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20th International Conference on Structural Mechanics in Reactor Technology (SMiRT 20)Espoo, Finland, August 9-14, 2009

SMiRT 20-Divis ion 1, Paper 3135

1

Fatigue curve and stress strain response for stainless steel

J ussi Solina, Gerhard Nagelb and Wolfgang Mayingerb

a

VTT, Finland, , e-mail: [email protected] Kernkraft GmbH, Tresckowstrasse 5, Hannover, Germany

Keywords: Fatigue, Stainless steel, Design curve.

1 ABSTRACT

Niobium stabilized austenitic stainless steel (X6CrNiNb1810 mod) was studied in strain controlled fatiguetests ranging from low cycle (LCF) to high cycle (HCF) regime in RT air. An experimental strain life curvewas determined as a base line for component specific evaluations and for comparison with the Langer andChopra curves, which are the basis of the ASME I II and NRC RG 1.207 design criteria. Stress strainresponses were carefully measured to clarify the strain amplitude dependent fatigue behaviour of this steel.

In the LCF regime our data lie within a common scatter band between the Langer and Chopra curves,but an endurance limit was observed at εa ≈ 0,19 %. This is due to notable secondary hardening at low strainamplitudes. In HCF regime the Chopra curve and NRC RG 1.207 become highly conservative for thismaterial. Relevance of all proposed design curves should be carefully considered.

2 INTRODUCTION

Fatigue design curves given in the ASME Code Section III are derived from reference mean curves proposedby Langer . They are based on strain controlled low cycle fatigue tests in room temperature. Arbitrary designmargins, 20 against life and 2 against strain, were considered appropriate for ensuring transferability of thedata to plant components, as by ASME (1972). Similar curves are included also in the German and Frenchdesign codes KTA and RCC-M.

The ASME III design by analysis philosophy and local strain approach assume that the designer hasrelevant material data available. Although generalized design curves have been included in the codes toreduce need for material testing, choice and applicability of the code curve – or an experimental curve – forthe particular application remains the responsibility of the designer. Consideration of operation environmentis a good example on this. The code itself does not give specific curves or quantitative factors for adoptinginfluence of reactor coolant to fatigue calculation. Moderate environmental effects are accounted for throughthe design curve definition (within the margin of 20 in life), but the responsibility of considering eventualenvironmental effects was left to the designer as clearly stated in the ASME (1972) Criteria Document forthe ASME III Design by analysis procedure as follows: “protection against environmental conditions such as

corrosion and radiation effects are the responsibility of the designer”

It is obvious that the designer may choose to use an appropriately determined and more relevantexperimental curve. This is literally recommended by STUK (2002) in the Finnish YVL guide for ensuringstrength of NPP pressure devices. According to YVL guide 3.5, “fatigue assessment shall be based on S-N -curves applicable to each material and conditions”.

The development of the ASME code was primarily aiming to prevent catastrophic fractures of pressurevessels and the fatigue assessment was focusing on severe but rare thermal transients that can cause notablelow cycle fatigue damage in heavy equipment. Later on, higher numbers of small stress cycles acting in thepiping have been addressed, in particular for small bore pipes. Simultaneously fatigue tests have beenconducted to longer lives. Encouraged by new experimental data and the proposal by Chopra and Shack(2007), the NRC (2007) endorsed a new air curve for stainless steels as part of a Regulatory Guide for new

designs un USA. Data and regression curves in line with the ANL/NRC reference curve (often referred as“Chopra curve”) have been published by Jaske (1977), Higuchi (2004), Solomon (2004), Faidy (2008) et al.

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It should be noted that in the LCF regime all air data for austenitic stainless steels lie practically within acommon scatter band and the later proposed curves do not much differ from the Langer curve. The

difference grows to an order of magnitude in life in the HCF regime (Nf >105 cycles), where the Chopracurve becomes more conservative. In other words, the ASME code stainless steel air curve has become asubject of debate because of gradually expanding scope of fatigue assessments to HCF region. But HCFtesting of stainless steels is not simple and valid data is still rare.

Aim of the current study is to test applicability of the existing and proposed design criteria. For thatpurpose we selected a material batch, which would have been ready for plant use, e.g., in a PWR surge line.An experimental strain life fatigue curve was determined to be used as a base line for component specificevaluations and for comparison with the Langer and Chopra curves, which are the basis of the ASME III andNRC RG 1.207 design curves. Monotonic and cyclic stress strain responses were carefully analyzed toclarify the deformation and fatigue mechanisms for this steel.

