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The effect of creep cavitation on the fatigue life under creep-fatigue interaction
Nuclear Engineering and Design 153 (1995) 213 221
ELSEVIER
Nuclear Engineering and Design
The effect of creep cavitation on the fatigue life under creep-fatigue interaction Soo Woo
N a m a, S o o C h a n
L e e a, Je M i n L e e b
a Department o f Materials Science and Engineering, Korea Advanced Institute of Science and Technology, 373-1 Kusong-Dong, Yusung-Gu, Taejon, South Korea b Defense Quality Assurance Agency, PO Box 24, Changwon, Kyungnam, South Korea
Abstract
Low cycle fatigue tests have been carried out with three different materials (1Cr-Mo-V steel, 12Cr-Mo-V steel and 304 stainless steel) for the investigation of the effect of surface roughness on the fatigue life. To see the effect systematically, we have chosen those materials which may or may not form grain boundary cavities. Test results show that the continuous fatigue life of 1Cr-Mo-V steel and aged 304 stainless steel with a rough surface is decreased compared with that of the specimens with a smooth surface. These two alloys are found to have no grain boundary cavities formed under creep-fatigue test conditions. On the contrary, the fatigue life of 12Cr-Mo-V steel and solutionized 304 stainless steel in which grain boundary cavities are formed under creep-fatigue test conditions is not influenced by the states of surface roughness. The characteristic test results strongly confirm that the fatigue life of the specimen under creep-fatigue interaction, during which creep cavities are forming, may be controlled by the cavity nucleation and growth processes rather than the process of surface crack initiation.
I. Introduction
In general, fatigue failure occurs by the initiation and propagation of surface cracks, and fatigue life is, therefore, expressed as a summation of the number of cycles to crack initiation and for propagation. In low temperature high cycle fatigue, nearly the entire fatigue life is spent in crack initiation (Laird, 1963), and the fatigue life depends greatly on surface roughness because surface grooves act as fatigue crack initiation sites owing to high stress concentration and fatigue cracks initiate
easily on them (Fluck, 1951; Juvinal, 1967). In high temperature fatigue, however, it has the character of low cycle fatigue (LCF) because it is concomitant with the plastic deformation during fatigue cycling. Also, the conditions of use of high temperature materials are such that creep deformation in addition to fatigue deformation will be undergone. In high temperature fatigue cycling with creep effect, grain boundary cavities are nucleated and the nucleated cavities grow to show high temperature brittle fracture behaviour. The high temperature fatigue behaviour of these materials may be
0029-5493/95/$9.50 © 1995 Elsevier Science S.A. All rights reserved SSDI 0029-5493(94)00829-N
214
S.W. Nam el al./ Nuclear Engineering and Design 153 (1995) 213 221
different from that of LCF behaviour. Even though a fatigue crack is formed at the surface and may propagate into the material, if creep cavitation prevails with a high value of the cavity nucleation parameter, the fatigue life may not be influenced by the fatigue crack but affected by the cavities. Recently, we have proposed a model which predicts the fatigue life of a material under creepfatigue interaction condition. This model explains that vacancies are formed by plastic deformation during the fatigue cycle and these vacancies are clustered to form cavities on the grain boundaries. These nucleated cavities grow during the hold time period. The nucleation of cavities on the grain boundaries is assumed to be strongly dependent on the characteristics of grain boundary precipitates, i.e. each material has its own nucleation parameter. In this study, to see the effect systematically we have chosen three materials (ICr M o - V steel, 12Cr M o - V steel and 304 stainless steel) which may or may not form grain boundary cavities, and LCF tests were conducted on these specimens with different surface roughness, i.e. very smooth and rough surfaces. The results obtained are discussed in terms of the importance of the crack initiation stage in continuous cycling and the effect of the cavitation process in creep-fatigue interaction.
2. Experimental method 1 C r - M o - V steel, 1 2 C r - M o - V steel, solutionized 304 stainless steel and aged 304 stainless steel were used for the fatigue tests. Solution treatment was undertaken at 1373 K for l h followed by water cooling. Some of the solution treated specimens were aged at 1033 K for 50 h. The grain size and microstructure of the materials used in this work are given in Table 1. The specimen was of a cylindrical type with a gauge length of 8 ram. Two different models of surface roughness of the gauge section (smooth and rough surface) were obtained by mechanically polishing the specimens perpendicular to the stress axis with emery papers of grades 80 and 1200 (1500) respectively.
Table 1 The grain size and microstructure of materials used in this work Material
Grain size
Microstructure
1Cr Mo V steel 12Cr Mo V steel 304 stainless steel
501am 100 ~tm 50 ~tm
Tempered bainite Tempered martensite Austenite
(a)
(b)
J
2.5o.m I
, L 2001sin
t
2.5t*m /
" ' 200~m
Fig. 1. Surface roughness profiles of 304 stainless steel specimens with two different surface roughnesses, polished with emery paper of grades (a) 80 and (b) 1200.
The surface roughness profiles in 304 stainless steel which were measured with a Talysurf instrument are shown in Fig. 1 as an example and the maximum values of the peak-to-valley roughness in all the specimens are shown in Table 2. Total strain-controlled uniaxial push-pull fatigue tests were carried out in an electromechanically driven closed-loop controlled Instron model 1362 device with a strain rate of 4 x 10-3 sFor high temperature fatigue tests, a radiant furnace was used to heat the specimen. For the investigation of the phenomenon of creep-fatigue interaction, a 10 30min hold at tensile peak strain was applied. The characteristic morphology of the fatigue surface and the cavitation behaviour during fagitue cycling were examined by scanning electron microscopy (SEM).
215
S.W. Nam et al. / Nuclear Engineering and Design 153 (1995) 213-221 Table 2 The maximum value of the peak-to-valley roughness height in the specimen used in this work Material
Roughness for emery paper grade 80
Roughness for emery paper grade 1200
ICr M o - V steel 1 2 C r - M o - V steel 304 stainless
8.0 ~tm 8.8 lam 13 I~m
0.4 lam 0.9 tam 0.3 lam "
Polished with emery paper grade 1500.
