COMMISSARIAT A L'ENERGIE ATOMIQUE

CENTRE D'ETUDES NUCLEAIRES DE SACLAY CEA-CONF —9363 Service de Documentation

F9J 191 GIF SUR YVETTE CEDEX

L5

TAVASSOLI A.A. -CEA Centre d'Etudes Nucléaires de Saclay, 91 - Gif—sur—Yvette ^FR), Dept. de Technologie

INFLUENCE Dt L'IRRADIATION SUR LES PROPRIETES DES JOINTS SOUDES

Communication présentée à : 3 , Internotionol conference on fusion reactor rcatericis

Karlsruhe (DE) 4-8 Oct 1987

COMMISSARIAT A L'ENERGIE ATOMIQUE ~~j

CstA TCCMNOLOO*

1IYIS10M M RETAlLiKIE «T »'ETUKS . K S CflMUSTIHES NUCLEAIRES _ J

THIRD INTERNATIONAL CONFERENCE ON FUSION REACTOR MATERIALS

1-8 Octobre 1987

KARLSRUHE (R-F.A.)

INFLUENCE DE L'IRRADIATION SUR LES PROPRIETES DES JOINTS SOUDES

A.A. TAVASSOLI

L'effet d'irradiation sur les propriétés mécaniques des maté­

riaux retenus pour la première paroi et la couverture des réacteurs à

fusion est passé en revue. L'accent ert mis sur l'acier inoxydable

aus té ni tique type 316 Let ses métaux d'apport et les paramètres de désigne

retenus sont semblables à ceux de Next European Torus (NET), à savoir doses

d'irradiation ^.15 dpa, température ^ 400 *C, nombre de cycles ^10^ et

temps de maintien 15 mn. L'effet d'irradiation sur d'autres matériaux y

compris l'acier inoxydable austénltique type 304 L, le métal d'apport 308 L

et les aciers ferrltiques/martensitiques (9-12 % Cr- Mo), ainsi que

d'autres conditions de service comme des températures jusqu'à 550 °C, sont

également brièvement discutés.

Les données regroupées et présentées dans cette revue sont ceux

qui sont habituellement mesurées par des essais de traction, resilience,

ténacité, fatigue, fatigue-fluage et la propagation de fissure. Dans

chaque cas, l 'effet des paramètres d'irradiation sur les changements

observés est discuté et des conclusions relatives sont tirées. L'obser­

vation la plus importante faite est le manque de données concernant l 'effet

des doses moyennes d'irradiation sur les soudures et notamment sur les

joints BE.

CENTRE D'ETUDES NUCLEAIRES DE SACLAY 91191 GlF-sui-YVETTE CEDEX

TECMKM.O0IC

MVtSlOi K ICTILUMME tT D'OMIS . KS COMUSTIRIS WCLEAIKS _ J

THIR INTERNATIOHAL CONFERENCE ON FUSION REACTOR MATERIALS

4-8 Octobre 1987 - KARLSRUHE R.F.A.)

THE INFLUENCE OF RADIATION ON THE PROPERTIES OF WELDS AND

JOINTS

A.A.TAVASSOU

Service de Recherches Métallurgiques Appliquées,

Département de Technologie, DMECN. IRDI, CEN-SACLAY

91191 GIF SUR YVETTE. FRANCE

ABSTRACT

The effect of radiation on mechanical properties of candidate structural materials for

the first wall and breeder blanket of fusion reactors is reviewed. The accent is placed on

austenitic stainless steel type 316L and its weld metals and design parameters

considered are similar to those currently specified for the Next European Torus,

namely, irradiation doses $ 1 5 dpa, temperatures $ 400°C, number of pulse cycles

*105 and hold times £15 mn. The effect of irradiation on other materials, including

austenitic stainless steel type 304L, weld metal type 308L and ferritic/martensitic

steels (9-12% Cr Mo), as well as other service conditions such as temperatures as

high as 550°C are also briefly discussed.