3 EXPERIMENTAL

Solution annealed Niobium stabilized austenitic stainless steel (X6CrNiNb1810 mod) was received as aφ 360x32 mm pipe, which fulfils all KTA material requirements for primary components in BWR and PWR.

Chemical composition of the test material is given in Table 1. The grain size in this pipe varies so that thematerial report classified 50 % to ASTM 0-1 and 50 % to ASTM 2-3.

Smooth round bar specimens were turned and polished from samples extracted all round thecircumference of the pipe. Tensile and CSSC test specimens were selected from each quarter segment andLCF specimens were randomly picked round the pipe. The LCF specimen dimensions are shown in Fig. 1.

Table 1. Composition of test material (wt %).

C N Si Mn Cr Ni Mo Nb P S

0,031 0,021 0,235 1,885 17,30 10,29 0,405 0,357 0,030 0,004

20

R

2

35

113

!

8

!

1 5

20

R

2

35

113

!

8

!

1 5

Figure 1. Fatigue specimen and dimensions in mm.

Tensile and fatigue testes were both performed in a MTS 100 kN rig with precision alignment grips anddigital control unit. Alignment of load train was adjusted with the help of strain gauged specimens according

to the ASTM E

1012-05 procedure. MTS extensometers with gage lengths of 50

mm (tensile tests) and 8

mm(fatigue) were used for measuring strains. The gauge section of the tensile specimens wasφ 8 mmx58 mm.

Five tensile tests were performed with nominal strain rates varying between 2·10-5 and 2·10-3. Partialunloadings were introduced to measure change of apparent elastic modulus as function of strain. Until onsetof necking true stress (σ =S·(1+e) and strain (ε =ln(1+e) were obtained from the nominal values (S and e).

Beyond instability the local strains were deduced from periodic optical measurements of necking.

Strain controlled low cycle fatigue tests were performed according to the ASTM E 606 procedure usingsinusoidal waveform. An average strain rate of 0,021/s was used for constant amplitude tests (e.g. 1 Hz for εa

=0,5 %). For longest HCF tests, the frequency was increased to 6 Hz during the late secondary hardeningphase with decreasing plastic strain.

For determination of cyclic stress strain curves (CSSC) a spectrum straining method developed by Solin(1986) was used. The frequency of sinusoidal ramps was varied to keep the average strain rate constant at0,011/s.

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A simplified definition of fatigue life (Nf,25) as the number of cycles to 25 % drop of peak stress from itsabsolute maximum was adopted to avoid practical problems with variable cyclic softening and hardeningbehaviour of stainless steels. However, all final load drop phases were so short that the obtained lives areinsensitive to the selected failure criterion.

4 RESULTS AND DISCUSSION

4.1 Stress strain response – tensile properties and hysteresis loops

The results of tensile tests together with the material specification report results for the particular pipe andaverages for the melt are summarized in Table 2. The elongation (A5) varied from 57 % to 74 %, the largestvalues obtained by the lowest applied strain rates. In spite of the well known correlation of instantaneousflow stress and pulling rate, an inverse correlation was seen in ultimate tensile strength, because of longerrange of strain hardening with low rate. Partial unloadings seemed to have minimal effect on the tensilecurves. The nominal and true stress strain curves are shown in Fig. 2.

Table 2. Tensile test results in current study and in the material specification report.

data source E Rp0,2 UTS

minimum of 5 tests 195 GPa 224 MPa 535 MPa

maximum of 5 tests 201 GPa 249 MPa 559 MPa

average of 5 tests 197 GPa 238 MPa 544 MPa

material report / pipe - 239 MPa 548 MPa

material report / melt - 251 MPa 544 MPa

X6CrNiNb1810 mod

0

250

500

750

1000

1250

1500

1750

2000

0 20 40 60 80 100 120 140 160 180

monotonic pull

intermediate rate

unloads, slow rate

true stress strain

n o m i n a l / t r u e s t r e s s

M P a

nominal / true strain %

true stress strain curve

nominal stress strain curves

X6CrNiNb1810 mod

0

250

500

750

1000

1250

1500

1750

2000

0 20 40 60 80 100 120 140 160 180

monotonic pull

intermediate rate

unloads, slow rate

true stress strain

n o m i n a l / t r u e s t r e s s

M P a

nominal / true strain %

true stress strain curve

nominal stress strain curves

Figure 2. Monotonic nominal and true stress strain curves.

