Irradiation Effects in a Highly Irradiated Cold Worked Stainless Steel

8/12/2019 Irradiation Effects in a Highly Irradiated Cold Worked Stainless Steel

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IRRADIATION EFFECTS IN A HIGHLY IRRADIATED COLD WORKED STAINLESS STEEL

REMOVED FROM A COMMERCIAL PWR

Joyce Conermann, Regis Shogan, Koji Fujimoto, Toshio Yonezawa, Yoichiro Yamaguchi3

Westinghouse Electric Company; 1332 Beulah Rd.; Pittsburgh, PA 15235, USAMitisubishi Heavy Industries, LTD.; 2-1-1, Shinhama, Arai-cho, Takasago-city, Hyogo Pref., 676-6868, JAPAN

3Nuclear Development Corporation; 622-12, Funaishikawa, Tokaimura, Ibaraki Pref., 319-1111, JAPAN

Keywords: IASCC, Swelling, PWR, Stainless Steel

Abstract

Mechanical and corrosion properties were measured on a cold

worked, Type 316 stainless steel tube removed from the core of aPressurized Water Reactor (PWR) after 23 years of service.

Neutron exposure levels of the material ranged from near 0 to65 dpa (~4.5x1022n/cm2, E>1 MeV) and irradiation temperaturesfrom 290 to 320C. As expected, the strength of the materialincreased and the ductility decreased with irradiation. Slow strain

rate and stressed O-ring crack initiation tests in PWR water wereused to determine the susceptibility of the material to irradiationassisted stress corrosion cracking (IASCC). The data indicate thatIASCC susceptibility may accelerate above 20 dpa and saturateafter 60 dpa. At the highest fluence, limited intergranular failure

was observed on the fracture surface of a specimen slow strainrate tested in inert gas.

Introduction

The assessment of irradiation effects in pressurized water reactor(PWR) internal materials has been limited by a lack of access to

high fluence materials for testing. While the end of extended lifefluence for reactor internal materials may exceed 150 dpa(displacement per atom), the highest fluence PWR irradiatedmaterial that has been tested to date was at the 35 dpa exposurelevel. While materials irradiated to higher fluences in test reactors

have been investigated, the data indicate that there may be asignificant difference between corrosion related irradiation effectsafter PWR versus test reactor irradiation [1,2]. At fluencesexceeding 10 dpa the stainless steels used in PWR reactorinternals construction will have already undergone significantembrittlement and may become susceptible to irradiation assisted

stress corrosion cracking (IASCC).

A type 316 stainless steel tube with fluences ranging from 0 to65 dpa became available. The material was a Type 316 cold

worked stainless steel, bottom mounted instrument (BMI, a.k.a.flux thimble) tube that was exposed in the center of a PWR corefor ~23 years. A test program to measure mechanical andcorrosion properties of an irradiated material over a wide fluence

range was designed based upon the use of this piece of material.The program included tensile testing, environmental slow strainrate tensile testing (SSRT), and environmental stressed O-ringcrack initiation testing.

Material

The program material was taken from a bottom mounted

instrument (BMI, a.k.a. flux thimble) tube that was exposed in thecenter of a PWR core for ~23 years. The tube outside diameterwas 7.65 mm and the wall thickness was 1.22 mm. The reportedchemistry of this material is shown in Table I along withspecification values for Type 316 SS. The process specificationfor this material requires ~10% to 12% cold work after the final

anneal to produce a yield strength between 480 and 620 MPa anda UTS of > 690 MPa.

Gamma spectroscopy was performed on the flux thimble, andfrom those results, a specimen removal plan was developed. The

plan included material from six locations on the thimble. Table IIshows the specimen blank cutting locations. These specimenlocations were selected to allow examination of differences inmaterial fluence and irradiation temperature combinations on the

material properties. One location outside the core was selected torepresent unirradiated material although the fluence at thislocation could have been as high as 0.01 dpa.

Experimental Procedures

Tensile Tests

Subsize, pin loaded tensile test specimens (Figure 1) weremachined from the flux thimble tubes. The design allowed three

specimens to be machined from around the circumference of athimble section. All of the specimens were machined so that theirlongitudinal axis was parallel to the longitudinal axis of the tube.

