Assessment of austenitic stainless steels

Assessment of austenitic stainless steels

Fusion Engineering and Design Fusion Engineering and Design 29 ( 1995) 371-390 Assessment of austenitic stainless steels A.A. Tavassoli SRMAIDECM...

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Fusion Engineering and Design Fusion

Engineering

and Design

29

( 1995) 371-390

Assessment of austenitic stainless steels A.A. Tavassoli SRMAIDECM,

CE-Saclay,

91191

Gif sur Yceite Cede.\-, Frunce

Abstract An assessment of stainless steels is performed in order to recommend the most suitable structural material for ITER, to identify areas where additional research and development work is needed for full qualification and licensing of the recommended material, and to estimate the cost of such a research and development programme to be completed within a five year time schedule. After a brief assessment of different stainless steel categories, the choice is narrowed down to austenitic stainless steels, and more precisely to the type 316LN composition retained in Europe for the fusion reactors, and used in four generations of fast breeder reactors in France. Material specifications for the recommended steel are presented, incorporating the most recent research and development results, obtained from work performed in Europe, Japan, and the USA. The available database is presented for the selected steel using the boundary conditions extracted from the most recent report by the Joint Central Team on ITER Outline Design, but also incorporating possible changes that may be made in these conditions either for the initial basic performance phase or for the future extended performance phase. For each evaluated property, whether physical, mechanical or environmental, reference values and their corresponding trend curves are presented and, when applicable, lower bound or upper bound design curves are proposed. Using end-of-life properties of unirradiated materials as the reference design critierion, effects of irradiation are evaluated. If the observed changes remain within the safety margins incorporated in design they are considered acceptable. This is shown to be the case for materials irradiated to doses up to 10 dpa at temperatures either less than 200 “C or around 400 “C. If the observed changes exceed or appear to exceed the safety margins incorporated in design they are considered inadequate. In this case additional experiments, and when applicable safety margins, are recommended. Areas where clear conclusions cannot be drawn, or where there is a lack of experimental data, are identified and a five year research and development programme is proposed for the establishment of a comprehensive database allowing full qualification and licensing of the recommended steel. The proposed research and development programme includes plans to use existing fast breeder reactors as a low temperature irradiation means to obtain doses greater than the 10 dpa.

1. Introduction The current reference design for the International Thermonuclear Experimental Reactor (ITER) is a water-cooled non-breeding concept and the recommended 0920-3796/95/$09.50 c 1995 Elsevier SSDI 0920-3796(94)00380-7

Science S.A. All rights reserved

structural material is stainless steel. However, stainless steel exists in various types and for a single type there are many variants. In order to reduce cost, to avoid duplication and to harmonize the research and development programme to be performed by the ITER part-

312

A.A.

Tavassoli 1 Fusion Engineering

nets, an ITER task the aim of which is a prior assessment of the austenitic stainless steels has been undertaken. The work presented here is extracted from the European contribution to this task and the reader is referred to that report (ITER Task BL-URD 3) for more details. The boundary conditions considered are those given in the ITER Joint Central Team document entitled “ITER Outline Design Report”, dated December 1993. The upper bound of these (i.e. the first wall operating conditions) with some approximations, are as follows: fusion power (nominal) surface heat flux burn time metal temperature number of cycles, up to peak end of life neutron damage peak end of life He production water temperature internal pressure

1.5 GW 50 Wcm-’ 1000 s < 300 “C 105 27.1 dpa 432 atppm < 200 “C 1.7-2.2 MPa

The operating conditions for the vacuum significantly less severe and in particular, damges are as follows: peak end of life neutron damage peak end of life He production

vessel are irradiation

0.047 dpa 0.3 at.ppm (value calculated for Inconel 625; for steel the concentration is 0.1 atppm)

However, since the ITER design options for the basic performance phase (BPP) and the extended performance phase (EPP) are subject to modification, the peak metal temperature is arbitrarily extended in our assessment to 400 “C or slightly higher, but not exceeding 450 “C. These two temperatures have particular significances for the stainless steels. At temperatures below 400 “C the accelerated irradiation swelling region is avoided [I]. At temperatures up to 450 “C, austenitic stainless steels operate in the negligible thermal creep region (Refs. [ 21 (see also Ref. [ 31) and [ 41) and therefore are not susceptible to the consequences of irradiationinduced grain boundary embrittlement. However, at such temperatures the water-cooled concept is unlikely to be used (excessive water pressure) and a liquid metal cooled concept is a most likely choice (as compared with a gas-cooled option). At temperatures less than 450 “C, austenitic stainless steels also remain compatible with liquid metal coolants such as lithium, l

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and Design 29 (1995) 371-390

and the amount of metal loss resulting from operating in such media is estimated to be negligible [5]. Notice that in this paper corrosion issues are not addressed in detail, and it is assumed that benchmark testing and component testing are treated elsewhere.

2. Selection of reference material Narrowing down the materials option from stainless steels to the austenitic steels is easily done on the basis of their wide and often unique use as a structural material in the nuclear industry. Likewise, it can easily be demonstrated that amongst different types of the austenitic stainless steels currently employed in the nuclear industry, i.e. types 304, 309, 321, 347, 348, and 316, the latter is the most promising. The feedback from service experience and the results obtained from more recent research and development work, including the fusion and the fast breeder reactor programmes, all favour the solution-annealed type 316LN steel, i.e. a low carbon grade where nitrogen is added to compensate the loss of strength due to a reduction in carbon content. Other advantages of type 3 16LN are as follows: . better irradiation database and service experience than other types of stainless steels used in structural components [6-IO]; less prone to irradiation assisted aqueous corrosion than type 304, or stabilized grades of the steel [ 111; less prone to delayed reheat cracking than the stabilized grades [ 121; less sensitive to irradiation hardening and irradiation embrittlement than type 304 steel [ 131; . better high temperature mechanical properties than type 304 steel [4,5]; its weld metals are less prone to irradiation embrittlement than those used for welding for instance type 304 steel (lower S-ferrite content) [ 131, e.g. see diagrams proposed in Refs. [ 14- 161 for calculating S-ferrite content. In comparison with the other grades of type 316 steel, type 316LN . is less prone to strees corrosion cracking than the conventional type 316 steel, and has higher mechanical properties than the plain type 316L steel [ 171. Probably the most important advantages of type 3 16LN are balanced compositions and characteristics backed up by years of service experience and research and development work [ 171, l

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A.A.