As the strain amplitude is fixed in fatigue tests, stress amplitude and mean stress depend on the materialresponse. The stress strain responses during constant amplitude tests were studied through hysteresis loopsand hardening softening curves. Hysteresis loops at half-life (N =N25 / 2) are shown in Fig. 3a. The generaltradition of reporting the half-life data is applied here, although this practice may be questioned for stainlesssteels, because the cyclic response doesn’t stabilize. A minimum stress response is found at 0,2 % strainamplitude due to secondary hardening beginning before half-life for smaller amplitudes.

The hysteresis loops in Fig. 3a are positioned to reveal different strain hardening paths in tensile

direction. According to Bayerlein et al. (1987) non-Masing behaviour (not coinciding loops) indicates thatthe dislocation microstructures and/or cyclic deformation mechanisms are amplitude dependent. But for thismaterial the deformation mechanisms change even during a single test. Fig. 3b shows evolution of hysteresis

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loops during a test at or just below the endurance limit. Inelastic strains are notable even in this HCF test,

which lasted over 12 million cycles. More than one third of the total strain ( εa =0,185 % ) was inelastic

before secondary hardening eliminated most of the hysteresis. See chapter 4.3 for further description of cyclic softening and hardening.

MPa

0

100

200

300

400

500

600

700

0,0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0

1,00%

0,80%

0,50%

0,40%

0,30%0,22%

0,20%

0,195%

0,185%

%

MPa

0

100

200

300

400

500

600

700

0,0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0

1,00%

0,80%

0,50%

0,40%

0,30%0,22%

0,20%

0,195%

0,185%

%

-300

-200

-100

0

100

200

300

-0,20 -0,15 -0,10 -0,05 0,00 0,05 0,10 0,15 0,20

begin 1-2hard 28

soft 80k

sec.h. 1M

EOL 10M

%

MPa

!a =0,185%

N25=12·106

-300

-200

-100

0

100

200

300

-0,20 -0,15 -0,10 -0,05 0,00 0,05 0,10 0,15 0,20

begin 1-2hard 28

soft 80k

sec.h. 1M

EOL 10M

%

MPa

!a =0,185%

N25=12·106

Figure 3. Hysteresis loops at half-lives of different tests (a) and at different stages of a single test (b).

4.2 Apparent elastic modulus

The elastic modulus was determined in tensile tests to 197 GPa, but unloading – reloading modulus is not

constant as shown in Fig. 4. The apparent elastic modulus reaches a maximum just when yielding starts atabout 0,1 % total strain and then gradually decreases by about 5 % until 1 % strain and by about 15 % until 10

% strain. Note that the reduction of specimen cross section was accounted for. The modules based onnominal stresses decrease even more. At high strains, changes of grain orientations may contribute, but atlow strains the changes must be mainly due to internal stresses. Austenitic stainless steels are known forstrain hardening and ability to generate large internal stresses by plastic strain.

X6CrNiNb1810 mod

( cross section corrected ) .

160

170

180

190

200

210

0,0 0,2 0,4 0,6 0,8 1,0

first pull 20-70 MPa

repeated unloads

single unloads

X6CrNiNb1810 mod

( cross section corrected )

130

150

170

190

210

0 10 20 30

repeated unloads

strain %

A p p a r e n t m o d u

l e

G P a

X6CrNiNb1810 mod


160

170

180

190

200

210

0,0 0,2 0,4 0,6 0,8 1,0


repeated unloads

single unloads

X6CrNiNb1810 mod


130

150

170

190

210

0 10 20 30

repeated unloads

strain %


l e

G P a

X6CrNiNb1810 mod


160

170

180

190

200

210

0,0 0,2 0,4 0,6 0,8 1,0


repeated unloads

single unloads

X6CrNiNb1810 mod


130

150

170

190

210

0 10 20 30

repeated unloads

strain %


l e

G P a

Figure 4. The measured apparent elastic modules as function of applied tensile strain.

Evolution of microstructure and internal stresses cause changes in hysteresis loop shape. Kuhlmann-Wilsdorf and Laird (1979) divided the shear stresses acting on the slip planes into thermal (rate dependent)

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“friction stress” and athermal (elastic type) “back stress” rising from short distance dislocation motion. Theyused friction stress and back stress components for modelling hysteresis loops for stainless steels.