The data collected and presented in this review are those usually measured, before

and after irradiation, through tensile, impact toughness, fracture toughness, fatigue,

creep-fatigue and fatigue crack propagation testing. In each case the influence of

irradiation parameters on the observed changes are discussed and relative conclusions

are drawn. The most important observation made is the lack of medium dose irradiation

data on the weld metal and in particular on the electron beam welded joints.

CENTRE D'ETUDES NUCLEAIRES DE SACLAY 91191 GIF-aui-ÏVETTE CEDEX

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1 .INTRODUCTION The materials requirements for the next generation of fusion reactors have been reviewed by several

authors (1-5]. These show that the performance of the First Wall/Breedir« Blanket (FW/BB) is amongst &e most critical factors determining reliability and economics of a fusion reactor. Imleed the materials selected for first wall application most retain a high degree of dimensional stability and integrity under the onerous service conditions [6,7]. These conditions include:

a) Cyclic thermal and mechanical stresses induced by high heat loads and pulsed reactor operation, coolant pressure, magnetic loads and temperature gradient.

b) Irradiation with energetic (14 MeV) neutrons, producing high number of displaced atoms and high helium and hydrogen concentrations, resulting in changes in the bulk properties.

c) Bombardment with ions and energetic neutral atoms, resulting in surface radiation and erosion of the first wall As a result most current experimental thermonuclear reactor designs envisage using removable protective tiles (such as graphite tiles) on the fust wail [8], see for instance figure 1. This measure reduces type (c) damage ar decreases therm: I loading on the first wall but the damage from types (a) and (b) remain significant In addition to the above requirements, the metal selected for the first wall must be readily fabricated and welded in required geometries (fig. 1) and be compatible with the coolant and tritium breeding materials.

The primary candidate for the FW/BB of most current designs is the stainless steel type 316L. The reason for this is thai this steel can be readily fabricated and welded, and that there is a long service experience with iu use in the nuclear industry. The most serious rival for 316L steel is probably a ferritic/ martensitic steel with 9-12%Cr-l%Mo [1-9]. In this range of chromium concentration, ferrioc/martensitic steels have a good corrosion and high temperature resistance [10], but most of all they offer superior resistance to swelling and lower thermal expansion coefficient than the austenitic steels [11-13].

In this paper, the existing data on the effect of radiation on mechanical properties of welded candidate materials for the FW/BB of fusion devices is reviewed. In particular the attention is focussed on the austenitic stainless steel type 3I6L and its weld metals. Throughout the paper numerous references are made to our, as yet unpublished, in-house data [14] which have been obtained over the past two decades as part of the studies carried out for evaluation of the behavior of fast breeder reactor internal structures.

The reference irradiation parameters which have been retained in this review are similar to those presently specified for the Next European Torus (NET) [8], that is:

- irradiation doses less than IS dpa,

- number of pulse cycles about 10 5 , - peak temperatures £400°C - hold times at peak strain;! less than about 15 mn.

2. MATERIALS Chemical composition of some of the materials referred to in this paper are presented in table 1. The

main microstructural differences between them is shown on a Schaaffler type diagram [15-18], figure 2. In this diagram the principal phases which are formed in the steels upon solidification are predicted in terms of steels equivalent chromium (Crg- ) and nickel (Ni„ ) contents (wt%):

C r „ - Cr + Mo + 1,5 Si + 0,5 Nb

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N i e q * N i + 3 0 C + 0 , 5 M n + 3 0 N

Notice that there is a slight difference between the diagrams used for predicting the phases in welds and in

base metils but in figure 2 this has been neglected and a unique diagram is used for the the base and weld

metals.