Tensile tests were performed by Studsvik Nuclear andWestinghouse and were per applicable portions of ASTMSpecification E8 and E21. Tests were conducted in air, at 21, 320and 340C at a strain rate of 5 x 10-3/min. The yield load, ultimateload, total elongation and uniform elongation were determineddirectly from the load-extension curve. The effective gauge

length was taken as the reduced section of the specimen. Theuniform elongation was calculated as the elongation to maximumload.

Corrosion Testing

Two types of corrosion tests were performed; slow strain ratetensile (SSRT) and constant load crack initiation tests (O-ringtesting). SSRT tests are used in corrosion testing to accelerate thefailure of susceptible materials to reasonable test times. Resultsare comparative between different materials or states within a

Proceedings of the 12th International Conference on

Environmental Degradation of Materials in Nuclear Power System Water Reactors Edited by T.R. Allen, P.J. King, and L. Nelson TMS (The Minerals, Metals & Materials Society), 2005

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Table I. Chemical Composition

Chemical composition (wt. %)

Material C Si Mn P S Ni Cr Mo Fe

Flux Thimble 316CW 0.045 0.43 1.70 0.026 0.01 13.3 17.4 2.69 Bal.

Nominal 316 SS


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testing. Tests were conducted at 340C. The O-ring test facilityhas four pull rods extending through the autoclave head. Each pullrod connects to a loading manifold which can accommodate up tosix specimens. Loads on the 24 specimens (maximum) are set by

varying the pull rod load or the lengths of the individualspecimens. Each pull rod is loaded with a Belleville washerspring stack which allows the set load to relax if a specimenshould fail. The load can then be reapplied to the remaining

specimens or the test stopped for specimen visual examination.Each load train contains a monitored load cell and LVDT position

transducer.

The specimen applied load is calculated using the ALGOR FiniteElement Analysis and Event Simulation program. The FEA

results were verified using strain gauged, unirradiated specimens.This configuration produces maximum tensile stresses on thespecimen inside diameter under the load application locations andon the outside diameter at 90 degrees to these locations. Thus four

areas are in high tensile stress during the test.

A typical load train of specimens is shown in Figure 2.

Figure 2. O-ring test assembly.

Results and Discussion

Tensile Test Results

The results of tensile tests of the program materials are shown inTable III. Figure 3 shows the tensile properties plotted as afunction of fluence. A representative photograph showing a pre

and post test tensile specimen can be found in Figure 4. All of thetensile specimens failed in the gauge section.

Irradiation greatly increases the strength and reduces the ductility

of the material. The yield strength approaches the ultimatestrength at high fluence which results in a reduction in theuniform elongation. There is little change in reduction of area ofthe necked region with irradiation. However, the length of thenecked region is reduced which contributes to the reduction intotal elongation. Even at the highest fluence and temperature the

material still exhibits significant (~7%) total elongation.

The trends in the effects of irradiation on the tensile properties arethe same at room and elevated temperatures. The extent of

embrittlement is greater at 320oC and 340oC than at roomtemperature. The strengthening and embrittlement effects werefound to be similar at the 28 and 64 dpa levels confirming

previous findings that the property changes mostly saturate in the

5 to 10 dpa range.

The fracture surfaces of the tensile specimens were examined byScanning Electron Microscopy (SEM). Most of the fracturesurfaces were completely ductile dimpled rupture. However,cleavage like fracture features appeared on fracture surfaces of the

high fluence material (65 dpa) tested at room temperature.

An example of this is shown in Figure 5. The cleavage likefeatures did not appear on the 65 dpa material tested at 320 or

340C. This difference in fracture mode may be due to different

Table III. Tensile Test Results

dpa

Test

Temperature

(C)

.

