and Halford, G.R., "Use o f strainrange partitioning to

predict high-temperature low-cycle fatigue life," NASA TN D-8072, 1976.

12

TABLE 1 . - STRESS-STRAIN PARAMETERS FOR PREDICTION OF

SRP FROM UNIVERAL-SLOPES EQUATION

~~

T e m p e r a t u r e , C E, MPa uUpp. MPa

%PP

EUCC

%CP

V P C

ouCc, MPa

uuCp, MPa

ouPc, MPa

P a r a m e t e r

MAR-M200

975 81 700

900 0.007

600 0.007

600 0.007

600 0.007

looor STRAIN RATE,

M a t e r i a 1

AF2-1 DA

7 60 161 600

1 041 0.006

965 0.005

834 0.01 1 5

1 062 0.02

8oo

G Oucc- . 61

U-700

7 60 162 700

1 320 0.122 1 060 0.135

950 0.036 1 320 0.122

S . S . 316

705 1 1 1 000

300 0.080

300 0.080

300 0.020

300 0.020

I I I I I 0 .04 .08 ,012 .016 .020 .024

TRUE STRAIN FIGURE 1. - TENSILE PROPERTIES OF MAR-M200+Hf AT 975 OC.

13

400 I- f I FREQUENCY = 0.02 Hz

c

-600 -.008 -.004 0 .004 .008 STRA I N

FIGURE 2. - HYSTERESIS LOOPS FOR MAR- M 2 W H f AT 760 OC.

1000

3 600 x 400

200

9 D

I I I I I I I I I 1 I I I I

I I I I IIIII I I 1 1 1 1 1 1 1 I I I I11111 I I I I I I111 I I I I 1 1 1 1 1 10-6 10-5 IO-^ 10-3 10-2 10-1

-02

STRAIN RATE, SEC-'

FIGURE 3. - MATERIAL PARAMETERS FOR NAR-M200+Hf AT 975 'C.

14

DRIGINAI; PAGE IS OF POOR QUALITY

s .02

d

.006

w

CC

c v,

CC I- O

0 cc 0 PP A PC 0 cp

UNIVERSAL SLOPES

,004 I I I I I 101 102 103 I 04 1 o5 106

CYCLES TO FAILURE

FIGURE 4. - SRP LIFETIME OF MAR-1200-Hf AT 975 'C.

( B ) CREEP-RANGE. (A) PLASTIC.

(C) ONE-CYCLE-TO-FAILURE PC. (D) ONE-CYCLE-TO-FAILURE CP.

FIGURE 5 . - TENSILE-TEST FAILURES OF AF2-IDA NICKEL ALLOY AT 760 OC.

15

( A ) PLASTIC.

( C ) ONE-CYCLE-TO-FAILURE PC.

(B) CREEP-RANGE.

(D) ONE-CYCLE-TO-FAILURE CP.

FIGURE 6 . - TENSILE-TEST FAILURES OF STAINLESS STEEL 316 AT 705 OC.

ORIGINAL PAGE IS OF POOR Q U A L m

16

i . SEC-1 -

1 . 8 ~ 1 0 - ~ 1-CYCLE PC

2 . 6 ~ 1 0 - ~ \

L2.4x10-5 1-CYCLE CP

1200

a

800 * * W oc I-

W 2 z I-

* 4oa

O ,004 ,008 .012 ,016 ( A ) AF2-1DA NICKEL ALLOY AT 760 'C.

l6O0 r 1. S K - 1

W

1 I I J 0 .08 .16 .24 .32

(B) UDIMET-700 NICKEL ALLOY AT 760 'C.

1, s c - 1 ,Oar 2.1x10-4- a B * * W z c v)

W T3 E

0 -08 .16 .24 - 3 2 TRUE STRAIN

( C ) STAINLESS STEEL 316 AT 705 'C.

FIGURE 7 . - TENSILE STRESS-STRAIN CURVES.

17

.02

,006 :: - 0 1

(A) AF2-IDA NICKEL ALLOY AT 760 O C .

10 6 4

2

1

000 0 MAXIMUM STRESS 0 FLOW STRESS

000

@

. I

(B) UDIMET-700 NICKEL ALLOY AT 760 OC.

600 r

a 2 r ;L a02 .01

10-6 10-5 10-4 10-3

200

100 !E 60 40 E

W

20 5 10

10-2

STRAIN RATE. SEC-'

( C ) STAINLESS STEEL 316 AT 705 OC.