Tuoassoli

/ Fusion Engineering and Design 29 (1995) 371-390

313

In Japan, the research and development work performed during the past few years has led to the same conclusions and the steel proposed for the next generation of their FBRs has a similar composition [10,20]. However, even in the most recent French FBR, i.e. SuperphCnix, there are several variants of type 316LSPH steel (Superphinix specification is similar to the recommended composition for ITER) (e.g. they have a slightly different upper limits for carbon or boron content), each adapted to meet the specific component requirements (welding, corrosion, irradiation). Although the differences in these variants are minor, the steel composition recommended for ITER is that one specified for welded structures and structures subjected to neutron irradiation (low carbon and low boron content). More precisely, the recommended composition incorporates post-SuperphCnix information (European FBR, EFR 316LN [ 171; European Community, EC 316LN [21]; Japanese 316LN [IO]) (Table 1). The final heat treatment for plates should be performed in the range from 1050 “C to 1150 “C, usually for 30-60 min at 1070-I 100 “C. The steel microstructure in the as-received state is characterized by a homogeneous grain distribution with the possible presence of very few s-ferrite stringers aligned in the rolling direction. Frequently, however, larger grains are observed at the surface of the plates and smaller grains in the centre. In either case, the ASTM number for the grain size is within the range 3-6. The amount of F-ferrite in the steel, calculated with the Pryce-Andrews diagram, is expected to be no

comprehensive database, including heat to heat variations and product size, and proven good ductility and toughness in various shapes and forms. In comparison with the cold-worked state, the solutionannealed state has higher initial ductility and toughness than the cold-worked state [ 18,191, is less prone to irradiation embrittlement than the cold-worked state at low doses, and is not subject to local loss of cold work during welding. With respect to the arguments advanced in favour of using cold-worked grades, either for their higher strength or for better irradiation swelling resistance, some cold working will inevitably be introduced in the solution-annealed material during fabrication and handling, * solution-annealed grades would undergo significant cyclic hardening during the first few reactor cycles, while cold-worked grades could undergo cyclic softening, and . at temperatures envisaged for ITER, swelling is not a major cause of concern [6].

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l

3. Specification of the reference steel Progress made during the past 20 years has allowed a better definition of the type 316LN steel. This has been particularly notable in France where the same steel (3 16LN) has been used in three generations of fast breeder reactors (FBRs): Rapsodie, PhCnix. SuperphCnix. The changes made from one reactor to another have incorporated results of the ongoing research and development and the return of experience. Most of the changes have, however, consisted of a stricter control of the impurity levels and the allowable deviations from the nominal specified compositions that in turn have become possible through advances in steel manufacturing practices.

more

than

1% [6].

Room and elevated temperature mechanical properties in the as-received condition and following an embrittling treatment (100 h at 750 “C) should meet the corresponding requirements given in the RCC-MR [2]. Specifications for weld metal compositions are given in Table 2. Two types of weld metals are recommended: type 16-8-2 (or its variants), mainly for TIG welding, and type 19-I 2-2 mainly for welding with filler metal. The specified range of S-ferrite in both cases is 3%7%.

Table 1 Chemical compositions of type 316LN steel recommended for ITER Metal

C

Mn

Si

P

s

Cr

Ni

MO

Nb”

Cu

B @pm)

Co

N,

316LN

0.015 0.03

1.60 2.00

0.50

0.025

0.005 0.010

17.00 18.00

12.00 12.50

2.30 2.70

0.15

0.30

20

0.25

0.060 0.080

“Nb+Ta+Ti.

374

A.A.

Tavassoli 1 Fusion Engineering and Design 29 (1995) 371-390

Table 2 Weld metal compositions Metal

used for welding Mn

C

type 316LN

steel

P

s

Cr

Ni

MO

Nb

Cu

._-si

B (ppm)

19-12-2 Coated electrode

0.045 0.055

1.20 1.80

0.40 0.70

0.025

18.0 19.0

11.0 12.0

1.90 2.2

-

-

0.020

16-8-2 Coated

0.045 0.055

1.80 2.5

0.50

0.025

15.5 16.5

7.5 8.5

1.8 2.5

-

0.020

0.1

8.0 9.0

1.8 2.2

-

0.1

-

8.0 10.0

1.95 2.50

-

~ 0.1

-

electrode

16-8-2 TIC

0.030 0.045

1.8 2.5

0.50

0.025

0.020

15.5 17

16-8-2 Submerged

0.035 0.045

2.0 2.60

0.50

0.025

0.020

16.0 17.0

arc

The above compositions are applicable to both horizontal and vertical welding and intermediate positions. Most ITER parts are thin sections and are likely to be joined together using electron beam (EB) or TIG welding. Residual stresses developed in such joints are usually low and may not require a high amount of S-ferrite. This is particularly true for EB welds where heat input is low, the molten zone is very small and the requriements for a minimum ferrite content could be waved. Brazing products that do not require long exposures to high temperatures (above 1000 “C) are preferred.

4. Design properties 4.1. Physical properties 4. I.

I. Reference

values

The physical properties commonly used in design are Young’s modulus, E (MPa), density p (kg m-‘), the average coefficient c(, ( x 10mh Km’) of linear thermal expansion between 20 “C and the considered temperature, the instantaneous coefficient C(~( x 10e6 K-‘) of linear thermal expansion, the coefficient L (W mm’ Km’) of thermal conductivity, the average coefficient a ( x IO-’ m2 s-‘) of thermal diffusivity and the specific heat C, (J kg-’ K-‘) [2]. To save space only equations describing variations in these properties as a function of temperature Q (“C) are given below; the reader is referred to Ref. [2] or the detailed assessment report for more details. l

l

l

l

l

l

l

rrn=

Co

N2

0.25

15.815 + 0.006 074 20 - 2.0796 20 c(, = 15.756 + 0.010 7460 - 4.1745 x 20 1 = 13.651 + 0.014 348 20 a = 3.8003 + 0.001 688 30 20 C, = 465.49 + 0.210 450 + 8.6641 x 20

0.25

x IO-%* “C < Q < 1000 “C lO-%’ “C < Q < 1000 “C “C < B < 1400 “C “C < Q < 1400 “C IO-‘0” “C < e < 1400 “C

Also Poisson’s ratio v = 0.3 is used within the elastic region. In addition to the conventional physical properties used in design there are those more specific to fusion, e.g. electrical resistivity and magnetic permeability. Electrical resistivity plays an important role in a fusion device because of the presence of strong magnetic fields used to confine the plasma. Austenitic stainless steels have intermediate resistivity (pzo 2 85 uR cm), i.e. higher than metals such as copper, but less than alloys such as Inconel 625 (p20 z 130 uQ cm). Toroidal one-turn electric resistance has been calculated by the JCT using simple calculations for 50 mm thick stainless steel back plate: 28.2 uR. This calculation uses 316SS SA at 150 “C: 86.7 x 10-s R m. Magnetic permeability also plays an important role and ferromagnetic materials with strong magnetic response are not desirable. Austenitic stainless steels are paramagnetic and from this point of view present an advantage over the ferritic-martensitic steels. However, the presence of s-ferrite in base metal and in particular in the weld metal induces some ferromagnetism.