Yield

Strength

(MPa)

t mate

Tensile

Strength

(Mpa)

Uniform

Elongation

(%)

Total

Elongation

(%)

0 21 620 790 17.2 33.5

0 21 644 788 15.5 30.4

0 320 539 638 7.3 14.5

0 320 544 636 7.6 15.7

0 340 498 604 7.8 17.3

0 340 514 596 7.2 17

33 21 1013 1202 1.5 13.7

28 21 1003 1186 1.8 14.2

33 320 931 1001 0.5 6.7

28 320 951 1009 0.5 6.8

35 340 946 954 0.3 965 21 988 1158 2.7 13

65 21 1083 1179 1.4 14.3

65 320 916 998 0.7 6.9

65 320 957 1000 0.6 7

65 340 929 946 0.3 7.8

deformation mechanisms at room temperature and elevatedtemperatures. Martensite formation during deformation issuggested to be the reason for the cleavage fracture at roomtemperature [4].

SSRT Test Results

The primary results of the SSRT tests are the fracture surfacemorphologies summarized in Table IV. Fracture morphology was

determined using a shielded Amray Scanning ElectronMicroscope (SEM).

A SSRT test of a relatively unirradiated portion of the BMI tubefrom outside the core region served as a control test (Specimen

F-2A). This material would have had a very low neutronexposure, estimated at ~0.01 dpa and would have been exposed toa 290oC water environment for 23 years. Figure 6 shows thefracture surface of specimen F-2A. This specimen had several

brittle, transgranular cleavage crack initiation sites which initiatedbefore final ductile failure. This behavior is not expected inType 316 stainless steel and may be an artifact from irradiationdespite the low fluence or from the long thermal exposure time.

Figure 7 illustrates the fracture surface of the 17 dpa specimen,E-1A, tested in PWR conditions. This specimen failed in the

gauge section. The fracture initiated as an intergranular crack andgrew intergranularly about half way across the gauge section untila final, dimpled rupture, ductile overload fracture occurred. Therewas a short mixed fracture mode section between the two areas.

PWR water SSRT test specimens which were irradiated to 35 dpaand above fractured in the head to gauge section radiuses or in thegrip pin holes of the specimens rather than in the gauge sections.

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0

200

400

600

800

1000

1200

0 10 20 30 40 50 60 70

Fluence (DPA)

0.2%

YieldStrength(MPa)

21C

320 or 340C

0

200

400

600

800

1000

1200

1400

0 10 20 30 40 50 60 70

Fluence (DPA)

UltimateTensileStrength(MPa)

21C

320 or 340C

0

5

10

15

20

25

30

35

40

0 10 20 30 40 50 60 70

Fluence (DPA)

%E

longation

Uniform (21C)

Uniform (320 or 340C)

Total (21C)

Total (320 or 340C)

Figure 3. Program tensile data vs. fluence.

Pre-Test

Post-test

Figure 4. Pre and post test photographs of a representative

tensile specimen.

Figure 5. Fracture face of high fluence tensile specimen (65

dpa) tested at room temperature, illustrating cleavage like

areas in the dimpled rupture matrix.

These fractures were generally 100% intergranular and indicatethe extreme susceptibility of this material to IASCC. A subsequentanalysis of the specimen design using finite element analysis

showed that a small highly stressed region develops at the pinlocation early in the test due to the curvature of the specimenhead. In a tensile test the gauge section stress would eventuallyexceed the stress in this region and failure would occur in the

gauge section. In the SSRT test, however, this small stressedregion in effect becomes a crack initiation test. Because of thelonger time required to develop a similar high stress in the gaugesection, this pin hole region develops a crack first and thesubsequent stress increase and high crack growth rate results inthe specimen head failing before the gauge section approaches theyield strength.