2-1. AUSTENITIC STAINLESS STEELS

As it can be seen in the figure 2, type 316L steel should form a fully austeuitic microstructure (fee)

whereas type 304L steel would contain some 5- ferrite (1-3%). In practice, however, a small amount of S-

ferrite (bec) is also frequently found in 316L( £1%). The chemical composition of weld metals usually used

for welding austenitic stainless steels (16-8-2,17-8-2,19-12-2,308L....) is deliberately chosen in a manner

to form some residual ferrite 3-9%) [19]. This is doe to the fact that the presence of 6- ferrite during weld

solidification reduces the risk of hot cracking [19] 1. Neverthelcs, it should be pointed out that the

presence of 5-ferrite in austenitic stainless steels, S -1% in 316L and 3-5% in its wild, may have some

undesirable effects even when the steel is used at low temperatures. For instance, upon aging at

températures between 300- 500°C, 6-ferrite would decompose and forms (o+a") phases. This reaction is

usually accompanied by an increase in hardness and embrittlement of 5- islets [20,21]. Although, one may

argue Aai the effect of small number of hard d-ferrite islets on the materials overall properties, in particular

the base metal properties, is negligible, some researchers have looked for compositions, such as 15-15 [22],

with fully austenitic microstructure.

Depending on the additional elements (Mo, Ti, V, etc.) added to austenitic reels other phases such as

carbides are formed in these steels which have important bearing on base and weld metal behaviors (see for

instance [23,24]). With regards to steel type 316L, two minor alloying elements play an important role.

One is nitrogen and the other is boron. Nitrogen is added «o 316L steel in order to compensate the loss of

strength provoked by reduction in carbon concentration (316L as compared to 316). Type 316L steel

containing N is called 316LN or as in the case of material specified for Superphcnix 316L-SPH. Boron h

added to steel type 316L to improve creep properties and fabricability. However it has two negative effects

one is on the weldability and the other is on the susceptibility to helium embrittlement through B 1 0 (n,a)

reaction with thermal neutrons [25]. As a result for most welded structures subjected to irradiation its

concentration is specified to less than 20 ppm.

2-2. FERRITIC/MARTENSITIC STEELS

Common ferritic/martensitic steels have in genera! less than about 13% chromium2, as in this range

they can be heat treated in the austenitic domain [26]. They also contain usually less than about 2%

molybdenum, as well as sometimes, other carbide forming elements such as tungsten, vanadium and

niobium in order to increase the creep resistance and tensile strengths.(table 1). Upon cooling they form

single or multiple microstructures which include ferritic, batnitic, inartensitic and tempered martensitic

1 In electron beam welded joints particular attention must be given therefore to the risk of hot

cracking. 2 An equation frequently used is (with composition in wt.%):

Net Cr^ » Cr + 6Si + 4Mo + 1.5W + 11V • 5Nb + 12Al(solubI<!) + 8Ti - 40C - 2Mn - 4Ni - 2Co -

30N-Cu

3

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[27]. In the case of alloys shown in table 1. MANET and rTO should form martensitic microstructures and EM12 duplex ferribc-mariensioc mkrostructure. see also figure 1.

Presently, martensitic steels with 9-12% chromium, are the primary feniuc/martensitic candidate materials for the fusion reactors (9J, since in this range they offer superior high temperature mechanical and corrosion i >istsnce properties than the lower chromium grades. Formation of 6-ferrite is considered for some steels desirable (EM12) yet for others undesirable (MANET). In any case avoiding formation of some S-ferrite specially on commercial scale would require stringent precautions.

Keeping the carbon concentration in the martensitic steels below about 0,2% allows the steel to be welded, but the steel must be preheated before welding and post-weld heat treated after welding to temper the microstructure. For instance, as reported by Lechtenerg [S], Sandvik recommends that the HT9 weld joint be preheated to 200-400°C in order to slow the weld cooling rate, reduce thermal shock to the weld region, and to drive any moisture and hydrocarbons absorbed on the metal surface. In the case of multipass welds an interpass temperature of 250°C is sometimes specified to prevent the transformation to untempered martensite which occurs upon cooling to room temperature. Finally, a post-weld heat treatment in the temperature range of 740°C-780°C for 1/2 to 2 hours is required immediately after welding. Parts must be handled carefully between the welding and tempering operation si. ce the regions of hard untempered martensite are present in the fusion zone and a port»- tftheHAZ.