E=l94000-81.48 20 “C < 0 < 700 “C p = 7968.2 - 0.376 328 - 7.2872 x 1O-582 20 “C
4.1.2. Irradiation EfSects Irradiation is expected to have some effect on the physical properties of type 316LN steel particularly

A.A. Tuvassoli

/ Fusion Engineering und Design 29 (1995) 371-390

through generation of point defects and nuclear transmutations. These can be divided into two cases. The first covers various coefficients, electrical resistivity and magnetic permeability. The second covers density and modulus of elasticity. In the first case very few experimental data are available, but one can reasonably assume that the changes would be small and for most negligible in the dose range envisaged for ITER. For instance, small changes in the permeability could occur under irradiation. This is very likely to be due to the phase transformation (ferrite, carbide and oxide formation etc.) and the irradiation-induced variations in the chemical composition. These changes have been shown to be small in type 316 steel. Measurements performed in France on a cladding material exposed to a fission fluence of 7 x lo** n cm-* have shown the change to be of the order of 6”/0 (magnetic permeability changes from a scale 1 to 1.06). In the second case there are some experimental data available. Density changes under irradiation have been widely used as a means to evaluate the level of swelling. Fig. 1 shows an example of immersion test results (density measurements) used for calculating swelling of

6

I

solution-annealed type 316LN when irradiated at 412 “C < T,,, < 545 “C [6]. In this temperature range (fission reactors), swelling is more marked than at higher or lower temperatures [ 11. A possible shift in the maximum swelling response can be expected in the presence of larger amounts of helium (Oak Ridge National Laboratory unpublished data), but it is unlikely to reach 300 “C (ITER). As a result it can be reasonably assumed that under ITER operating conditions changes in the density and hence the amount of swelling for type 316LN steel would be negligible. Irradiation is not expected to have an effect on the physical value of the modulus (barring massive nuclear transmutations). However, it could have an effect on the apparent value of modulus if swelling is present or large quantities of helium are formed, i.e. because of a reduction in the material’s effective cross-section. 4.1.3. Recommendations Unirradiated physical property values can be used in design for ITER, at least in the initial phase of EDA and until more fusion related experimental data are available.

I

316LN

i i

and Welds 412°C < Tirr < 545°C

..

5

1

315

i Experimental i $data needec& p/ 0 j

.

shift’ in trend ..__. ;for ITER (Tirr c 400°C)

I

0 20

Irradiation

30

Fig. 1. Swelling of type 316LN and welds vs. irradiation dose [6]. With decreasing is suppressed

and can be considered

negligible

under

ITER

operating

40

50

dose, dpa conditions.

irradiation

temperature

(below 400 ‘C) swelling

A.A. Tavassoli / Fusion Engineering and Design 29 (1995) 371-390

376 4.2. Mechanical properties

ing standards, such as ASTM or AFNOR [27,28] and have strain rates of the order of (l-5) x 10m4 s-r (mainly about 4 x 10m4 s-r). The relatively large scatter observed in the tensile properties is mainly due to the fact that the data bank includes a variety of grades and product sizes and forms, and incorporates effects of long-term aging, up to 50 000 h at temperatures as high as 650 “C. When the data are limited to SPH (Superphenix) grades (similar to the recommended specification), the scatter is considerably reduced and the minimum conventional yield strength and ultimate tensile strength curves proposed in the RCC-MR, Section A3.1S.3.1, 1985 edition, are validated. Equations representing the average data bank yield and ultimate tensile strength curves vs. test temperature are given below:

Since there is no well established rule in the code [2,4] for dealing with the irradiation effects, the following procedure will be used throughout this paper. If the irradiation-induced effects do not exceed those estimated for end of life properties, then the current design practice is recommended. If they do exceed the estimated end of life properties, then additional safety margins have to be introduced and probably additional experiments performed to validate the practice. l

l

4.2. I. Tensile properties 4.2.1.1. Reference values. The database is excellent for Type 3 16LN base metal; see for instance Refs. [ 6,22-261. It is also good for the weld metal except in the temperature range of 50-500 “C where it is less abundant. An example of the data available is shown in Fig. 2 for the base metal yield and tensile strengths. The data shown are in genera1 obtained using international test-

! i

;

YS = 289.75 - 0.904 310 + 0.003 499 20’ - 1.0297 x IO-%’ + 1.7738 x 10~x84 - 1.4968 x IO-“H5 + 4.7016 x 10~‘506 (YS in megapascals,

i :

; :

! :

6 in degrees

Celsius)

316LN

i ; Data; B$se, (iyA,ding

ageein

1

$I ‘7”’

600

500

4

400

200

100

0 0

100

200

300

400

500

600

Test Temperature,

700

800

900

1000

“C

Fig. 2. YS and UTS vs. temperature for type 316LN steel. Data used are for several variants and a wide range of product forms and sizes, and include effect of aging up to 50 000 h at temperatures up to 650 “C (2454 tests). Also shown are minimum YS and UTS and S, values specified in RCC-MR.

A.A. Tavassoli / Fusion Engineering and Design 29 (1995) 371-390

UTS = 609.11 - 1.8168 + 0.007 805 70* - 1.3904 x IO-58

+ 9.4773 x 10-W

- 2.1127 x lo-‘285 (UTS

in megapascals,

0 in degrees

Celsius)

4.2.1.2. Irradiation efects. For irradiation doses up to 10 dpa, the available database is relatively good and adequately covers the effect of irradiation at temperatures up to 400 “C. Doses investigated in the European Fusion Technology programme at low irradiation temperatures (2575 “C) extend to 5 dpa [23,24], at irradiation temperatures of 220-400 “C to 10 dpa [25.26], and at higher temperatures (400-450 “C) to 40 dpa [6]. At temperatures higher than 450 “C the database for austenitic stainless steels is extensive as it benefits from the work done for wrappers and fuel claddings. Post-irradiation experiments have been performed on the specimens taken from base metal, electron beam welds, manual metal arc welds, and weld deposits. In addition some tests have been performed to determine the effect of strain rate and He:dpa ratio. Despite this,

371

it is admitted that neither at low nor at high temperatures does the database adequately represent the effect of 14 MeV neutrons and the high He:dpa ratio expected in the fusion environment. Assuming that the present database can be used for ITER, the effects of irradiation at temperatures less than 400 “C can be summarized for the base metal as follows: * significant increase in yield stress and ultimate tensile strength at low irradiation doses (below 1 dpa), saturating at above about 3 dpa; significant reduction in uniform elongation, total elongation and reduction of area at low doses, saturating at above about 3 dpa. Similar effects are observed for the weld metal but changes are less marked than those noted for the base metal because the starting weld metal strength is higher and its ductility lower. In fact, at higher doses, ductilities of base and weld metals converge. On a finer scale, the observed changes vary with the irradiation temperature. Fig. 3 shows plots of tensile yield strength vs. test temperature for three irradiation temperatures. Irradiation dose in all cases is kept constant and equal to 10 dpa (data mainly from Ref. [26]). l

1200 \3i6LN [IO

dpa

000

d z

::

F Z

600

, 0

100

200

300

400

Test Temperature, Fig. 3. Variation

in the yield stress vs. temperature

for type 316LN steel following

500

600

700

000

“C irradiation

to 10 dpa at temperatures

up to 425 ‘C.