Fukuya experienced the same specimen head failure in a similarirradiated BMI tube material and was able to modify the specimendesign and use head section radius loading to force failure into the

gauge section [5]. The SSRT failures occurred after the specimenstress exceeded the yield strength and some plastic deformationhad occurred. In Fukuyas study the specimens exhibited near butnot 100% intergranularity. In the tensile mode it would be

expected that failure would not be fully intergranular since theexpected low fracture toughness of irradiated stainless steel wouldresult in a final ductile overload failure when the crack lengthreached a critical size. The degree of intergranularity would then

be influenced by the hardness of the test machine. In the presentstudy the 100% intergranular failures reported are the result of the

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Table IV. SSRT Test Results

Specimen

Number

Fluence

(dpa)

Irradiation

Temperature

(C)

Test

Temperature

(C)

Test

Environment

%

Intergranular

%

Transgranular Failure Location

A-3B 65 320 340 inert gas 12 0 gauge section

A-3C 65 320 320 inert gas < 3 0 gauge section

E-1A 17 325 340 PWR 50 0 gauge section

B-1A 35 325 340 PWR 100 0 near radius

C-1A 61 295 340 PWR 100 0 in head

A-3A 65 320 320 PWR 100 0 in head

A-2A 65 320 340 PWR 100 0 in head

D-1A 35 290 340 PWR 75 0 in head

F-2A 0 290 340 PWR 0 13 gauge section

Figure 6. Fracture surface of Specimen F-2A, 0 dpa, 340oC, PWR showing several transgranular crack initiation sites.

Figure 7. Fracture face of specimen E-1A, 17 dpa, 340C, PWR intergranular features on the fracture face.

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Figure 8. Specimen A-3B tested in an inert gas atmosphere, exhibiting intergranular features on the fracture

face mixed with a small amount of dimpled rupture.

bending mode loading once the crack initiates at the pinhole. An

overload final fracture does not necessarily occur in bending.

SSRT tests were also performed on 65 dpa material at 320 and340C in dry, inert gas. Figure 8 shows the fracture face of one of

the high fluence specimens, A-3B, tested in inert gas at 340C.The fracture faces of these specimens showed 3% and 12%intergranularity respectively. This observation indicates thatanother mechanical, brittle failure mechanism, possibly unrelated

to the stress corrosion mechanism, is also operating in thesehighly irradiated materials.

O-ring Test Results

The time to failure of the program crack initiation specimens is

given in Table V. Crack initiation time was found to be stronglydependent on irradiation fluence and stress. Above 60 dpastresses at the yield stress caused failure in less than 12 hours.

Even at 65% of the yield stress failure occurred in 2 to 3 days. At17 dpa failures could be induced only near the yield stressalthough even here the time to failure was short (52 hrs.). Thedata indicates that at a given dpa level there may be a lower bound

for the stress at which IASCC can occur.

Representative SEM photographs of a typical o-ring specimenfracture face are shown in Figure 9. Cracks are found at all fourareas of maximum stress. All fractures were found to be nearly100% intergranular. A very thin layer of ductile overload wasfound on the surface opposite the initiation site. From the test

records it was determined that the cracks typically grewcompletely through the tube wall thickness in about 1-2 days.

Although the data is limited, a proposed plot of applied stress

versus time-to-failure for the various neutron exposures examinedin this program is shown in Figure 10. From this plot Figure 11was developed which presents the data in a format that predictstime to failure as a function of stress and fluence. In Figure 11two predictive curves are shown. In this depiction the upper curveis based on approximately the stresses and fluences for failure in

100 hours and might apply to short term stress increases after

Table V. Crack Initiation Test Results

Exposure

time

hrs

0 3183 No failure

17 162 17 1311 No failure

17 498 53 2077 No failure

17 723 77 3183 No failure

17 888 94 52 52

35 158 17 1311 No failure

35 162 17 1311 No failure35 437 46 619 619

35 444 47 638 638

35 707 75 41 41

35 723 76 104 104

35 880 93 19 19

35 911 96 24 24

61 164 18 1311 No failure

61 594 64 55 55

61 732 79 115 115

61 976 105 ~12 ~13

65 168 18 1311 No failure65 312 34 262 No failure

65 605 65 73 73

65 749 81 76 76

65 928 100 ~12 ~13

Time to

Failure, hrs

Yery high

dpa

Test Max

Stress,

MPa % of Y.S.

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Figure 9. Fracture surface of a representative O-ring specimen (340C).

0

20

40

60

80

100

120

140

0 500 1000 1500 2000 2500 3000 3500

Time to failure, hrs

%

ofyieldstres

s

0 dpa

35 dpa

65 & 61 dpa

17 dpa

Figure 10. Semi-quantitative plot of the O-ring crack

initiation data.