Theoretically speaking it is advantageous to use austenitic electrodes or nickel based alloys to weld manensiDC steels since the risk of hydrogen diffusion from weld metal to HAZ is reduced and reduced proof stress of the filler metal results in lower residual stress. However, in the context of FW/BB where thermal fatigue is of a major concern, these advantages are overshadowed by the large differences in the coefficient of

4 thermal expansion. Furthermore Sanderson [26] sites limited testing experience which indicate that the fully relieved joints are mechanically stronger than the parent material which invariably fails in the region of tempering overlap whfch results in a heavily tempered microstructure; he adds that this is also the case for brazed joints where the braze metal flow temperature is 1200°C giving similar microstructural graduation to that from the weld heat affected zone.

3. RADIATION EFFECTS, RESULTS and DISCUSSION Radiation effects presented in this review are essentially induced by: 1- atomic displacements 2- nuclear transmutations, in particular (n,ct) [28-30].

At low and medium dose levels (<15 dpa) the risk of swelling is small [9-11,31]. As a result swelling is not discussed in this review, although Sawai and co-workers [32] have reported more void swelling in the electron beam melted specimen of type 316 stainless steel and its heat affected zone when irradiated in microscope,

3.1. ON TENSILE PROPERTIES The effect of low dose neutron irradiation (<2 dpa) on mechanical properties of austenitic stainless

steels and their weldments has received considerable attention during the past few years. This has been mainly due to their utilization in the fixed internal structures of fast breeder reactors (14,28,33-34). The effect of medium dose neutron irradiation has been receiving more attention recently, although sonw results have been available from studies on cladding and hexagonal tubing [see for instance 35-41],

4

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In general it has been observed that at above a certain "incubation" dose, austenibc steels and their weMments when irradiated at temperatures below about 50tTC. exhMt a rise in yield sness and in ultimate tensile strength widt an associated decrease in total and uniform elongations (Figs 3a and 3b). The loss of ductility is attributed to the matrix hardening and to the fact that die UTS is increased by a lower amotuit than die yield stress. This implies tint die ability of steel to work harden is reduced and hence die onset of plastic instability is advanced. At low and medium temperatures, however, some damage from irradiation induced segregation should also be taken into consideration [42].

The dose at above which tensile properties are modified is temperature dependent [14]. For austenitic stainless steels irradiated at temperatures between 300-500*0 it is situated between 0.S to 1 dpa. A saturation in hardening and embrittlement is observed at about 20-30 dpa (figures 3a and 3b ) while Ehrikh [22] has reported even a subsequent slight softening at lower temperatures (230°Q.

In general weld metal property changes are less significant don die base metaL This is due to die fact Uiat weld metal has before irradiation a higher yield strengtit and a lower ductility. Nevertheless as shown in figure 3b die weld metal ductility attains more critical levels at much lower doses.

Apart from die above major changes, there are several lesser known effects of irradiation on tensile properties which merit particular attention. For instance:

a) die hardening caused by low do»; neutron irradiation is reduced or eliminated at higher irradiation temperatures (> 55v Q or following a short post irradiation exposure to temperatures as low as 330°C [14].

o) when the test temperature is increased to above about 600°C a tendency towards intergranular rupture of irradiated specimens is observed, which provokes a reduction in ductility,

c) this latter phenomenon is not eliminated by simple aging at 550°C and appears to be related to helium embrittlement; in fact it occurs at very low helium levels (Si appm), figure 4 [14].

There are fewer published post-irradiation tensile data available on die manensitic steels and even fewer on its weldments. The available results however show similar effects as those presented hereabove for die austenitic steels [9]. For instance, Lechtenberg [9] has compared tensile property changes occurring in HT9 after exposure to 9-13 dpa at temperatures ranging from 50-550°C. He shows that die ferritic steels exhibit an increase in tensile strength when irradiated at temperatures below approximately 500°C.