3. Experimental results
10 304 S-S 873K
o. CO
3.1. Fatigue life with surface roughness
<3
3.1.1. I C r - M o - V steel The continuous LCF and creep-fatigue lives of the 1 C r - M o - V steel with two different surface roughness modes at 823 K are shown in Fig. 2. In the continuous LCF and creep-fatigue tests (10 min hold at tensile peak strain), the fatigue life of the specimen with the rough surface decrased compared with that of the smooth surface as shown in Fig. 2. 3.1.2. 304 stainless steel The continuous LCF lives of the solutionized and the aged 304 stainless steel with different roughnesses at 873 K are shown in Fig. 3. The fatigue life of the specimen with a rough surface decreased with increasing surface roughness as shown in Fig. 3. 10 -'
c
1 C r - M o - V steel Ar a t m o s p h e r e Temperat, ure 823K
~
[12
t. = 0 •
o
l Omin. o #1200
•
10-3
I
I
o
I
10
I
"~ I.-
#80 I I III 10
Number
~
I
I
I
I
cont., cont., cont., cont.,
sol., aged., $o1., aged.,
& C O rY
10
.c_ 0 l,q
0
ft. 1°,0
. . . . . . .
;4'
'
. . . . . .
io'
Critical Number of Cycles ( N c r ) Fig. 3. Relationship between plastic strain range and fatigue life in 304 stainless steel specimens with two different surface roughnesses under continuous cycling.
In the creep-fatigue test (10 min hold at tensile peak strain) the fatigue life of the solutionized specimen is found to be independent of the surface roughness conditions; however, that of the aged specimen is observed to decrease with increasing roughness as shown in Fig. 4.
3.1.3. 1 2 C r - M o - V steel The test results of continuous LCF and creepfatigue (30 min hold at tensile peak strain) of the 12Cr-Mo-V steel with two different surface roughness modes at 873 K are shown in Fig. 5. In the continuous LCF and creep-fatigue tests, the fatigue lives of the specimens are also found to be independent of the surface conditions as shown in Fig. 5.
.c__ ~ 10 -,
_o
o •
I I ''01
~
'
of Cycles to Failure
Fig. 2. Relationship between plastic strain range and fatigue life in I C r - M o - V steel specimens with two different surface roughnesses.
3.2. Observation of the fracture surface To investigate the nature of the crack propagation behaviour of the specimens with the different
216
S.W. Nam et al. / Nuclear Engineering and Design 153 (1995) 213 -221
1 0 -I
,504 S - S , Hold Time
a. co
875K 10 min. ,~ . : * t
emery poper ~1500 emery poper #80
E ~ 1 0 -~ .C 0 L ~q
0
~
, , , t
i
i
i
i
i
l l l l
10 ~
,
,
I0 ~
Criticol
Number
of
Cycles
, , ,,
10 ' (
Nor )
Fig. 6. OM images of the crack propagation modes of l Cr Mo V steel at 823 K: (a) continuous cycling; (b) creep fatigue cycling.
Fig. 4. Relationship between plastic strain range and fatigue life in 304 stainless steel specimens with two different surface roughnesses under creep fatigue cycling.
10 -'
12~Cr Mo V Steel Cont. Cycling 875K, Air C
C (/3
52 o
0:
O 10 "~0 2
th=50mi n, ~(1200 th=50min, #80 I 10 ~
10 '
Number of Cycles to Foilure Fig. 5. Relationship between plastic strain range and fatigue life in 1 2 C r - M o - V steel specimens with two different surface roughnesses.
surface roughness modes, the specimens were fatigued up to the number of cycles required to attain an about 5% drop in tensile peak load and then polished. For the observation of the crack propagation mode, the specimens were observed by optical micrography (OM) and optical micrographs of solutionized 304 stainless steel and I C r - M o - V steel, as the typical examples, are shown in Fig. 6 and Fig. 7 respectively. All the other results are listed in Table 3, It is assumed that the crack initiation phase is up to the crack size of 100 lam (one or two grain sizes).
Fig. 7. OM images of the crack propagation modes of solutionized 304 stainless steel at 823 K: (a) continuous cycling: (b/ creep fatigue cycling.
In the continuous cycling, fatigue crack propagation modes of all the materials tested in this study were transgranular without any grain boundary cavitation as shown in Fig. 6(a) and Fig. 7(a). In the creep-fatigue cycling, however, intergranular crack growth was observed in all the materials except 1Cr-Mo-V steel, in which cracks propagate transgranularly without cavitation damage as shown in Fig. 7(b) and Fig. 6(b). The nucleation of cavities on the grain boundaries is considered to be strongly related to the inter-
S.W. Nam et al. / Nuclear Engineering and Design 153 (1995) 213-221
217
Table 3 Variation in the fatigue life and fracture mode of materials with testing conditions Material
Temperature
Continuous cycling
Creep-fatigue cycling
1Cr-Mo V steel
823 K
12Cr-Mo-V steel
873 K
Solutionized 304 stainless steel Aged 304 stainless steel
873 K
Transgranular, Nr(R) < Nr(S) Transgranular, Nf(R) = Nf(S) Transgranular, Nr(R) < Nr(S) Transgranular, Nr(R) < Nr(S)
Transgranular, Nr(R) < Nr(S) Intergranular, Nr(R) = Nr(S) Intergranular, Nr(R) = Nf(S) Transgranular, Nr(R) < Nr(S)*
873 K
Nr(R), fatigue life of the specimen with a rough surface; Nr(S), fatigue life of the specimen with a smooth surface; Nr(S)*, intergranular.
granular crack growth modes in the creep-fatigue cycling.