We assume that – together with the real elastic modulus – microscopic or mesoscopic internal (back)stresses play a role in evolution of modulus (linear part of hysteresis loops) during cyclic deformation. Adrop of modulus was systematically observed at all strain amplitude levels and during the whole fatigue lifeas demonstrated in Fig. 5. Reduction of modulus scaled off or even turned to increase very late during the

longest tests, but it is worth of noting that even there the reducing trend continued also during the secondaryhardening phase. The modulus change was not in phase with softening or hardening.

170

175

180

185

190

195

200

1,E+00 1,E+01 1,E+02 1,E+03 1,E+04 1,E+05 1,E+06 1,E+07 1,E+08

up 0,185% dwn 0,185%

up 0,200% dwn 0,200%

up 0,220% dwn 0,220%

up 0,300% dwn 0,300%

up 0,500% dwn 0,500%

up 0,800% dwn 0,800%

up 1,000% dwn 1,000%

GPa

1 10 100 1000 104 105 106 107 108N

170

175

180

185

190

195

200

1,E+00 1,E+01 1,E+02 1,E+03 1,E+04 1,E+05 1,E+06 1,E+07 1,E+08

up 0,185% dwn 0,185%

up 0,200% dwn 0,200%

up 0,220% dwn 0,220%

up 0,300% dwn 0,300%

up 0,500% dwn 0,500%

up 0,800% dwn 0,800%

up 1,000% dwn 1,000%

GPa

1 10 100 1000 104 105 106 107 1081 10 100 1000 104 105 106 107 1081 10 100 1000 104 105 106 107 108N

Figure 5. Change of modulus in increasing and decreasing ramps during cyclic straining.

4.3 Cyclic softening and hardening

Change of cyclic stress response during a test is best visualized by softening / hardening curves. They showstress responses (Sa =ΔS / 2) as function of number of cycles during constant strain amplitude tests, Fig. 6.

MPa

200

220

240

260

280

300

320

340

360

380

1,E+00 1,E+01 1,E+02 1,E+03 1,E+04 1,E+05 1,E+06 1,E+07

1,0 0,8%

0,5 0,4 0,3%

0,22 0,2%

0,195%

0,190%

0,185%

1 10 100 1000 104 105 106 107N

εaMPa

200

220

240

260

280

300

320

340

360

380

1,E+00 1,E+01 1,E+02 1,E+03 1,E+04 1,E+05 1,E+06 1,E+07

1,0 0,8%

0,5 0,4 0,3%

0,22 0,2%

0,195%

0,190%

0,185%

1 10 100 1000 104 105 106 1071 10 100 1000 104 105 106 107N

εa

Figure 6. Cyclic hardening and softening curves.

Consistent and well repeatable cyclic stress strain responses including initial hardening, softening andsecondary hardening were observed. Dislocation density and internal stresses are increased during the first

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10 to 20 hardening cycles. Subsequent dislocation activity utilizes the driving force and repeatedopportunities to optimize the dislocation structures. This is seen as cyclic softening, which lasts for 2000 to100 000 cycles depending on the strain amplitude. During softening the stress response may be loweredbelow that for the soft annealed material.

Softening is followed by secondary hardening. It is not clear, how onset of secondary hardening isdetermined after up to 105 softening cycles, but it is astonishingly well predictable as a function of

amplitude. Secondary hardening is soon interrupted by crack growth and fracture at strain amplitudes above0,2 %, but below this threshold, most of the fatigue life is spent in secondary hardening and the stressresponse will notably increase. As shown in Figs. 3b and 6 (see lowest curve in Fig. 6), the secondaryhardening phase began after 80 000 cycles and occupied 99 % of the 12 million cycles endurance of a

specimen tested at εa =0,185 %.

4.4 Cyclic stress strain curves

Stress strain data can also be presented as cyclic stress strain curves (CSSC), which give the stress amplitudeas function of strain amplitude. Cyclic stress strain curves (CSSC) are traditionally defined as half liferesponse of the material in constant amplitude as shown in Fig. 7a. The half life is assumed to representstabilized response occupying most of the fatigue life. But such stabilisation doesn’t occur for austenitic

stainless steels. A peak is observed below 0,2 % strain amplitude, where secondary hardening is prominent.