3.2. ON IMPACT TOUGHNESS ~ The changes in yield stress and ultimate tensile strength discussed earlier have ramifications to other

properties such as toughness and in particular to die ductile to brittle transition temperature (DBTT). Although diere is no quantitative method of using DBTT ( or FATT ) in design, impact toughness testing is a simple and fast technique. Some attempts have been made to draw a correlation between impact toughness and fracture toughness results[44,4S], but this has been often for ferritic/martensitic steels and its extension to austenitic steels is less evident This is due to die fact that die austenitic steels do not exhibit a DBTT and their measured impact toughness values (CV) remain almost unchanged between 20 to 550°C while as it will be shown later an fracture toughness values ()\Q) decrease widi increasing temperature.

The effect of low dose neutron irradiation on die impact toughness of austenitic steels base metal is more marked than on their weld metals [14,46], This is believed to be, once again, due to die fact that die weld metal has much less to loose: CV (weld) - 1/3 of CV (base). Furthermore, weld metal properties remain lower dian diose of die base metal after irradiation to low and medium doses.

The effect of fluence and irradiation temperature on DBTT temperature of ferritic/martensitic steels has •

5

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been reviewed by several investigators, see >or instance (9). It has been observed that the DBTT of these

steels increases more significantly at temperatures below about 400*C than at higher temperatures. The

increase in DBTT usually begins to stow down after about IS dpa and tends to saturate at 20-30 dp*.

Although recent developments have considerably reduced sensitivity c. manensttic steels » irradiation,

some available data show that the DBTT of these steels may increase to temperatures as high as

150°C.rb»lowing exposure to 15 dpa at 390°C

3.3. ON FRACTURE TOUGHNESS

The reported information on die effect of radiation on facture toughness of austenftfcsteeb and ferrioc

steels and in particular their welds are very few [933.47-54]. Furthermore different steels (316,316L, 316

H, 304.304L. 321.348,347...) and testing methods or specimen sizes have been used which sometimes

render their comparison more difficult. This is particularly true for the base metal which has a high ductility

and as a result needs large size specimens (lack of space in die irradiation rig) to obtain valid test results.

Our in-house data [14], obtained with compact tension (CT20) specimens, show that there is a

reduction in fracture toughness of type 304L after exposure to doses less than about 1 dpa while there may

be a slight improvement in that of the weld metal tested at 400*C (figures 4a and 4b) 3. However as it can

be sen the observed changes are small and the base metal's toughness remains well above that of the weld

metal. In contrast fracture toughness of both weld and base metals significantly decline with increasing test

temperature [1433,52].

Michel and Gray [51], using compact tension specimens, have tested several alloys at 427°C including

steels type 304L, 316L (annealed and 20% cold worked) and 308L weld metal at higher doses (8.2 to 12

dpa). Their observations are similar to those mentioned here above, but they show that after 11 dpa weld

fracture toughness can decrease to as low as Jjç=30 KJ/nr ( figs. 4a and 4b). An important point shown

by these figures is that the weld metal exhibits a low tearing modulus after irradiation (slope of the curves

during stable crack propagation).

There is a sharp reduction in die fracture toughness of unirradiated martensitk steels at temperatures

about and less than DBTT. Odette and co-workers [53] and Dahms and Lechtenberg [54] have investigied

more recently the plane-strain fracture toughness of the 12Cr-lMoWV steel in the lower shelf region. The

purpose of their tests has been to show whether the ferritic s'eels can be reliably operated in the cleavage

fracture region, assuming that the irradiation has no effect on the lower shelf toughness.

3.4. ON FATIGUE, FATIGUE-CREEP

Most existing creep-fatigue data suggest that the phenomenon can be divided in three domains [56]: a)

fa Jgue dominated, b) simultaneous creep-fatigue, c) creep dominated.