4. Discussion
4.1. Effect o f surface roughness on the continuous low cycle fatigue
In the continuous LCF tests, the fatigue life of all the materials tested in this study is decreased with increasing surface roughness except for 1 2 C r - M o - V steel in which it is independent of surface modes as shown in Fig. 5. In 1 2 C r - M o V steel as a special case, fatigue cracks are nucleated at the grain boundaries of the specimen regardless of the surface roughness; there is no difference in the fatigue life with surface roughness. Also, the slopes of both plots with surface roughness in each material are approximately the same over the plastic strain range tested. If the fatigue life is dominated by the crack propagation the slope for the test results with a rough specimen should be steeper than that of a smooth specimen because of the general concept (Laird, 1963; Maiya, 1975a) that the life fraction for crack initiation decreases with increasing strain range. This is inconsistent with the present results. The experimental results suggest that LCF life is strongly influenced by the crack initiation process even for a high strain range and the fraction of the fatigue life for crack initiation is nearly the
same over the plastic strain range tested. This agrees qualitatively with Dowling's observations (Dowling, 1977) that the life fraction for any given surface crack size (0.01 inch) is independent of the life in A533B steel. To date many authors (Magnin, 1985, 1988; Maiya, 1957b) have investigated surface roughness effects to evaluate the importance of the crack initiation stage in LCF. Maiya and Busch tested type 304 stainless steel at 866 K and suggested that the fatigue life decreased with increasing surface roughness owing to the reduction in the number of cycles to the crack initiation (Maiya, 1975b). Magnin et al. reported that the fatigue cycles up to nucleation and evolution of micro-cracks at the specimen surface of austenitic and ferritic stainless steel represented most of the LCF life regardless of the strain amplitude applied (Magnin, 1985, 1988). However, Wareing and Vaughan argued that the LCF behaviour was dominated by the crack propagation process (Wareing, 1977) and they insisted that the difference in fatigue life between machined and electropolished specimens of type 316 stainless steel resulted from the different initiated crack shapes for the two different surface finish modes at 673 K (Wareing, 1979). To investigate the importance of crack initiation in high temperature LCF of 1 C r - M o - V steel, the smooth surface specimen was fatigued at a plastic strain range of 0.7% for a number of cycles corresponding to 60% of the fatigue life at
218
S.W. Nam et al. Nuclear Engineering and Design 153 (1995) 213- 221
Fig. 8. SEM image of the fracture surface of 1Cr Mo V steel specimen with smooth surface, fatigued for 60% of the fatigue life at 823 K with a plastic strain range of 0.7% (the arrows indicate a crack propagated at 823 K).
this strain range. The specimen was fractured by impact in liquid nitrogen and the fracture surface was observed by SEM, as shown in Fig. 8. A few microcracks (a crack size of 80 ~tm) nucleated at the specimen surface are shown on the fracture surface. From these results, the fatigue crack initiation is very important in the continuous LCF in which grain boundary cavities are not nucleated.
4.2. Effect of surface roughness on the creep-fatigue test In the creep-fatigue test holding at tensile peak strain, fracture modes are divided into two groups. One is the case in which a fatigue crack propagates transgranularly similarly to the case of continuous cycling because of no cavitation damage during creep-fatigue cycling; the other is the case in which a fatigue crack propagates in an intragranular mode due to the cavitation damage during creep-fatigue cycling. The reduction in the fatigue life in the case of no cavitation damage is due to the earlier crack nucleation and the accelerated crack propagation rate because of the recovery effect ahead of the crack tip during a hold time period. Crack nucleation is also important in this case. The life of the specimen with a rough surface is nearly half that of the specimen with a smooth surface; this is considered to be due to the earlier crack initiation at the surface roughness grooves
in the case of 1 C r - M o - V steel. In this steel, creep cavitation damage is not observed during the creep-fatigue test because of the non-existence of the cavity nucleation sites, MnS particles, in the grain boundaries (Wang, 1985). Similarly, the fatigue life of the aged 304 stainless steel is decreased with increasing surface roughness because of the delayed cavitation process owing to the large spacing between the grain boundary precipitates ( M 2 3 C 6 ) which act as cavity nucleation sites. Also, the crack propagation mode of the specimen with the rough surface is transgranular because of the earlier crack initiation at the surface roughness grooves and the delayed cavitation process. However, the other materials such as 12Cr M o - V steel and solutionized 304 stainless steel show a different aspect in the surface roughness effect because of the creep cavitation damage. In case of 1 2 C r - M o V steel, the cavities are nucleated not only on prior austenite grain boundaries but also on lath martensite boundaries, and the nucleated cavities grow during every tensile hold time period. The carbides in the prior austenite grain boundaries and the lath martensite boundaries in 12Cr Mo V steel act as the cavity nucleation site. Similarly, in solutionized 304 stainless steel, many carbides (M23C6) are precipitated at the grain boundaries during the creep-fatigue cycling, and these carbides act as cavity nucleation sites. This high density of grain boundary carbide accelerates a cavitation process (Lee, in press). In the above two materials, the reduction in the fatigue life with surface roughness is considered to be caused by the increased crack propagation rate due to the interaction between the fatigue damage as surface crack initiation and growth and the creep damage as cavity nucleation and growth at the grain boundaries during hold time period. However, the surface condition which strongly influences the fatigue crack initiation is known to have little effect on the fatigue life because the cavitation damage is more effective than the fatigue damage in these two materials. It is thought that the fatigue life is nearly the same as shown in Figs. 4 and 5, because the creep cavitation process takes a decisive role in the fatigue life.