0

100

200

300

400

500

600700

800

900

0 20 40 60 80

Neutron exposure, dpa

Stress,

MPa

Short term

IASCC failure

possible

No IASCC

Long term

IASCC failure possible

Figure 11. Prediction of IASCC time-to-failure failure versus

stress. Failure would be predicted after a significant time

above the Long term curve and in a short time above theShort time curve.

fluence has accumulated in the material. The lower prediction

curve is taken from the lower bound (Figure 10) of stress forwhich IASCC can occur at a given fluence level. This mightrepresent a stress below which IASCC would never occur even atend of life fluences. A component operating at 600 MPa stress, for

example, would be susceptible to IASCC above 20 dpa but wouldnot have a high probability of failure unless held at stress forseveral months (from Figure 10). However, if this component

was stressed to 600 MPa after 40 dpa, failure would be expectedin a very short time. The reader should be aware that these

predictive curves are based on limited data for one heat of

material. They are intended to be illustrative of a methodologythat might be applied to crack initiation data for predicting the lifeof internals components in PWR reactors. The InternationalIASCC Advisory Committee has initiated a new testing program

to better quantify this methodology in the future. The dpa axis inFigure 11 can also be translated to a time axis for any givenlocation in the reactor using the neutron flux at that location.

Summary

1. Saturation in irradiation induced hardening and ductility lossabove 10 dpa after PWR irradiation was verified to 65 dpa.

2. Limited ductility (~7%) was retained in tensile tests even at

fluences as high as 65 dpa. The uniform elongation was veryminimal at high fluences suggesting that the notched tensilestrength and fracture toughness may be adversely affected.

3. SSRT data was limited by IASCC failure in the heads of the

specimens for material with a fluence of 35 dpa and above.Qualitatively this indicated the high susceptibility of thismaterial to IASCC and the need to consider stressconcentrations in analyzing component effects.

4. Cleavage like features appeared in the 65 dpa tensilespecimens tested at room temperature. These features may bethe result of the formation of martensite during thedeformation. It appears that a high fluence is required for

this cleavage fracture to occur.

5. At very high fluences, limited intergranular cracking wasfound in inert gas SSRT tests. This failure mechanism isdifferent than IASCC although it may have the sameunderlying causes.

6. IASCC crack initiation was induced at stresses as low as46% of the yield strength at 35 dpa. At higher stresses andfluences cracks could be initiated after a few days of

exposure.

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7. A methodology was developed for using crack initiation datafor predicting IASCC susceptibility in PWR austeniticcomponents as a function of stress, fluence and time.

8. The property changes in this program were correlated withirradiation induced microstructural changes, chemicalcomposition analysis near the grain boundaries and retainedgas analysis by Fujimoto et al. [6]. IASCC susceptibility

changes did not correlate with grain boundary chemicalchanges or mechanical property changes. A correlation was

suggested with hydrogen and/or helium gas evolution.

Acknowledgements

The authors wish to acknowledge the encouragement and fundingfor this project by the International IASCC Advisory Committee:

Electric Power and Research Institute/MaterialsReliability Project representing the US PWRutilities

The 5 Japanese PWR utilities: The Kansai Electric

Power Co., Inc., Hokkaido Electric Power Co.,Inc., Shikoku Electric Power Co., Inc., Kyushu

Electric Power Co., Inc., and The Japan AtomicPower Company

Tractebel Energy Engineering

Nordostschweizerische Kraftwerke ( NOK )Vattenfall/Ringhals supplied the test material through Studsvik

Nuclear. Studsvik Nuclear performed most of the tensile testing.Laboratory testing was carried out by the Westinghouse ElectricCo., Science and Technology Department Hot Cell staff, in

particular Greg Kustra, Bruce Lingenfelter, and Bob Rees.

References

1. R. P. Shogan and T. R. Mager, Susceptibility of Type 316Stainless Steel to Irradiation Assisted Stress CorrosionCracking in a PWR Environment, Proceeding of the 10th

International Conference on Environmental Degradation ofMaterials in Nuclear Systems-Water Reactors, NACE, 2001.