A.A. Tavassoli 1 Fusion Engineering and Design 29 (1995) 371-390

378

14 i 316LN

._____. i......__.__.........(_. ____...........__ _

12

0

T{irr)

=

i 10dpa “E (225”c) UE (325°C) UE (400%)

T&t)

2

0 0

100

200

300

400

Test Temperature, Fig. 4. Effect of irradiation

on uniform

elongation

of type 316LN.

This figure shows that irradiation hardening is higher at temperatures of about 300 “C, than at lower temperatures (below 220 “C), both of which are significantly higher than at 400 “C and above [6]. In fact increasing the irradiation temperature to 550 “C, or exposing a material irradiated at lower temperatures to high temperatures, can result in a recovery of irradiation hardening. Associated with the irradiation hardening is irradiation embrittlement that is not necessarily recovered after exposure to high temperatures (total elongation does not increase). A discrepancy is reported [26,29], however, at low irradiation temperatures (below 400 “C) between the values of total elongation, which remain appreciable (above 5%), and the values of uniform elongation which are reduced to very low values (below 2%) even at low doses (Fig. 4). The cause of this discrepancy has been traced to the presence of an early peak in the stress-strain curve that does not necessarily represent the true ultimate tensile strength of the start of necking. In fact the material retains a non-negligible work-hardening potential after the first peak in stress, except at about 300 “C where the work-hardening capability is not evident.

Arrows

600

link points

of equal irradiation

and test temperatures.

also a primary criterion used in ITER outline design (through the 3S,, rule) for determining maximum allowable wall thickness for a given material in the first wall. The values of S,,, as specified in the RCC-MR (Table 5.1.2, Section A3.15.5.1) and validated by the present work are given below and in Table 3: S,,,=147Nmm~2

forH
and for 0 in the range

160-700 ‘C,

S,,, = l98( 1.0453 - 2.5053 x IO-% + 4.1763 x 10m6Q2 - 2.5069 x lO-9Q’) Since irradiation increases the tensile resistance, it would also increase S,,,. As a result the reference S,,, values can be used as a conservative measure. However, higher yield stress values need to be taken into account during elastoplastic analysis. Effect of irradiation on ductility is picked up later with the fracture mechanics data. 4.2.1.4. Additional work. Main areas requiring additional work are tensile properties at doses greater than IO dpa and * verification of irradiation effects on tensile ductility at around 300 “C. l

4.2.1.3. Recommendations. S, is the most important design criterion derived from tensile test results. It is

500 “C

A.A. Tavassoli

379

1Fusion Engineering and Design 29 (1995) 371-390

Table 3 S, values vs. temperature

S,

20 197

25 195

50 184

75 174

100 165

125 157

150 149

175 143

200 137

225 132

250 127

275 123

300 119

325 116

0 (“C) xl?

315 111

400 109

425 107

450 106

41s 105

500 104

525 103

550 102

515 101

600 100

625 99

650 98

615 96

700 95

0 (“C)

4.2.2. Impact properties 4.2.2.1. Reference values. Impact properties do not figure in the conventional design criteria for ductile materials, such as the austenitic stainless steels. However, acceptance tests include specifications for minimum values of 120 J cm-* in the as-received state for base metal (100 J cm-* following heat treatment at 750 “C for 100 h), and 70 J cm-* in the as-welded state for weld metal (30 J cm-* following heat treatment at 750 “C for 100 h). The database put together in the present work for several variants of type 316LN steel and weld metals [6,30], including the effects of cast-to-cast variability, product size and product form, shows that most of the base metal results are situated above 200 J cm-* and most of the weld metal results are situated above 70 J cm-*. In fact, the results obtained for base metals satisfying specifications recommended for ITER are so ductile that they frequently give absorbed energies beyond the capacity of the conventional impact testing machines (above 350 J cm-*). l

l

4.2.2.2. Irradiation effects. The database is poor for irradiation temperatures less than 400 “C and does not adequately cover doses greater than 5 dpa. Most of the higher dose results are for T,,, > 400 “C [6], where a slight reduction in the impact toughness of base and weld metals is observed with increasing irradiation dose. More recently, room temperature Charpy impact toughness tests have been performed in EU [31] on precracked specimens irradiated at 77 “C to 5 dpa. The results obtained show a reduction of about 3O”/u in toughness with weld deposits showing the lowest values. 4.2.2.3. Recommendations. Impact tests are essentially used for acceptance tests but, since they are easier to perform than the fracture mechanics tests, they could be used to complement the latter. This is particularly useful in the embrittled state where a correlation may

350 113

exist between Charpy impact test results and the fracture toughness results (e.g. J,,). As an interim measure one can use 30 J cm-* as a lower bound for irradiated materials. This limit matches the specified minimum Charpy toughness after an embrittling heat treatment of 100 h at 750 “C. Notice that because the service temperature for ITER is less than 400 “C, it is unlikely that there will be an additional embrittlement due to aging. 4.2.2.4. Additional work. Main areas requiring additional work are * impact toughness at doses greater than 10 dpa and irradiation embrittlement at around 300 “C.

l

4.2.3. Fracture

toughness

4.2.3.1. Reference values. Austenitic stainless steel type 316LN is very ductile and in general yields fracture toughness values that do not satisfy the plane strain criterion according to standards such as ASTM E 81389, even when using specimens as large as CT25 [6,19,32-351. This is particularly the case when the tests are performed at room temperature where JO,*nL values greater than 500 kJ m-* are often obtained. Exceptions to the above behaviour have been reported for products that do not meet the specifications or whose fabrication has not followed the specified routes (e.g. for air-melted products) [ 191. In contrast, weld metals exhibit lower fracture resistance (about 50% of base metal), except EB welds [6]. Unlike base metal, results obtained from tests performed on adequately sized weld specimens are often valid. Fracture toughness of base and weld metals decreases with increasing temperature at least up to 600 “C. From the extensive analysis performed on the EFR program [32,36], lower bound curves have been proposed for end of life behaviour of type 316LN and its welded joints that incorporate effects of long-term aging. (a) Equations describing lower bounds for type 3 16N plates are as follows:

380

A.A. Tavassoli / Fusion Engineering and Design 29 (1995) 371-390

test temperature 20 “C T < 370 “C 370 “C < T < 550 “C

fracture J = 237 J = 217 J = 129

The average curve for the welded joint at T < 400 “C is

toughness Aa0 334 Aao.334 Aa0.480

J = 215 Aao.s57

where J is in kilojoules per square metre and Aa is in millimetres. Because there is an intersection between the curves proposed for 20 “C and T < 370 “C at about Aa = 2 mm, the more conservative part of the curves is recommended for intermediate temperatures. For ITER design the curve at T < 370 “C is recommended and its validity is extended to 400 “C to cover extended ITER boundary conditions. In the same temeperature range the mean curve is given by J = 280 Aao.479 (b) Equations describing 316LN welded joints are 20”C
lower

bounds

for

J = 130 Aa”.s55 J = 115 Aao.665

type

The applicable Aa range for the above equations is 0.2-3.0 mm, where Aa represents the total crack extension including blunting. Corresponding values of J0,2BL for base and weld metals are respectively about 100 kJ mm2 and about 50 kJ me2. However, as stated earlier this does not necessarily mean that toughness values less than these values are unacceptable. For instance, toughness values admitted for cladding and wrapper materials are significantly lower. 4.2.3.2. Irradiation eflects. The database inadequately covers the effects of irradiation at low temperatures [6,19,31-351. In particular, very few experimental results are available for materials irradiated to high doses and tested at the intermediate temperatures (T,,, z 300 “C), where ductility is low. Some recent point to possible channel

1000

600

_.

I

. I I

Lower

___----

1

1.5

2

2_2_dp_” __ - -

2.5

Aa (mm) Fig. 5. Effect of irradiation on fracture toughness of type 316LN base metal.

3

A.A. Tavassoli / Fusion Engineering and Design 29 (1995) 371-390 500 1

381

I I ,

I I

400

I I I I I

-

T T

I

_

300

-

“E =3

_

5

-

irr

316LN Weld = 422 - 450°C =

test

-

427°C

I I I I I I I I I I I

_________I--29 dpa _____________

______w-----

_________------___________________________

0

400

0.5

1

1.5 Aa

2

2.5

3

(mm)

Fig. 6. Effect of irradiation on fracture toughness of type 316LN weld metal. fracture with increasing irradiation dose at these temperatures, but these are yet to be analysed. At irradiation and test temperatures less than 400 “C, the database is limited to doses up to a few displacements per atom. At irradiation and test temperatures near or just above 400 “C, the database includes results for doses up to 30 dpa (for instance, Figs. 5 and 6). Additional tests are in progress in Europe on specimens irradiated at lower temperatures. US results [ 191 reported for reference 316LN steel, following exposure to 3 dpa (50 at.ppm He) at temperatures of 60- 125 “C and 200-300 “C and tested at 25-250 “C, confirm the high toughness of the recommended steel. JQ values varying from 350 to 800 kJ mm’ were obtained with the lower values being obtained at 250 “C. The same work also indicates lower resistance for a cold-worked and an air-melted heat of 316 steel. The US work also includes tests on irradiated welded joints, but test temperatures do not exceed 90 “C. Reported values for irradiated weld metals are of the order of 350 kJ mm” (252 MPa m’12).

Combining all available information, and noting that fracture toughness decreases with increasing test temperature, the following preliminary observations are made. Base metal fracture toughness should remain acceptable and above the unirradiated lower bound curve for doses up to 10 dpa, but should fall below this limit at 20 dpa. * Weld metal toughness exhibits large scatter and reported fracture toughness values vary according to the quality of the weld. As a result, fracture toughness values could fall below the unirradiated lower bound curve at 5 dpa. Although the JO,,, value of the irradiated materials usually remains above the lower bound limit for doses up to 10 dpa, the J resistance curve (dJ/da) of irradiated materials is lowered and irradiated materials could give lower J values at higher crack extensions. l

l

4.2.3.3. Recommendations. Use as an interim step the lower bound curves presented for unirradiated base and

382

A.A.

Tavassoli / Fusion Engineering and Design 29 (1995) 371-390

weld metals for doses up to 10 dpa and 5 dpa respectively. These bounds are expected to cover satisfactorily effects of irradiation at temperature less than 200 “C or greater than 370 “C, but may not be sufficient for irradiation effects at intermediate temperatures. Additional experimental data are needed before a reliable recommendation can be made. Special attention has also to be paid to the quality of welds. For ITER, where most of the welds are EB welds or TIG welds with thin sections, it can reasonably be expected that the welds will be of good quality. 4.2.3.4. Additional work. Main areas requiring additional work are fracture toughess at doses greater than 10 dpa and fracture toughness at around 300 “C.

l

l

4.2.4. Fatigue

resistance

4.2.4. I. Reference

values

(a) With respect to endurance, the database for deformation-controlled fatigue properties (more relevant to ITER than load controlled) is excellent and covers a wide range of testing conditions and product forms and sizes [6,35,37-551.

An example of the available results is shown for the weld metal in Fig. 7, where the reference base metal curve is also plotted. The correlation between weld and base metals resistance to fatigue is good. In both cases lowering the test temperature slightly increases the resistance. Since the observed changes in the temperature interval 20-450 “C are small, the general practice is to use a single curve for these temperatures. However, additional data may be needed to validate this practice fully. Also thermal fatigue results obtained by cycling base metal specimens between an upper and a lower temperature, to simulate the first wall operation, have revealed excellent correlation with isothermal fatigue test results [50]. Experimental thermal fatigue data for welds are, however, lacking. Equations describing isothermal behaviour of base and weld metals at 550 “C are shown below. Data used in deriving these equations are fully validated and are for parallel-sided cylindrical specimens [9,48]. All tests were performed under fully reversible triangular cycles, at a strain rates of (2-5) x 1O-3 SK’, using the preferred axial extensometry for strain measurement. For base metal the Langer equation is AEt = 16 .63N-0.42’ + 0.26 25

10'

Base Metal (550°C)

1 0” 10’

1 o2

103

1 o4

lo5

1 o6

1 o7

Cycles to Rupture Fig. 7. Fatigue resistance of type 316LN weld metal; notice that the weld metal behaviour curve at 550 “C).

is close to that of the base metal (fitted

A.A. Tavassoli

For Weld metal the Langer

equation

is

As t = 1 1.848N-0.3708 + 0.222 R and the Basquin,

Manson

AEt =0 .8032N-0.0R468+ R

383

/ Fusion Engineering and Design 29 (1995) 371-390

and Coffin correlation

is

11.761N,042’58

These equations (from Refs. [Sl-531) can be used to describe a conservative curve for fatigue properties at 20 “C-400 “C, as the data in the 20-400 “C temperature range (ITER conditions) are not as extensive but generally lie slightly above the 550 “C data. The effects of hold time in tension and in compression have been investigated for type 316LN with hold times as long as 1400 min [26]. The results obtained show that the effect of short hold times at low temperatures, such as those encountered in ITER (1000 s at less than 400 “C), will be negligible. (b) With regard to cyclic deformation curves, solution-annealed type 316LN steel exhibits cyclic hardening. With decreasing strain range, the extent of cyclic hardening decreases but the trend, i.e. rapid consolidation during the first few cycles followed by a plateau, remains.