219
s.w. Nam et al. / Nuclear Engineering and Design 153 (1995) 213-221 4.3. Life prediction in c r e e p - f a t i g u e test
Recently, we have proposed a model (Hong, 1985) which predicts the fatigue life of a material in which the creep cavitation damage is dominant under creep-fatigue cycling. This model explains that vacancies are formed by plastic deformation during the fatigue cycle and these vacancies are clustered to form cavities on the grain boundaries. Also, it is assumed that the number of cavities generated in one cycle is proportional to the plastic strain range. The number n of nucleated cavities during cyclic loading per unit area of the grain boundary is represented by (Lee, in press) n = P A% N
(1)
where P, AEp and N are the nucleation factor per unit area of grain boundary, the plastic strain range and the number of cycles respectively. These generated cavities are assumed to grow during the hold time period in tension straining by grain boundary diffusion of vacancies. The Hull and Rimmer model (Hull, 1959) for diffusional growth of cavities at the grain boundary provides a good approximation of cavity growth and the stress term of the model has to be modified as a function of hold time because of load relaxation during tensile strain hold (Lee, in press): d A /dt = ~Dg&f2a( t) /lk T
(2)
where A is the area of a given cavity, l is the cavity spacing, Ogis the grain boundary diffusion coefficient, f2 is the atomic volume, a(t) is the relaxed stress variation during the hold time, and k and T have their usual meanings. From the total cavitated area per unit area of grain boundary up to N cycles and a failure criterion, which is that the load-carrying capacity is drastically reduced by coalescence of grain boundary cavities, the final form of the fatigue life prediction for a material of which creep-fatigue failure is controlled by creep damage rather than by a process of fatigue crack initiation and propagation is given by (Lee, in press)
(3)
lSkv
15gk~
67P
gl¢
Fig. 9. SEM image showingthe fatigue crack initiation mode in solutionized304 stainlesssteel at 873 K after ( 1/3)N~,.under creep fatigue cycling. where Qg is the activation energy of the grain boundary diffusion, t is the hold time and C is a constant including the critical cavitated area. The proposed equation has been checked using the experimental results of 304 stainless steel and 12Cr M o - V steel in which failure is controlled by the creep cavitation damage. The fatigue cracks in solutionized 304 stainless steel are nucleated at grain boundaries owing to grain boundary cavitation as shown in Fig. 9. To investigate whether the cavities are nucleated continuously in every cycle, the specimens of solutionized 304 stainless steel were fatigued up to 1/3 and 2/3 of the fatigue life and near fatigue life respectively and then the specimens chilled in liquid nitrogen were broken by impact. The impact fracture surfaces were observed by SEM and are shown in Fig. 10. Fig. 10 shows that the number of cavities increased with increasing creep-fatigue cycling, and this result means that cavities are nucleated continuously with the fatigue cycles. All the above results mean that the proposed equation is applicable to the two materials. Also, the nucleation of cavities on the grain boundaries is assumed to be strongly dependent on the characteristics of grain boundary precipitates, i.e each material has its own nucleation parameter. The nucleation parameters of the 1 2 C r - M o - V steel, the solutionized 304 stainless steel and the aged 304 stainless steel are 4.5 x 10 ~°, 6.43 x 101~ and 1.27 × l0 II respectively. Using the above nucleation parameters, we predict the creep-fatigue life of each material with the experimental conditions and the predicted
220
S. I,IL Narn et al. / Nuclear Engineering and Design 153 (1995) 213-221
/
304 S-S , 873K sol aged t.(rnin.) 10 3
.~
*
30
•
50 /
/ / /
.
•
n
10 2
,
,
,
,
~
, ~,~
10 3
I0
Experimental
Life
Fig. 11. Predicted life by creep cavitation only vs. experimental life on 304 stainless steel under creep fatigue cycling at 873 K.
1000
8OO Q
,._e Fig. 10. SEM images showing grain boundary cavities on intergranular fracture surfaces produced at liquid nitrogen temperature by impact in solutionized 304 stainless steel, fatigued at 873 K: (a) (1/3)Nor; (b) (2/3)N~; (c) near Nor.
~.~ ._u ~XD 400600
5. Conclusions (1) The fatigue life of the specimens with a rough surface decreases compared with that of the
o
o
EL ~/
200
creep-fatigue lives are compared with experimental lives, as shown in Fig. 11 and Fig. 12. These plots show that the predicted lives are in good greement with the experimental lives. From these results, it may be understood that creep-fatigue failure is controlled by creep damage only rather than by a process of fatigue crack initiation and propagation. It is thought that the proposed equation in which only cavity nucleation and growth are considered is good for life prediction in creep-fatigue cycling.
o
0
/7 /
873K, Air
/
I 200
o : th=lOmin.
:
I 400
..
~ 600
Experimentol
I 800
1000
Life
Fig. 12. Predicted life by creep cavitation only vs. experimental life on 1 2 C r - M o - V steel under creep-fatigue cycling at 873 K.
specimens with a smooth surface in continuous LCF. This decrease in the fatigue life with surface roughness is thought to be due to the reduction in the number of cycles to the crack initiation caused by the surface roughness. Therefore, it is thought that the crack nucleation is considered to be an important process in the fatigue life for the materials without grain boundary cavitation.
S. IV. N a m e t al. / Nuclear Engineering and Design 153 (1995) 213-221
(2) The fatigue life of the materials in which the creep cavitation process takes place is independent of the surface roughness conditions because the creep-fatigue failure is controlled mostly by the creep damage rather than by a process of fatigue crack initiation. From this result, it is thought that the proposed life prediction formula based on considering the creep cavitation only is reasonable.
References N.E. Dowling, Crack growth during low-cycle-fatigue of smooth axial specimens, ASTM 637, 1977, p. 97 (ASTM, Philadelphia, PA). P.G. Fluck, The influence of surface roughness on the fatigue life and scatter of test results of two steels, ASTM 51 (1951) 584. J.W. Hong, S.W. Nam and K.T. Rie, A model for life prediction in low-cycle fatigue with hold time, J. Mater. Sci. 20 (1985) 2763. D. Hull and D.E. Rimmer, The growth of grain-boundary voids under stress, Philos. Mag. 4 (1959) 673. R.C. Juvinal, Effect of load type, specimen size and surface finish, Stress, Strain and Strength, McGraw-Hill, New York, 1967, Chap. 12, p. 234.
221
C. Laird and G.C. Smith, Crack propagation in high stress fatigue, Philos. Mag. 8 (1963) 847. J.M. Lee and S.W. Nam, Effect of thermal aging on high temperature low cycle fatigue behavior in AISI 304 stainless steel, Int. J. Damage Mech., in press. T. Magnin, L. Coudreuse and T.M. Lardon, A quantitative approach to fatigue evolution in FCC and BCC stainless steel, Scr. Metall. 19 (1985) 1487. T. Magnin and C. Ramade, Proc. 2nd Int. Conf. on Low Cycle Fatigue and Elasto-plastic Behavior of Materials, Munich, September 7-11, 1987, Elsevier, Amsterdam, 1988, p. 354. P.S. Maiya, Consideration of initiation and crack propagation in low-cycle fatigue, Scr. Metall. 9 (1975a) 1141. P.S. Maiya and B.E. Busch, Effect of surface roughness on low-cycle-fatigue behavior of type 304 stainless steel, Metall. Trans. A6 (1975b) 1761. Z.G. Wang, C. Laird and K. Rahka, The deformation and fatigue characteristic of 1Cr-lMo-0.25V steel cycled at room temperature and an elevated temperature, Mater. Sci. Eng. 73 (1985) 113. J. Wareing and H.G. Vaughan, The relationship between striation spacing, macroscopic crack growth rate, and the low-cycle-fatigue life of a type 316 stainless steel at 625 °C, Met. Sci. 11 (1977) 439. J. Wareing and H.G. Vaughan, Influence of surface finish characteristic of type 316 stainless steel at 460 °C, Met. Sci. 13 (1979) 1.