2. J. Conermann, et al., Characterization of Baffle FormerBolts Removed from Service in US PWRs, Proceedings of

the 10th International Conference on EnvironmentalDegradation of Materials in Nuclear Systems-WaterReactors, NACE, 2001.

3. J. F. Williams, et al, Microstructural Effects in AusteniticStainless Steel Materials Irradiated in a Pressurized Water

Reactor, Proceedings of the Eighth InternationalSymposium on Environmental Degradation of Materials in

Nuclear Power Systems-Water Reactors, American NuclearSociety, August, 1997.

4. A. Jenssen, and R. Jakobsson, Mechanical Properties,Hardness and Microstructure of a Flux Thimble Irradiated ina PWR, (STUDSVIK/N(K)-02/016, October 2002).

5. K.Fukuya, et al., IASCC Susceptibility and Slow Tensile

Properties of Highly-irradiated 316 Stainless Steels,Journalof Nuclear science and Technology, Vol. 41, No. 6,

p 673-681, June 2004.

6. K. Fujimoto, et al., Effect of Accelerated Irradiation andNuclear Transmuted Gas on the IASCC Characteristics forHighly Irradiated Austenitic Stainless Steels, The 12th

International Conference on Environmental Degradation of

Materials in Nuclear Systems-Water Reactors, edited byTMS (The Minerals, Metals & Materials Society), 2005.

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Session Name: IASCC I

Session Day/Time: Wednesday, 8/17/05, 8:00 am-noon

Irradiation Effects in a Highly Irradiated Cold WorkedStainless Steel Removed from a Commercial PWRPresenter: Joyce Conerman

Name of Person Asking Question: V. Ehrnstain

Affiliation of Person Asking Question: VTT, Finland

Question: In the SSRT-specimens tested in PWR-water with IG and ductile fast

final fracture, was there any other type of fracture morphology between the IGand ductile areas?

Response: No, there was no other type of fracture morphology observedbetween the IG and ductile areas.

Name of Person Asking Question: Martin Widera

Affiliation of Person Asking Question: RWE Power

Question: Concerning your tensile tests, did you also evaluate the fracture areareduction? And, if yes, what are the approximate values?

Response: The fracture area reduction was not evaluated.

Name of Person Asking Question: Raj Pathania

Affiliation of Person Asking Question: EPRI

Question: You showed an IASCC time to failure-stress-fluence plot based on yourtests. Is it consistent with the field experience with cold-worked Type 316SSbaffle bolts in PWRs?

Response: Data from field experience with col-worked Type 316SS baffle boltshas not yet been compared with the prediction model. However, experimentaldata generated from testing of irradiated 304SS baffle plates has been consistentwith the prediction model (see next comment by Massoud).

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Comment from Jean-Paul Massoud (EDF): Just a comment to confirm thatadditional experimental data (on SA 304 material) confirms the time to failurediagram.

Name of Person Asking Question: Al Strasser

Affiliation of Person Asking Question: Aquarius Services

Question: 1) What stresses was the tube exposed to in service? 2) At whattemperature were the ductility measurements made? Particularly the 30 dpa) doses. Can you speculateon the mechanism by which cracking is occurring in Ar at high dose?

Response: Intergranular cracking of high fluence stainless steel tested in inertgas has been observed by others. It is not clear by which mechanism thiscracking is occurring. However, it may be related to the high concentration ofbubbles along the grain boundaries of the high fluence specimens or perhaps ahydrogen effect.

Name of Person Asking Question: Regis Shogan

Affiliation of Person Asking Question: Westinghouse Electric

Comments made by co-author to discussion after presentation:

The stress vs. ppa to failure curves will also be a function of material variability.

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SSRT tests are valuable in low susceptible (IASCC) materials where high strainsare needed to produce failures. Constant load crack initiation tests are useful inmore susceptible materials where cracks initiate near or below the yield strength.

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Irradiation Effects in a Highly Irradiated Cold Worked Stainless Steel