600

I 316L-SPH

Weld metal also exhibits cyclic hardening at high strain ranges but undergoes cyclic softening at low strain ranges. Monotonic and cyclic hardening curves at 20 “C and 550 “C are presented for type 316LN steel in Fig. 8. Here, monotonic data are taken from the first quarter cycles, and the cyclic data from the half-life stress amplitudes. At 550 “C cyclic hardening is more pronounced than at 20 “C, but because the initial strength is higher at 20 “C, the curves for cyclically hardened materials at the two temperatures are only slightly different. Thus cyclic hardening curves at 20 “C-550 “C can be approximated by a series of curves, between these extremes, or they can be expressed by an average equation. Equations describing cyclic behaviour of type 3 16LN base and weld metal at 550 “C are shown below. AE~= As, + Asp The plastic component is deduced equations: for base metal, G’. d = 456.037~‘:‘~~ PA I4

1

I

-

;

0.7

0.3

Total Strain Amplitude, E

ia

Fig. 8. Cyclic and monotonic

from the following

hardening

curves

,%

for type 316LN

at 20 “C and 550 “C.

384

A.A. Tavassoli / Fusion Engineering and Design 29 (1995) 371-390

for weld metal, c.d = 385.158&?‘944 PA The elastic component E, =

is calculated

using uar i.e.

o,/E

and finally E, = a,/E + (a,/456.037) where E, is the percentage).

natural

“‘.‘s4 4’ total

strain

amplitude

(not

4.2.4.2. Irradiation efects. The database at temperatures of interest is modest and is essentially for low irradiation doses [6,37-431. In addition, a great number of reported results do not satisfy standards set for the present assessment and cannot be compared with unirradiated data bank data. As a part of the European fusion technology programme four major studies have been performed. In the first work [35], specimens taken from base metal and TIG weld metals of type 316LN steel have been irradiated at around room temperature to 0.3 dpa. Post-irradiation tests were performed at 75, 250 and 450 “C and strain ranges of 0.75%, 1.O% and 1.5%. The results obtained do not show an effect of irradiation on endurance. In the second work [ 381, type 3 16LN steel and welds have been irradiated at 250 “C, together with US-Ref., US-PCA and Japanese PCA materials to doses up to 9.2 dpa (about 50 at.ppm He). Post-irradiation tests were performed at 250 “C. The results were obtained using miniature hour-glass specimens (gauge length, 5 mm; gauge diameter, 3.2 mm) that do not satisfy criteria set for the unirradiated database in the present work. Nevertheless, they do display minor effects of irradiation except at low strain ranges where a reduction in endurance is noted. In the third work [41], effects of neutron irradiation (3-4 dpa, 30-40 atppm He) have been investigated at 600, 700 and 800 K. In this work also, specimens are hour-glass shaped but they are considerably larger than those used in the previous study (gauge length, 21 mm; gauge diameter, 8.8 mm). Here also no significant effect of irradiation on fatigue endurance is observed under continuous cycling. In contrast, an effect of strain rate is observed that becomes more pronounced after irradiation. In the same work, using the striation spacing, measured on the fracture surfaces, it is shown that an increasing fraction of fatigue life is spent for crack initiation with decreasing strain range (90% at N, = 10’

for base metal at 700 K). More recently the work in the same laboratory has been extended to 10 dpa, and the results obtained confirm little effect of irradiation, even at this dose, on the continuous cycling fatigue properties. Finally, in the fourth work [8], the effects of irradiation doses up to 40 dpa have been investigated at temperatures of 400 “C-550 “C, with and without a tensile hold time (10 min), for base metal and several weld metals. The results obtained are fully compatible with the conditions set for the unirradiated database, but the irradiation temperatures are high, above 400 “C. Nevertheless, conclusions drawn from this work are in agreement with observations made earlier, i.e. in the absence of physical damage (e.g. cavities, etc) and when time-dependent properties are not involved, irradiation has little effect on continuous fatigue properties. This is the situation for continuous fatigue testing at conventional strain rates and low temperatures. When time-dependent properties are involved (high temperature, low strain rate, long hold time), which are not fully relevant to ITER operation, then irradiation has an effect on the fatigue properties through grain boundary embrittlement. On a finer scale the above work shows that irradiated specimens could undergo cyclic softening. This observation, together with in-pile experimental data [56], indicates that dislocations generated during fatigue interact with the irradiation defects. As a result, one can reasonably assume that if the irradiation defects are not stabilized (e.g. voids and cavities) they may be annihilated during simultaneous irradiation and fatigue cycling. An important consequence of such an interaction would be a better swelling resistance of type 3 16LN steel and hence its utilization at higher irradiation doses, a point that merits further attention. 4.2.4.3. Recommendations. RCC-MR recommendations are currently being updated (1993 edition) to incorporate additional data acquired since the 1985 edition. Recommendations for ITER are made following the same guidelines, but with additional data and when necessary adjustment to meet specific requirements of ITER. RCC-MR [2] recommended design fatigue values for 316L-SPH at 20 “C, 425 “C and 550 “C- are given in Tables 4 and 5. Equivalent strain range Act(%) is used instead of AE~(%) to represent both strain-controlled fatigue (e.g. thermal fatigue) and stress-controlled fatigue (e.g. vibrations). For small temperature ranges (thermal striping), enhancement of plasticity is negligible and the following equation is to be applied:

A.A. Tuvassoli

Table 4 RCC-MR

recommended 10

fatigue

design strain

385

/ Fusion Engineering and Design 29 (1995) 371-390

range

values for 316L-SPH

steel

N % = 20 ic % = 425 -C 8 = 550 “C

4.9 I 3.59 2.92

2 x 10 3.34 2.42 I .95

4x IO 2.38 1.72 1.40

lo2 1.57 1.17 0.971

2 x 10’ 1.19 0.920 0.783

4 x 102 0.924 0.753 0.655

10’ 0.671 0.579 0.513

2 x 103 0.533 0.475 0.427

4x 10’ 0.448 0.382 0.368

lo4 0.335 0.329 0.302

2 x 104 0.304 0.293 0.260

N H = 20 “C % = 425 ‘C H = 550 ‘C

4x104 0.261 0.253 0.228

IO5 0.227 0.215 0.199

2x 105 0.214 0.206 0.188

4 x 10s 0.204 0.196 0.179

1Oh 0.19 0.183 0.167

5x106 0.171 0.164 0.150

10’ 0.163 0.155 0.142

5 x 10’ 0.143 0.137 0.124

lox 0.139 0.133 0.121

5 x 10” 0.125 0.119 0.109

10’ 0.120 0.114 0.104

Equivalent strain range &(‘S) is not applicable for stress controlled factors 2 on Aat and 20 on N, decreasing by a coefficient (log k = -0.007 768[(log N) - l]).)