FIGURE 8. - DEFINITION OF PLASTICITYICREEP BOUNDARY.

18

CALCULATED TEST

,001 102 103 104 io5

CYCLES TO FAILURE

AT 760 OC. ( A ) AF2-1DA NICKEL ALLOY

€

102 103 104 105 CYCLES TO FAILURE

(B) UDIMET-700 NICKEL ALLOY AT 760 OC.

I- i .--PC.CP

102 103 104 105 CYCLES TO FAILURE

(C) STAINLESS STEEL 316 AT 705 OC.

FIGURE 9. - PREDICTED SRP LIFETIMES.

19

NASA 1. Report No.

NASA TM-100796

Report Documentation Page 2. Government Accession No. 3. Recipient's Catalog No.

4. Title and Subtitle

R e l a t i o n s h i p Between F a t i g u e L i f e i n t h e Creep- F a t i g u e Region and S t r e s s - S t r a i n Response

5. Report Date

6. Performing Organization Code l------ 7. Author@)

A . B e r k o v i t s and S. Nadiv 8. Performing Organization Report No.

E-3972

10. Work Unit No

506-63-1 B 9. Performing Organization Name and Address

11. Contract or Grant No. N a t i o n a l Ae ronau t i cs and Space A d m i n i s t r a t i o n Lewis Research Center

7. Key Words (Suggested by Author@))

F a t i g u e ; L i f e - p r e d i c t i o n ; S t r e s s - s t r a i n ; F a t i g u e damage; High- temperature; M i c r o s t r u c t u r e ; Creep; Supera l l oys

13. Type of Report and Period Covered Cleve land , Oh io 44135-3191

18. Distribution Statement

U n c l a s s i f i e d - U n l i m i t e d Sub jec t Category 39

Technica l Memorandum 2. Sponsoring Agency Name and Address

9 Security Classif (of this report)

U n c l a s s i f i e d

N a t i o n a l Ae ronau t i cs and Space A d m i n i s t r a t i o n Washington, D.C. 20546-0001

20 Security Classif (of this page) 21 No of pages 22 Price'

U n c l a s s i f i e d 20 A02

I

14. Sponsoring Agency Code : I

5. Supplementary Notes

Prepared f o r t h e V I I n t e r n a t i o n a l Congress on Exper imenta l Mechanics sponsored by t h e S o c i e t y o f Exper imenta l Mechanics, P o r t l a n d , Oregon, June 5-10, 1988. A Berkovi t s , N a t i o n a l Research Counci l - NASA Research A s s o c i a t e on l e a v e from Technion - I s r a e l I n s t i t u t e o f Technology, Dept . o f Aero- n a u t i c a l E n g i n e e r i n g , H a i f a , I s r a e l : S . N a d i v , Technion - I s r a e l I n s t i t u t e o f Technology, D e p t . o f M a t e r i a l s E n g i n e e r i n g .

6. Abstract

On t h e b a s i s o f mechanical t e s t s and m e t a l l o g r a p h i c s t u d i e s , s t r a i n r a n g e p a r t i - t i o n e d l i v e s were p r e d i c t e d by i n t r o d u c i n g s t r e s s - s t r a i n m a t e r i a l s parameters i n t o the U n i v e r s a l Slopes Equat ion . Th is was the r e s u l t o f c o r r e l a t i n g f a t i g u e damage mechanisms and d e f o r m a t i o n mechanisms o p e r a t i n g a t e l e v a t e d temperatures on t h e b a s i s o f observed mechanica l and m i c r o s t r u c t u r a l behav io r . C o r r e l a t i o n between h igh- tempera ture f a t i g u e and s t r e s s - s t r a i n p r o p e r t i e s f o r n i cke l -base s u p e r a l l o y s and s t a i n l e s s s t e e l s u b s t a n t i a t e d t h e method. Parameters which must be eva lua ted f o r PP- and CC- l i f e a r e t h e maximum s t r e s s ach ievab le under e n t i r e l y p l a s t i c and c reep c o n d i t i o n s r e s p e c t i v e l y and cor respond ing i n e l a s t i c s t r a i n s , and t h e e l a s t i c modulus. For p l a s t i c i t y l c r e e p i n t e r a c t i o n c o n d i t i o n s (PC and CP) two more p a i r s o f s t r e s s - s t r a i n parameters must be a s c e r t a i n e d .