The fatigue dominated region covers areas where creep is negligible, that is when test temperature is

low, imposed strain and strain rate is high, or hold time is short. In this domain low dose neutron

irradiation (£2 dpa, He/dpa - 1) has very little damaging effect on the fatigue properties [14]. In fact often a

slight improvement is observed. At higher doses however the effect of irradiation is more complex. For

instance Grossbeck and Liu [58,59] h2ve reported a decrease in fatigue resistance of 20% cold worked AISI

316L irradiated and tested at 430°C following exposure to doses ranging from S to 15 dpa and He/dpa -60

appm/dpa. They have attributed this to a matrix hardening caused by radiation induced formation of

3 This may be due to relaxation of stresses in the weld durng irradiation period which lasted four years.

6

1-23

precqHUKs.

The data reported an the anstenkic weM metals mdicatt that w«U metal has a lower fatit^iesisaance> but their behavior is more or less similar to that of the base metal (fig.I). For instance Micbd (59] has reported a slight reduction in fatigue resistance of type 308 weM metal at 4 8 2 ^ ami doses above atout 10 2 1

ivtm2 E> 0,1 MeV(« 2-5 dpa).

The effect of irradiation on fatigue properties of both base and weld metids becomes more marked with increasing role of creep, that is, lower strain, lower strain rate, longer hold time and higher temperature. In domains b) and c) even small doses have significant effects mainly through a change in rupture mode from transgranolar to mtergranular rupture. With regards to FW/BB operating conditions die effect of <15 mn hold at temperatures less than 400"C should be small, in fact hold times longer than 30 mn at 550°C are required to provoke a significant change in the rnicrostructure [601.

Finally the elevated temperature low-cycle fatigue and creep properties of 9%Cr-Mo variants have been reviewed by several authors (see for instance (26]). They show in general similar effects bat as it was mentioned earlier die advantage of these steels lies in their low thermal expansion coefficients.

3.5. ON FATIGUE-CRACK GROWTH For both LMFBR and fusion reactors fatigue crack growth resistance is considered to be of critical

importance. Most of die available irradiation fatigue crack propagation data have been reported by shahinian, James, Mitchel and de Vries [61-66]. Recently deVries and Mitchel [65] have compared their experiments in order to link existing data from RNL tests on High Fluence Experimental Breeder Reactor II (EBRII) irradiated single-edge-notch-cantiliver specimens of type 316 with those from ECN tests on low fluence High Flux Reactor (HFR) irradiated compact-tension specimens (w=50 mm) of type 304 stainless steel. Their result provide a good understanding of the effect of irradiation on die base metal.

As an example figure 6 shows die effect of 0,2 and S dpa on die fatigue crack growth of 316 type steel.[56]. Also shown in this figure is an upper limit proposed in [66] for both weld and base metals irradiated up to about 12 dpa at 380°C. The conclusions which can be drawn from this figure and global results reported in die literatue are similar to those drawn for fatigue. In essence they indicate that irradiation does not significantly affect the fast loading fatigue crack propagation, whereas it does reduce the FCG resistance under slow cyclic loading, in particular at high temperatures and when hold time is introduced.

4. CONCLUSIONS The foregoing review of die influence of radiation on die properties of welds and joints for application

in the first wall and breeding blanket of fusion reactors leads to die following conclusions: 1 - mechanical properties of austenitic stainless steel type 316L, considered as die primary candidate for

these applications, are modified under die influence of radiation; these changes are observed at doses above about 0,8 dpa and continues at least up to 15 dpa.

2- mechanical properties of weld metal in most, if not in all cases, are inferior to those of die base metal and remain so after irradiation exposure,

3- die need for experimental data is more acute at medium doses and in particular on die weldments.

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[35] MX Grossbeck and PJ. Maziasz, J.Nucl.Mater.. 85 & 86 (H. B), (1979) 883. [36] C. Allbertini, A. Del Grande and M Montagnani. in "Effect of Radiation on Materials", 12th Int.