case for N > 106. (Design values are obtained using the usual k equal to 1.3-1.5 at lo9 cycles and 1 at 10 cycles

Table 5 RCC-MR

RCC-MR also proposes the following cyclic curves (see also Fig. 10):

recommended

fatigue design steel (stress controlled for N > IOh)

values

for 316L-SPH

-

106 0.190 0.183 0.167

-

Act (thermal

striping)

10’ 0.147 0.140 0.128

21+V

= 3 l_~

2x 10s 0.111 0.106 0.097

4 x 10” 0.107 0.102 0.094

a, [email protected]

Design values proposed in Table 4 are in good agreement with the experimental data on 316LN steel used in the present work (Fig. 9). This is somewhat expected: (a) the difference between the fatigue behavior of types 316, 316L and 316N is small; (b) data used in the present work are partially included in the RCC-MR data bank, and the RCC-MR recommendations are made for data obtained using parallel-sided specimens. However, it should be noted that the data used in the present work are sorted for type 316LN and include additional results generated since the source document was prepared for RCC-MR. Unlike the 550 “C results, values proposed in Table 4 for temperatures of 20 “C and 425 “C, and in particular, extrapolations of these to intermediate temperatures, i.e. 100%350 “C, cannot be verified with adequate experimental data. The latter are therefore considered as interim recommendations. Use unirradiated RCC-MR A&,-Nr design curves as an interim step for doses up to 10 dpa. Substantiate irradiated and unirradiated properties at temperatures between 20 “C and 400 “C with additional experimental data.

for

‘irn

As,=1OOAo---N % = 20 ‘C % = 425 C H = 550 “C

equation

where K = 718 and m = 0.319 (constant over the temperature range). Experimental data for weld and base metals at 550 “C are plotted in Fig. 10. Considering the usual scatter in fatigue results, there is a good agreement between the proposed equation and the experimental data for base metal. Weld metal cyclic behaviour, as expected, is different from that of the base metal and is not represented by the above equation. Notice that the proposed cyclic hardening curve may not apply to irradiated specimens, since the latter could undergo cyclic softening, depending on the prior level of the irradiation hardening or the applied strain range during fatigue testing. However, like weld metal which also undergoes cyclic softening, and which cyclic curve converges towards that of the base metal, the irradiated specimen’s curve could also present the same trend. 4.2.4.4. Additional work. Main areas requiring additional research and development are thermal fatigue test results for welded joints, isothermal experimental data at intermediate temperatures, isothermal experimental data at doses greater than 10 dpa, and in-pile fatigue tests for extended performance. l

l

l

l

4.2.5. Fatigue 4.2.5.1.

crack propagation

Reference

steel is very good

values. The database

and covers

for type 316LN a wide range of testing

A.A. Taoassoli /Fusion Engineering and Design 29 (1995) 371-390

386

I’ . .._..

I0

.I

~

10'

lo=

10'

10'

Cycles to rupture, Fig. 9. Validation of RCC-MR 2 GAEL)and 20 (N,) factors.

fatigue design curve at 550 “C with design curves derived from experimental

conditions (R ratio, frequency, temperature) as well as various product forms and sizes [6,36,57-591. From the extensive analysis performed for the EFR an upper bound for fatigue crack growth rate and a lower bound for fatigue crack growth threshold have been proposed for type 316LN and welds (Fig. 11). These bounds are valid for temperatures of 20 “C400 “C. The data measured in the Paris range are expressed by n da -_=C $ dN ( > where da/dN is in millimetres per cycle, AK in megapascals metres to the power l/2, and Tin degrees Celsius. The stress ratio R = K,,,,,/K,,, varies in the limits 0.05 < R < 0.5 and AK = Km,, - AK,,,,,. The exponent n can also be calculated for various values of R from n = 4.2 exp( - 2.45R) The proportionality factor various values of R from

N 25

C can be calculated

for

data using conventional

R - 0.05

log,, C = G-t 2.25 C’, = -4 in the 20-400 “C range (for temperatures in the range 400-550 “C, C, = -5.6 + 0.0047). In general the above equation is valid from the fatigue crack growth threshold values to 35 MPa ml/‘, although higher limits have been reported for room temperature tests. The recommended lower bound threshold for fatigue crack growth is with

AKth = 6.5 - 4.5R

The equation is valid for analytical reasons in the range 0 < R < 0.95 4.2.5.2. Irradiation e&&s. At present the database is inadequate. The results available are mainly obtained following irradiation at 400 “C and above, and the doses investigated do not fully cover ITER service conditions [5,57,58]. Nevertheless, it is reasonable to assume that, provided that the irradiation damage does not change the

A.A. Tavassoli 1 Fusion Engineering and Design 29 (1995) 371-390

Effective Fig. 10. 316L-SPH cyclic curves with experimental

deformation mechanism (e.g. channel fracture), it does not have a significant effect on the low temperature fatigue crack growth response of type 316LN steel. Examples of this are shown in Fig. 11, where irradiation doses up to 36 dpa do not apparently have a significant effect on the fatigue crack growth rate and the fatigue crack growth threshold of type 316LN and welds. Irradiated results, however, do not cover the full range of the Paris law. Under testing conditions where time-dependent properties are involved (e.g. high temperature, low frelong hold times) irradiation does have a quency, significant effect on fatigue crack growth, but these conditions are not relevant to ITER. Despite the above optimism, the intermediate temperature region (250-325 “C), where ductility is low, could present a different response and needs to be investigated. 4.2.5.3. Recommendations. Use as an interim step the recommended lower bound fatigue crack growth threshold and the upper bound fatigue crack growth rates in design. The main area requiring additional research and development work is

Total Strain data

l

Range,

obtained

387

%

at 550 “C for weld and base metals.

experimental data on specimens irradiated in the intermediate temperature range (200- 325 “C).

5. Conclusions 5.1.

Reference

values

The database for unirradiated type 316LN and weld metal is relatively good and does not require significant efforts. In the intermediate temperature range, 20 “C < T -C 400 “C, there is a need for some additional work to substantiate or fill in the missing data, cf. Section 3. The bulk of the effort, however, needs to be directed towards validation of the proposed ITER design concept. This includes thermal fatigue tests on welded joints, on Be-coated surfaces, and composite specimen testing representative of the first wall and other critical components. Welds usually constitute the weakest element of a joint. Factors affecting weld quality need to be investigated. The use of welds in highly irradiated areas should be avoided and the advantages of lower ferrite

,...L T;;

A.A. Tavassoli 1 Fusion Engineering and Design 29 (1995) 371-390

388

lo2

--+-

10’ loo

10” 1 $ 2 5

z ;

lo-* 1o-3

..qp.e ......0.0

_:

.......__..__ j j j ij/jj : : :;jj: ....... ._...: __...._: ..___ 1.... :...L..:..2.;

1 om4

)

ii;;;;

;

ji:

:

:

: : : j . . . . . . . . . . . . . . . . . . . . . ..__ j . . . . . j

:

:

;vp

;.+.;.;

:

::

_‘..“‘..“i

i

/ :j

:’

10‘”

;

z

&.;

i.”