ELSEVIER
Nuclear Engineering and Design
The effect of creep cavitation on the fatigue life under creep-fatigue interaction Soo Woo
N a m a, S o o C h a n
L e e a, Je M i n L e e b
a Department o f Materials Science and Engineering, Korea Advanced Institute of Science and Technology, 373-1 Kusong-Dong, Yusung-Gu, Taejon, South Korea b Defense Quality Assurance Agency, PO Box 24, Changwon, Kyungnam, South Korea
Abstract
Low cycle fatigue tests have been carried out with three different materials (1Cr-Mo-V steel, 12Cr-Mo-V steel and 304 stainless steel) for the investigation of the effect of surface roughness on the fatigue life. To see the effect systematically, we have chosen those materials which may or may not form grain boundary cavities. Test results show that the continuous fatigue life of 1Cr-Mo-V steel and aged 304 stainless steel with a rough surface is decreased compared with that of the specimens with a smooth surface. These two alloys are found to have no grain boundary cavities formed under creep-fatigue test conditions. On the contrary, the fatigue life of 12Cr-Mo-V steel and solutionized 304 stainless steel in which grain boundary cavities are formed under creep-fatigue test conditions is not influenced by the states of surface roughness. The characteristic test results strongly confirm that the fatigue life of the specimen under creep-fatigue interaction, during which creep cavities are forming, may be controlled by the cavity nucleation and growth processes rather than the process of surface crack initiation.
I. Introduction
In general, fatigue failure occurs by the initiation and propagation of surface cracks, and fatigue life is, therefore, expressed as a summation of the number of cycles to crack initiation and for propagation. In low temperature high cycle fatigue, nearly the entire fatigue life is spent in crack initiation (Laird, 1963), and the fatigue life depends greatly on surface roughness because surface grooves act as fatigue crack initiation sites owing to high stress concentration and fatigue cracks initiate
easily on them (Fluck, 1951; Juvinal, 1967). In high temperature fatigue, however, it has the character of low cycle fatigue (LCF) because it is concomitant with the plastic deformation during fatigue cycling. Also, the conditions of use of high temperature materials are such that creep deformation in addition to fatigue deformation will be undergone. In high temperature fatigue cycling with creep effect, grain boundary cavities are nucleated and the nucleated cavities grow to show high temperature brittle fracture behaviour. The high temperature fatigue behaviour of these materials may be
0029-5493/95/$9.50 © 1995 Elsevier Science S.A. All rights reserved SSDI 0029-5493(94)00829-N
214
S.W. Nam el al./ Nuclear Engineering and Design 153 (1995) 213 221
different from that of LCF behaviour. Even though a fatigue crack is formed at the surface and may propagate into the material, if creep cavitation prevails with a high value of the cavity nucleation parameter, the fatigue life may not be influenced by the fatigue crack but affected by the cavities. Recently, we have proposed a model which predicts the fatigue life of a material under creepfatigue interaction condition. This model explains that vacancies are formed by plastic deformation during the fatigue cycle and these vacancies are clustered to form cavities on the grain boundaries. These nucleated cavities grow during the hold time period. The nucleation of cavities on the grain boundaries is assumed to be strongly dependent on the characteristics of grain boundary precipitates, i.e. each material has its own nucleation parameter. In this study, to see the effect systematically we have chosen three materials (ICr M o - V steel, 12Cr M o - V steel and 304 stainless steel) which may or may not form grain boundary cavities, and LCF tests were conducted on these specimens with different surface roughness, i.e. very smooth and rough surfaces. The results obtained are discussed in terms of the importance of the crack initiation stage in continuous cycling and the effect of the cavitation process in creep-fatigue interaction.
2. Experimental method 1 C r - M o - V steel, 1 2 C r - M o - V steel, solutionized 304 stainless steel and aged 304 stainless steel were used for the fatigue tests. Solution treatment was undertaken at 1373 K for l h followed by water cooling. Some of the solution treated specimens were aged at 1033 K for 50 h. The grain size and microstructure of the materials used in this work are given in Table 1. The specimen was of a cylindrical type with a gauge length of 8 ram. Two different models of surface roughness of the gauge section (smooth and rough surface) were obtained by mechanically polishing the specimens perpendicular to the stress axis with emery papers of grades 80 and 1200 (1500) respectively.
Table 1 The grain size and microstructure of materials used in this work Material
Grain size
Microstructure
1Cr Mo V steel 12Cr Mo V steel 304 stainless steel
501am 100 ~tm 50 ~tm
Tempered bainite Tempered martensite Austenite
(a)
(b)
J
2.5o.m I
, L 2001sin
t
2.5t*m /
" ' 200~m
Fig. 1. Surface roughness profiles of 304 stainless steel specimens with two different surface roughnesses, polished with emery paper of grades (a) 80 and (b) 1200.
The surface roughness profiles in 304 stainless steel which were measured with a Talysurf instrument are shown in Fig. 1 as an example and the maximum values of the peak-to-valley roughness in all the specimens are shown in Table 2. Total strain-controlled uniaxial push-pull fatigue tests were carried out in an electromechanically driven closed-loop controlled Instron model 1362 device with a strain rate of 4 x 10-3 sFor high temperature fatigue tests, a radiant furnace was used to heat the specimen. For the investigation of the phenomenon of creep-fatigue interaction, a 10 30min hold at tensile peak strain was applied. The characteristic morphology of the fatigue surface and the cavitation behaviour during fagitue cycling were examined by scanning electron microscopy (SEM).