Solin (1989) demonstrated applicability of spectrum straining to determine CSSC for stainless steels,which do not stabilize in cyclic deformation. By spectrum straining method with a single specimendetermined CSSC represents a certain microstructure (dislocation density and configuration), which dependsmostly on the largest strains in spectrum. Therefore, realistic amplitude scales should be applied. Resultswith different amplitude scales are shown together with the constant amplitude data in Fig. 7a.

By the spectrum straining method cyclic hardening or softening can be modelled as evolution of theCSSC in different phases of the test as demonstrated in Fig. 7b. The six first cycles are in line with themonotonic tensile curve. A reasonably stabilized CSSC for the early test condition is reached already within40 first cycles, but cyclic hardening gradually increases the curve until crack growth takes over.

0

50

100

150

200

250

300

350

0,0 0,2 0,4 0,6 0,8 1,0

1,00% half-life 3650

0,50% half-life 180000,25% half-life 125000

const.ampl. half-life

MPa

%

0

50

100

150

200

250

300

350

0,0 0,2 0,4 0,6 0,8 1,0




MPa

%

0

50

100

150

200

250

300

350

0,0 0,2 0,4 0,6 0,8 1,0




MPa

%

0

50

100

150

200

250

300

350

0,0 0,2 0,4 0,6 0,8 1,0

end-of-life (N=7300)

hardest (N=5750)

half-life (N=3650)

early phase (N=500)

first block (N=39..100)

first cycles (N=21..38)



monotonic tensile test

MPa

%

0

50

100

150

200

250

300

350

0,0 0,2 0,4 0,6 0,8 1,0

end-of-life (N=7300)

hardest (N=5750)

half-life (N=3650)

early phase (N=500)

first block (N=39..100)




monotonic tensile test

MPa

%

Figure 7. Summary of half-life CSSC’s (a) and evolution of CSSC in a spectrum straining test (b).

4.5 Fatigue lives

The obtained fatigue lives are summarized in Figs. 8-9. The individual data points are shown in Fig. 8together with the data set used by Langer to derive the stainless steel reference curve and the ASME designcurve. Our data is within the scatter band of Langer’s data set and well above the ASME design curve.

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0,1

1,0

1,E+03 1,E+04 1,E+05 1,E+06 1,E+07 1,E+08N

s t r a i n a m p l i t u d e

%

Langer data

Raw data N(25%)

Interrupted test

Mean curve

Langer curve

ASME design curve

104 105 106 107 108103 N

0,1

1,0

1,E+03 1,E+04 1,E+05 1,E+06 1,E+07 1,E+08N

s t r a i n a m p l i t u d e

%

Langer data

Raw data N(25%)

Interrupted test

Mean curve

Langer curve

ASME design curve

104 105 106 107 108103 N104 105 106 107 108103 104 105 106 107 108103 N

Figure 8. Strain life data compared to ASME 3 design curve and air curve in US NRC Reg.Guide 1.207.

104 105 106 107 108103 N

1,0

0,7

0,5

0,2

0,1

0,07

0,05

%

εaN/12

/2εa

[NUREG 6909 ][NRC RG 1.207 ]

[Current data & Chopra margins ]

104 105 106 107 108103 N

1,0

0,7

0,5

0,2

0,1

0,07

0,05

%

εaN/12

/2εa

104 105 106 107 108103 N104 105 106 107 108103 104 105 106 107 108103 N

1,0

0,7

0,5

0,2

0,1

0,07

0,05

%

εa

%

εaεaN/12

/2εa/2εaεa

[NUREG 6909 ][NRC RG 1.207 ]

[Current data & Chopra margins ]

Figure 9. Strain life data compared to ASME 3 design curve and air curve in US NRC Reg.Guide 1.207.

The scatter in obtained fatigue lives (and endurance limit) is small, except at the secondary hardening

transition regime ( εa ≈

0,2 % ; N25≥ 105 ). At 0,2 % strain amplitude the secondary hardening begins at 25 to

50 % of N25. As notable fraction of LCF tests is usually spent in crack growth, we may assume that crackshad initiated before the hardening was effective. This may explain, why the fatigue lives are shorter at thislevel. Furthermore, three specimens of five tested at 0,2 % broke at or near the altitude of a extensometer

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knife edge. Two other specimens broke just outside of the strain measurement gauge section (at 0,4 % and0,185 %). These five specimens were marked with validity concerns, because such crack locations may leadto early acceleration of crack growth and reduce life. However, to stay in conservative side, even those datapoints are included in the geometric averages and for determining the mean curve.