&

. .........._...

1 OS7

..

. ..

i

:

:

i

:

is

: / +_T

,;w

i

1 o‘Q

! i ij

1

>___~._.~_.~..~

f

. .

;

..__...__.

f j j

< . . . . . . . .

i

j

. . . i

. . . . . . . .

. . . . . :...

; j ;

..____...___...___~ _..........; . . . . . .:;..,.. \

; i : j ;;

;

i

; ;

;

.....

ij:

:

__,_.._...___~

: : :, ,..___..

L””



:‘.....:

;

1 o‘0

/

___..).“.

.._:if _...____._: :__..___ f . ..__++~._.;..;..;

:

:jj

.A __., c..i .i..i..i

;

j

pa+ ; ,;,i j j i

:::I: :::

j:

,___..L

:

:I___.__..... :i..____i ::j;I I I

; i _..........._.. a;__........;

,‘&

Irr.

:..

... ..

. . . . . . .. ,.... j. . . . .~..f_.~,..~

;

d

:...

ijjj;

: : : :::i\ _..............: i

los

. : : : :,,,, ,,,,,,,,..... y-j ; j : ,‘~..‘+j . . .._. + . . . . . . . . . .. . . . . . . . . . . . . . +

::A”:!; ;

,...........,..

q .....

:

: : : : i ..........j.......j......j.

(36

dpa/448”C)

..I,

.

.

.

1 o-l0 I’0

IhO

AK (MPadm) Fig. 1 I. Upper and welds.

concentrations

bound

for fatigue

crack growth

rate and lower bound

(below SO/u)in thin section welds investi-

gated. 5.2. Irradiation

effects

The work to be performed can be divided in two parts: (1) doses up to 10 dpa and (2) doses greater than 10 dpa. For doses up to 10 dpa, a moderate effort is needed to cover missing data and to investigate effects of irradiation on toughness and fatigue properties at around 300 “C. The work should also include reweldability and irradiation assisted stress corrosion cracking problems and cover the effects of dpa: He ratios comparable to those found in the fusion environment. For doses greater than 10 dpa, the database is inadequate and needs to be established for almost all the properties. In addition to the above work, investigations should be undertaken for a better understanding of the damag-

for fatigue

crack growth

threshold

proposed

for type 3 16LN

ing mechanisms involved. This is particularly needed for validation of experimental results in two areas: effect of 14 MeV neutrons; interaction of moving dislocations with irradiation defects simultaneously created (see, for example, Ref. [ 561). Finally it is important that all four parties involved in the ITER-EDA agree to a common program, including common materials and common experimental procedures. The recommendations made here are therefore preliminary and are intended to be used as a basis for four-party discussion.

References [I] F. Garner, Irradiation performance of cladding and structural steels in liquid metal reactors, section 6, pp. 419-543, to be published.

A.A.

PI

Taoassoli

/ Fusion

Engineering

RCC-MR. Rtgles de Conception et de Construction des Materials Mecaniques des ilots nucleaires RNR. Vols. I and 11, AFCEN, 1993. of DCRC recom[31 European Fast Reactor, Compendium mendations, EFR B401 5 II 39 D, July 1992. [41 ASME Boiler and Pressure Vessel Code, Code Cases, Nuclear Components N-47-24, Class 1 Components in Elevated Temperature Service, Section III, Div. I. 1986. t51 A. Terlain. EC Structural Materials meeting, Garching, March 10~11, 1993. of 316L steel welded [61 A.A. Tavassoli, Characterization joints irradiated between 15 to 41 dpa, Effects of Radiation on Materials, ASTM STP 1125, 1991, pp. 110331121. PSM 2/7 and Mat 1 (PSM 3/7) Reps. Effect of Neutron Radiation on Mechanical Properties of Permanent Near Core Structures, ASTM STP 1046, 1990. pp. 6844699. [71 C. Picker, J. Wareing and A.A. Tavassoli, Data collection on the effect of irradiation on the mechanical properties of austenitic stainless steels and weld metals, IWGFR Specialists Meet., Influence of Low Dose Irradiation on the Design Criteria of Fixed Internals in FBRs, Gif sur Yvette, December l-3. 1993. Irradiation induced [81 M. de Vries and A.A. Tavassoli, segregation of fracture toughness of stainless steel plate and weld materials at about 700 K, IWGFR Specialists Meet., Influence of Low Dose Irradiation on the Design Criteria of Fixed Internals in FBRs, Gif sur Yvette. December I-3, 1993. 191 M.G. Horsten, M. de Vries, R. Schmitt, C. Picker and A.-A. Tavassoli. Effect of low dose neutron irradiation on creep and creeppfatigue properties of type 316LN steel and welds, IWGFR Specialists Meet., Influence of Low Dose Irradiation on the Design Criteria of Fixed Internals in FBRs. Gif sur Yvette, December l-3. 1993. 1101 K. Aoto. Y. Abe, I.E. Shibahara and Y. Wada, Effects of neutron irradiation on creep properties of FBR grade 3 16 stainless steel, IWGFR Specialists Meet., Influence of Low Dose Irradiation on the Design Criteria of Fixed Internals in FBRs, Gif sur Yvette, December l-3, 1993. Y. Wada. Y. Kawakami and K. Aoto, A statistical approach to fatigue life prediction for SUS 304, 316, and 321 austenitic stainless steels. Pressure Vessel and Piping Conf., San Diego, CA, June 288July 2. 1987. 1111 M. Kodama, J. Morisawa, K. Asano, S. Shima. and K. Nakata, Stress corrosion cracking and intergranular corrosion of austenitic stainless steels irradiated at 323 K. Poster Session 3. M004, Box 56, ICFRM-6 on Fusion Reactor Materials, Stresa, Steptember 27- October 1. 1993. 1121 C. Picker, A.A. Tavassoli. and W. Dietz, Survey about failures in European FBRs, Europe-CEI Exchange Meeting about Sodium Technology, Bensberg (Siemens), December 14-18, 1992. [I31 M. de Vries and A.A. Tavassoli, ICFRM-6 on Fusion Reactor Materials, Stresa, September 27-October 1, 1993. of [I41 L. Pryce and K.W. Andrews, Practical estimation composition balance and ferrite content in stainless steel. J. Iron Steel Inst. (August 1960) 4155417.

and Design

29 (1995)

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389

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