215
S.W. Nam et al. / Nuclear Engineering and Design 153 (1995) 213-221 Table 2 The maximum value of the peak-to-valley roughness height in the specimen used in this work Material
Roughness for emery paper grade 80
Roughness for emery paper grade 1200
ICr M o - V steel 1 2 C r - M o - V steel 304 stainless
8.0 ~tm 8.8 lam 13 I~m
0.4 lam 0.9 tam 0.3 lam "
Polished with emery paper grade 1500.
3. Experimental results
10 304 S-S 873K
o. CO
3.1. Fatigue life with surface roughness
<3
3.1.1. I C r - M o - V steel The continuous LCF and creep-fatigue lives of the 1 C r - M o - V steel with two different surface roughness modes at 823 K are shown in Fig. 2. In the continuous LCF and creep-fatigue tests (10 min hold at tensile peak strain), the fatigue life of the specimen with the rough surface decrased compared with that of the smooth surface as shown in Fig. 2. 3.1.2. 304 stainless steel The continuous LCF lives of the solutionized and the aged 304 stainless steel with different roughnesses at 873 K are shown in Fig. 3. The fatigue life of the specimen with a rough surface decreased with increasing surface roughness as shown in Fig. 3. 10 -'
c
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cont., cont., cont., cont.,
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. . . . . . .
;4'
'
. . . . . .
io'
Critical Number of Cycles ( N c r ) Fig. 3. Relationship between plastic strain range and fatigue life in 304 stainless steel specimens with two different surface roughnesses under continuous cycling.
In the creep-fatigue test (10 min hold at tensile peak strain) the fatigue life of the solutionized specimen is found to be independent of the surface roughness conditions; however, that of the aged specimen is observed to decrease with increasing roughness as shown in Fig. 4.
3.1.3. 1 2 C r - M o - V steel The test results of continuous LCF and creepfatigue (30 min hold at tensile peak strain) of the 12Cr-Mo-V steel with two different surface roughness modes at 873 K are shown in Fig. 5. In the continuous LCF and creep-fatigue tests, the fatigue lives of the specimens are also found to be independent of the surface conditions as shown in Fig. 5.
.c__ ~ 10 -,
_o
o •
I I ''01
~
'
of Cycles to Failure
Fig. 2. Relationship between plastic strain range and fatigue life in I C r - M o - V steel specimens with two different surface roughnesses.
3.2. Observation of the fracture surface To investigate the nature of the crack propagation behaviour of the specimens with the different
216
S.W. Nam et al. / Nuclear Engineering and Design 153 (1995) 213 -221
1 0 -I
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Fig. 6. OM images of the crack propagation modes of l Cr Mo V steel at 823 K: (a) continuous cycling; (b) creep fatigue cycling.
Fig. 4. Relationship between plastic strain range and fatigue life in 304 stainless steel specimens with two different surface roughnesses under creep fatigue cycling.
10 -'
12~Cr Mo V Steel Cont. Cycling 875K, Air C
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Number of Cycles to Foilure Fig. 5. Relationship between plastic strain range and fatigue life in 1 2 C r - M o - V steel specimens with two different surface roughnesses.
surface roughness modes, the specimens were fatigued up to the number of cycles required to attain an about 5% drop in tensile peak load and then polished. For the observation of the crack propagation mode, the specimens were observed by optical micrography (OM) and optical micrographs of solutionized 304 stainless steel and I C r - M o - V steel, as the typical examples, are shown in Fig. 6 and Fig. 7 respectively. All the other results are listed in Table 3, It is assumed that the crack initiation phase is up to the crack size of 100 lam (one or two grain sizes).
Fig. 7. OM images of the crack propagation modes of solutionized 304 stainless steel at 823 K: (a) continuous cycling: (b/ creep fatigue cycling.
In the continuous cycling, fatigue crack propagation modes of all the materials tested in this study were transgranular without any grain boundary cavitation as shown in Fig. 6(a) and Fig. 7(a). In the creep-fatigue cycling, however, intergranular crack growth was observed in all the materials except 1Cr-Mo-V steel, in which cracks propagate transgranularly without cavitation damage as shown in Fig. 7(b) and Fig. 6(b). The nucleation of cavities on the grain boundaries is considered to be strongly related to the inter-
S.W. Nam et al. / Nuclear Engineering and Design 153 (1995) 213-221
217
Table 3 Variation in the fatigue life and fracture mode of materials with testing conditions Material
Temperature
Continuous cycling
Creep-fatigue cycling
1Cr-Mo V steel
823 K
12Cr-Mo-V steel
873 K
Solutionized 304 stainless steel Aged 304 stainless steel
873 K
Transgranular, Nr(R) < Nr(S) Transgranular, Nf(R) = Nf(S) Transgranular, Nr(R) < Nr(S) Transgranular, Nr(R) < Nr(S)
Transgranular, Nr(R) < Nr(S) Intergranular, Nr(R) = Nr(S) Intergranular, Nr(R) = Nf(S) Transgranular, Nr(R) < Nr(S)*
873 K
Nr(R), fatigue life of the specimen with a rough surface; Nr(S), fatigue life of the specimen with a smooth surface; Nr(S)*, intergranular.
granular crack growth modes in the creep-fatigue cycling.