The in this study determined mean curve is shown in Fig. 9 together with the mean curve modeldeveloped by Chopra (2007) for 304, 316 and other grades of stainless steels (310, 347, 348, …). Our data is

within the same scatter band in LCF regime, but scales off in HCF. A remarkable difference is seen for longlives ( N25>105). For comparison with the design curve endorsed by NRC (2007) for new designs in US, wedetermined an experimental design curve with the design margins proposed by Chopra (2007). We can see inFig. 9 that the NRC design curve would be very conservative for this particular material batch in HCF.

5 DISCUSSION ON STRAIN LIFE CURVES

5.1 Endurance limit

The obtained data in room temperature air demonstrates a typical endurance limit behaviour for this material.Note that all data below 0,195 % (at amplitudes 0,17 %; 0,185 %; 0,19 %) were run-outs against a definitionof endurance limit for 107 cycles. Two tests were interrupted and one lasted for 12,3 million cycles, before

fracturing outside of gauge section. Three tests at εa =0,195 % resulted to less than one million cycles, butthe test at εa =0,190 % was interrupted after ten million cycles. This specimen was actually tested till failure

at εa = 0,22 % and it lasted more than double the longest life received at that amplitude – clearly

demonstrating that the specimen was not near its end of life when interrupted, and also that the preloadsecondary hardening was effective even at the higher amplitude. It is naturally possible that several repeats

could bring finite lives at εa =0,190 %, but all three tests below εa =0,195 % showed remarkable secondary

hardening and we have good reason to assume that there is an effective endurance limit below, but not much

below εa =0,195 %.

The endurance limit behaviour is enhanced by secondary hardening in a synergistic manner. As thespecimen doesn’t fail, secondary hardening gets a chance to decrease the damaging inelastic strains, thuseffectively ensuring that the specimen endures even longer. Further details and discussion on the secondaryhardening phenomena go beyond this presentation, but it is worth of noting that it may occur also inoperation temperatures. The GE/EDF programme actually resulted to extreme secondary hardening in 300 ºCas convincingly reported by Solomon et al. (2004). Chopra (2007) also observed secondary hardening inelevated temperature.

5.2 Strain life curves and different stainless grades

The Langer and Chopra curves are not much different for LCF and our data lie within a common scatter bandin between. The curves deviate in HCF regime, where Chopra curve becomes more conservative.

Our current data for the stabilized steel has lower slope and extends even above the Langer curve inhigh cycle region, Fig. 8. In addition to secondary hardening, good HCF properties for this steel can be

explained by a hardening effect of Niobium carbides and – in general terms – with the classic modelscorrelating material strength and strain life curve slope.

In recent decades, mitigation of susceptibility to stress corrosion cracking has been a major challengefor material scientists in the nuclear industry. Sensitization has been prevented mainly by reducing carboncontent of the steels. As a side effect, high cycle fatigue strength of the nuclear grade stainless steels haveprobably decreased. If this is true, it might explain part of the differences between recent experimental dataand the Langer curve. On the other hand, our current results provide supporting arguments for use of stabilized stainless steels in components, where high cycle fatigue is a concern.

Together with parallel results by Solin (2006, 2009) this study clearly demonstrated that different gradesof stainless steels exhibit different fatigue performance, and that the new air curve endorsed by the NRC isdirectly applicable only to part of the stainless steels. Based on general experience, it seems probable, that

many soft grades developed particularly to exclude sensitivity to stress corrosion cracking will comply withthe Chopra curve, but many grades used in the existing plants will not. So, as the NRC also states, theRegulatory Guide 1.207 as such is applicable to new (US) designs only. Furthermore, one might consider

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determining more relevant experimental curves for material batches to be used in building of new nuclearpower plants and use code curves, when relevant experimental data is not available.

6 CONCLUSIONS

Strain controlled fatigue tests for an austenitic Niobium stabilized (X6CrNiNb1810 mod) stainless steel in

room temperature air led to the following conclusions:• The data demonstrates an endurance limit behaviour for this material. Fatigue failures were not

obtained within ten million cycles at strain amplitudes below 0,195 %.

• The endurance limit coincides with a marked transition in the cyclic stress response.

• Secondary hardening is pronounced at strain amplitudes below 0,195 %. Secondary hardeningreduces inelastic strains and thus eliminates fatigue damage.

• Determination of cyclic stress strain curve (CSSC) for this material is complicated by continuouschange of stress response through initial hardening, softening and secondary hardening. The data forCSSC should be based on largest strain amplitudes relevant in plant operation.