4. Discussion
4.1. Effect o f surface roughness on the continuous low cycle fatigue
In the continuous LCF tests, the fatigue life of all the materials tested in this study is decreased with increasing surface roughness except for 1 2 C r - M o - V steel in which it is independent of surface modes as shown in Fig. 5. In 1 2 C r - M o V steel as a special case, fatigue cracks are nucleated at the grain boundaries of the specimen regardless of the surface roughness; there is no difference in the fatigue life with surface roughness. Also, the slopes of both plots with surface roughness in each material are approximately the same over the plastic strain range tested. If the fatigue life is dominated by the crack propagation the slope for the test results with a rough specimen should be steeper than that of a smooth specimen because of the general concept (Laird, 1963; Maiya, 1975a) that the life fraction for crack initiation decreases with increasing strain range. This is inconsistent with the present results. The experimental results suggest that LCF life is strongly influenced by the crack initiation process even for a high strain range and the fraction of the fatigue life for crack initiation is nearly the
same over the plastic strain range tested. This agrees qualitatively with Dowling's observations (Dowling, 1977) that the life fraction for any given surface crack size (0.01 inch) is independent of the life in A533B steel. To date many authors (Magnin, 1985, 1988; Maiya, 1957b) have investigated surface roughness effects to evaluate the importance of the crack initiation stage in LCF. Maiya and Busch tested type 304 stainless steel at 866 K and suggested that the fatigue life decreased with increasing surface roughness owing to the reduction in the number of cycles to the crack initiation (Maiya, 1975b). Magnin et al. reported that the fatigue cycles up to nucleation and evolution of micro-cracks at the specimen surface of austenitic and ferritic stainless steel represented most of the LCF life regardless of the strain amplitude applied (Magnin, 1985, 1988). However, Wareing and Vaughan argued that the LCF behaviour was dominated by the crack propagation process (Wareing, 1977) and they insisted that the difference in fatigue life between machined and electropolished specimens of type 316 stainless steel resulted from the different initiated crack shapes for the two different surface finish modes at 673 K (Wareing, 1979). To investigate the importance of crack initiation in high temperature LCF of 1 C r - M o - V steel, the smooth surface specimen was fatigued at a plastic strain range of 0.7% for a number of cycles corresponding to 60% of the fatigue life at
218
S.W. Nam et al. Nuclear Engineering and Design 153 (1995) 213- 221
Fig. 8. SEM image of the fracture surface of 1Cr Mo V steel specimen with smooth surface, fatigued for 60% of the fatigue life at 823 K with a plastic strain range of 0.7% (the arrows indicate a crack propagated at 823 K).
this strain range. The specimen was fractured by impact in liquid nitrogen and the fracture surface was observed by SEM, as shown in Fig. 8. A few microcracks (a crack size of 80 ~tm) nucleated at the specimen surface are shown on the fracture surface. From these results, the fatigue crack initiation is very important in the continuous LCF in which grain boundary cavities are not nucleated.
4.2. Effect of surface roughness on the creep-fatigue test In the creep-fatigue test holding at tensile peak strain, fracture modes are divided into two groups. One is the case in which a fatigue crack propagates transgranularly similarly to the case of continuous cycling because of no cavitation damage during creep-fatigue cycling; the other is the case in which a fatigue crack propagates in an intragranular mode due to the cavitation damage during creep-fatigue cycling. The reduction in the fatigue life in the case of no cavitation damage is due to the earlier crack nucleation and the accelerated crack propagation rate because of the recovery effect ahead of the crack tip during a hold time period. Crack nucleation is also important in this case. The life of the specimen with a rough surface is nearly half that of the specimen with a smooth surface; this is considered to be due to the earlier crack initiation at the surface roughness grooves
in the case of 1 C r - M o - V steel. In this steel, creep cavitation damage is not observed during the creep-fatigue test because of the non-existence of the cavity nucleation sites, MnS particles, in the grain boundaries (Wang, 1985). Similarly, the fatigue life of the aged 304 stainless steel is decreased with increasing surface roughness because of the delayed cavitation process owing to the large spacing between the grain boundary precipitates ( M 2 3 C 6 ) which act as cavity nucleation sites. Also, the crack propagation mode of the specimen with the rough surface is transgranular because of the earlier crack initiation at the surface roughness grooves and the delayed cavitation process. However, the other materials such as 12Cr M o - V steel and solutionized 304 stainless steel show a different aspect in the surface roughness effect because of the creep cavitation damage. In case of 1 2 C r - M o V steel, the cavities are nucleated not only on prior austenite grain boundaries but also on lath martensite boundaries, and the nucleated cavities grow during every tensile hold time period. The carbides in the prior austenite grain boundaries and the lath martensite boundaries in 12Cr Mo V steel act as the cavity nucleation site. Similarly, in solutionized 304 stainless steel, many carbides (M23C6) are precipitated at the grain boundaries during the creep-fatigue cycling, and these carbides act as cavity nucleation sites. This high density of grain boundary carbide accelerates a cavitation process (Lee, in press). In the above two materials, the reduction in the fatigue life with surface roughness is considered to be caused by the increased crack propagation rate due to the interaction between the fatigue damage as surface crack initiation and growth and the creep damage as cavity nucleation and growth at the grain boundaries during hold time period. However, the surface condition which strongly influences the fatigue crack initiation is known to have little effect on the fatigue life because the cavitation damage is more effective than the fatigue damage in these two materials. It is thought that the fatigue life is nearly the same as shown in Figs. 4 and 5, because the creep cavitation process takes a decisive role in the fatigue life.
219
s.w. Nam et al. / Nuclear Engineering and Design 153 (1995) 213-221 4.3. Life prediction in c r e e p - f a t i g u e test
Recently, we have proposed a model (Hong, 1985) which predicts the fatigue life of a material in which the creep cavitation damage is dominant under creep-fatigue cycling. This model explains that vacancies are formed by plastic deformation during the fatigue cycle and these vacancies are clustered to form cavities on the grain boundaries. Also, it is assumed that the number of cavities generated in one cycle is proportional to the plastic strain range. The number n of nucleated cavities during cyclic loading per unit area of the grain boundary is represented by (Lee, in press) n = P A% N
(1)
where P, AEp and N are the nucleation factor per unit area of grain boundary, the plastic strain range and the number of cycles respectively. These generated cavities are assumed to grow during the hold time period in tension straining by grain boundary diffusion of vacancies. The Hull and Rimmer model (Hull, 1959) for diffusional growth of cavities at the grain boundary provides a good approximation of cavity growth and the stress term of the model has to be modified as a function of hold time because of load relaxation during tensile strain hold (Lee, in press): d A /dt = ~Dg&f2a( t) /lk T
(2)
where A is the area of a given cavity, l is the cavity spacing, Ogis the grain boundary diffusion coefficient, f2 is the atomic volume, a(t) is the relaxed stress variation during the hold time, and k and T have their usual meanings. From the total cavitated area per unit area of grain boundary up to N cycles and a failure criterion, which is that the load-carrying capacity is drastically reduced by coalescence of grain boundary cavities, the final form of the fatigue life prediction for a material of which creep-fatigue failure is controlled by creep damage rather than by a process of fatigue crack initiation and propagation is given by (Lee, in press)
(3)
lSkv
15gk~
67P
gl¢
Fig. 9. SEM image showingthe fatigue crack initiation mode in solutionized304 stainlesssteel at 873 K after ( 1/3)N~,.under creep fatigue cycling. where Qg is the activation energy of the grain boundary diffusion, t is the hold time and C is a constant including the critical cavitated area. The proposed equation has been checked using the experimental results of 304 stainless steel and 12Cr M o - V steel in which failure is controlled by the creep cavitation damage. The fatigue cracks in solutionized 304 stainless steel are nucleated at grain boundaries owing to grain boundary cavitation as shown in Fig. 9. To investigate whether the cavities are nucleated continuously in every cycle, the specimens of solutionized 304 stainless steel were fatigued up to 1/3 and 2/3 of the fatigue life and near fatigue life respectively and then the specimens chilled in liquid nitrogen were broken by impact. The impact fracture surfaces were observed by SEM and are shown in Fig. 10. Fig. 10 shows that the number of cavities increased with increasing creep-fatigue cycling, and this result means that cavities are nucleated continuously with the fatigue cycles. All the above results mean that the proposed equation is applicable to the two materials. Also, the nucleation of cavities on the grain boundaries is assumed to be strongly dependent on the characteristics of grain boundary precipitates, i.e each material has its own nucleation parameter. The nucleation parameters of the 1 2 C r - M o - V steel, the solutionized 304 stainless steel and the aged 304 stainless steel are 4.5 x 10 ~°, 6.43 x 101~ and 1.27 × l0 II respectively. Using the above nucleation parameters, we predict the creep-fatigue life of each material with the experimental conditions and the predicted
220
S. I,IL Narn et al. / Nuclear Engineering and Design 153 (1995) 213-221
/
304 S-S , 873K sol aged t.(rnin.) 10 3
.~
*
30
•
50 /
/ / /
.
•
n
10 2
,
,
,
,
~
, ~,~
10 3
I0
Experimental
Life
Fig. 11. Predicted life by creep cavitation only vs. experimental life on 304 stainless steel under creep fatigue cycling at 873 K.
1000
8OO Q
,._e Fig. 10. SEM images showing grain boundary cavities on intergranular fracture surfaces produced at liquid nitrogen temperature by impact in solutionized 304 stainless steel, fatigued at 873 K: (a) (1/3)Nor; (b) (2/3)N~; (c) near Nor.
~.~ ._u ~XD 400600
5. Conclusions (1) The fatigue life of the specimens with a rough surface decreases compared with that of the
o
o
EL ~/
200
creep-fatigue lives are compared with experimental lives, as shown in Fig. 11 and Fig. 12. These plots show that the predicted lives are in good greement with the experimental lives. From these results, it may be understood that creep-fatigue failure is controlled by creep damage only rather than by a process of fatigue crack initiation and propagation. It is thought that the proposed equation in which only cavity nucleation and growth are considered is good for life prediction in creep-fatigue cycling.
o
0
/7 /
873K, Air
/
I 200
o : th=lOmin.
:
I 400
..
~ 600
Experimentol
I 800
1000
Life
Fig. 12. Predicted life by creep cavitation only vs. experimental life on 1 2 C r - M o - V steel under creep-fatigue cycling at 873 K.
specimens with a smooth surface in continuous LCF. This decrease in the fatigue life with surface roughness is thought to be due to the reduction in the number of cycles to the crack initiation caused by the surface roughness. Therefore, it is thought that the crack nucleation is considered to be an important process in the fatigue life for the materials without grain boundary cavitation.
S. IV. N a m e t al. / Nuclear Engineering and Design 153 (1995) 213-221
(2) The fatigue life of the materials in which the creep cavitation process takes place is independent of the surface roughness conditions because the creep-fatigue failure is controlled mostly by the creep damage rather than by a process of fatigue crack initiation. From this result, it is thought that the proposed life prediction formula based on considering the creep cavitation only is reasonable.
References N.E. Dowling, Crack growth during low-cycle-fatigue of smooth axial specimens, ASTM 637, 1977, p. 97 (ASTM, Philadelphia, PA). P.G. Fluck, The influence of surface roughness on the fatigue life and scatter of test results of two steels, ASTM 51 (1951) 584. J.W. Hong, S.W. Nam and K.T. Rie, A model for life prediction in low-cycle fatigue with hold time, J. Mater. Sci. 20 (1985) 2763. D. Hull and D.E. Rimmer, The growth of grain-boundary voids under stress, Philos. Mag. 4 (1959) 673. R.C. Juvinal, Effect of load type, specimen size and surface finish, Stress, Strain and Strength, McGraw-Hill, New York, 1967, Chap. 12, p. 234.
221
C. Laird and G.C. Smith, Crack propagation in high stress fatigue, Philos. Mag. 8 (1963) 847. J.M. Lee and S.W. Nam, Effect of thermal aging on high temperature low cycle fatigue behavior in AISI 304 stainless steel, Int. J. Damage Mech., in press. T. Magnin, L. Coudreuse and T.M. Lardon, A quantitative approach to fatigue evolution in FCC and BCC stainless steel, Scr. Metall. 19 (1985) 1487. T. Magnin and C. Ramade, Proc. 2nd Int. Conf. on Low Cycle Fatigue and Elasto-plastic Behavior of Materials, Munich, September 7-11, 1987, Elsevier, Amsterdam, 1988, p. 354. P.S. Maiya, Consideration of initiation and crack propagation in low-cycle fatigue, Scr. Metall. 9 (1975a) 1141. P.S. Maiya and B.E. Busch, Effect of surface roughness on low-cycle-fatigue behavior of type 304 stainless steel, Metall. Trans. A6 (1975b) 1761. Z.G. Wang, C. Laird and K. Rahka, The deformation and fatigue characteristic of 1Cr-lMo-0.25V steel cycled at room temperature and an elevated temperature, Mater. Sci. Eng. 73 (1985) 113. J. Wareing and H.G. Vaughan, The relationship between striation spacing, macroscopic crack growth rate, and the low-cycle-fatigue life of a type 316 stainless steel at 625 °C, Met. Sci. 11 (1977) 439. J. Wareing and H.G. Vaughan, Influence of surface finish characteristic of type 316 stainless steel at 460 °C, Met. Sci. 13 (1979) 1.