• The apparent elastic modulus decreases as function of the applied strain. This was consistentlyobserved after partial unloads during tensile tests and also in cyclic hysteresis loops.

Our LCF data for Niobium stabilized austenitic stainless steel extends to very long lives, where it is inbetter agreement with the Langer curve than some newly proposed air curves. This means that the originalbasis of ASME 3 and KTA design criteria is valid for this material batch, which is completely relevant forprimary loops in certain operating PWR’s. Applicability of the new air curve given in NRC RG 1.207, whichis based on different fatigue data, should be questioned also for new designs utilizing this kind of material.

Acknowledgements. The reported work is part of the Technical Programme of E.ON Case on Thermal Transients funded by E.ON Kernkraft GmbH. The experiments were carried out at VTT by Mr. J . Alhainenand E. Arilahti. The analysis of results was partly funded by the Academy of Finland decision 117700.

REFERENCES

ASME 1972. Criteria of the ASME Boiler and Pressure Vessel Code for design by analysis in sections IIIand VIII division 2. Pressure Vessels and Piping: Design and Analysis, A Decade of Progress, Volume One.P. 61-83.

Bayerlein, M., Christ, H-J ., Mughrabi, H., 1987. A critical evaluation of the incremental step test. In: Rie, K- T. (ed.), Low Cycle Fatigue and Elasto-plastic Behaviour of Materials. Essex, Elsevier. pp. 149-154.

Chopra, O., Shack, W., 2007. Effect of LWR Coolant Environments on the Fatigue Life of ReactorMaterials, Final Report. NUREG/CR-6909, ANL-06/08, Argonne National Laboratory. 118 p.

Faidy, C., 2008. Status of French fatigue analysis procedure. Proceedings of PVP 2008-ICPVT11 2008ASME Pressure Vessel and Piping Division Conference, July 27-31, 2008, Chigago, USA. Paper PVP2008-61911. 9 p.

Higuchi, M., 2004. J apanese program overview. 3rd International Conf. on Fatigue of Reactor Components,3-6. 10. 2004, Seville, Spain, EPRI/OECD, 15 p.

Jaske, C.E., O’Donnell, W.J ., 1977. Fatigue Design Criteria for Pressure Vessel Alloys, Trans. ASME J.Pressure Vessel Technol. 99, pp. 584-592.

Kuhlmann-Wilsdorf, D., Laird, C., 1979. Dislocation behavior in fatigue II: friction stress and back stress asinferred from an analysis of hysteresis loops. Materials Science and Engineering 37, ss. 111-120.

NRC, 2007. U.S. Nuclear Regulatory Commission Regulatory Guide 1.207, 2007. ”Guidelines for evaluatingfatigue analyses incorporating the life reduction of metal components due to the effects of the light-waterreactor environment for new reactors”. 7 p.

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Solin, J ., 1986. A Spectrum straining method for low-cycle fatigue resistance determination (C 245/86).Fatigue of Engineering Materials and Structures, IMechE, 1986, vol 1, pp. 273-280.

Solin, J ., 1989. Cyclic stress strain behavior of an austenitic stainless steel Polarit 778 under variableamplitude loading. VTT Research reports 647, Espoo 1989, 42 p.

Solin, J., 2006. Fatigue of stabilized SS and 316 NG alloy in PWR environment. Proceedings of PVP 2006-ICPVT11 2006 ASME Pressure Vessel and Piping Division Conference, July 23-27, 2006, Vancouver, BC,

Canada. Paper PVP2006-ICPVT11-93833.

Solin, J., Alhainen, J ., Karlsen, W., 2009. Fatigue of primary circuit components (Fate-Safir) – cyclic behav-iour and fatigue design criteria for stainless steel. SAFIR2010 - The Finnish Research Programme onNuclear Power Plant Safety 2007-2010, Interim Seminar, 12-13. 3. 2009. 13 p.

( see http://virtual.vtt.fi/virtual/safir2010/ )

Solomon, H., DeLair, R.E., Vallee, A.J ., Amzallag, C., 2004. 3rd Int. Conf. on Fatigue of Reactor Compo-nents, 3-6. 10. 2004, Seville, EPRI/OECD, 22 p.

STUK, 2002. YVL-guide 3.5, Ensuring the strength of nuclear power plant pressure devices, issue 5.4.2002.(in Finnish, but translations exist)

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