Conf. ASTM STP 870, F. A. Garner and J.S. Perrin. Eds.. Am. Soc. for Testing Materials, Philadelphia, 1985.783.

[37] JP. Johnson, F. A. Gamier, H. R. Hragger and R. L. Fish, HEDL-S A-2045 A (1980). [38] L. Schafer, KFK-Report, 3029 (1980). [39] E.E. Bloom, J. Nucl. Mat 85486 (1979) 795. [40] J. Dupouy, J. Erler and R. Huittey, in "Radiation Effects in Breeder Reactor Structural Materials",

AIME, (1977) 83. [41] R. L. Klueh, M. L. Grossbeck, in "Effect of Radiation on Materials", 12th Int Conf. ASTM STP

870, F. A. Gamer and J.S. Perrin, Eds., Am. Soc. for Testing Materials, Philadelphia, (1985) 768. [42] See articles in "Radiation-Induced Sensitisation of Stainless Steels", Pro. Of the Symp. Held at

Berkeley Nuclei Lab.CEGB, on 23 Sept. 1986, Ed. by D.LR. Noms. CEGB 1987. [43] J. Vitek and R. Klueh, ADIP, Sem. Progress Report 1984, DOE/ER-0045/12 (1984) 110. [44] J.M. Barson and S.T. Rolfe, "Correlation Between KJC and Charpy V Notch Test results in the

Transition Temperature Range", ASTM, STP 466,1970. [45] R.H. Sailors and H.T. Corten, "Relationship Between Material Fracture Toughness Using Fracture

Mechanics and Transition Temperature Tests", ASTM, STP 514,1972. [46] J.R. Hawthorne and H.E. Watson, Welding Research, June (1973) 255s.Res US 15 dpa [47] M. Hamilton, F. A. Gamer and WJ.S. Yang, Fusion Technology, 10(1986)405. [48] E. A. Little, in "Effect of Radiation on Materials', 12th Int Conf. ASTM STP 870, F. A. Gamer

and J.S. Perrin, Eds., Am. Soc. for Testing Materials, Philadelphia, (1985) 563. [49] F. M. Hagg&g, W. L. Server, W. G. Reuter and J.M. Beeston, in "Effect of Radiation on

Materials",12th Int Conf. ASTM STP 870, F. A. Garner and J.S. Perrin, Eds., Am. Soc. for Testing Materials, Philadelphia, 1985,548.

[50] E. A. Little, J.Nucl.Mater.139 (1986) 261. [51] D. J. Michel and R.A. Gray, J.Nucl.Maier. 148 (1987) 194, [52] F. H. Huang, Int J. Fract, 25 (1984) 181. [53] G. Odette, G. Lucas, R. Maui, and J. Sheckherd, DAFS Quarterly Report, (1984). [54] C. Dahms and T. Lechtenberg, ADD» sem. Progress Report DOE/ER0045/14 (1984) 21. [551 M. I. de Vries, in "Effects of radddiation on Materials", Scottsdale Ar. 23-28 June 1982, ASTM

STP 782, H.R. Brager and J.S. Perrin, Eds., ASTM 1982. [56] D.S. Wood, "Fatigue Behavior of Type 316 Steel Relevant to Fast Reactors", AIEA Specialist

9

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meeting on Effect of Environment Other Than Radiation. Budapest. 198 [57] W. Sceibe. R. Schmitt and H.U. Borgstedt, KFk-Nachrichten 15.2/83 (1983) 143. [58] M.L. Grossebeck and K.C Liu, J. Nucl. Mater. 103-104 (1981) 853 [59] M.L. Grossebeck and K.C. Liu. Nucl. Tech. 58 (1982) 53. [60] A.A. Tavassoli. Phil. Mag.A.54.4 (1986) 52!. [61] P. Shahinian, H.E. Watson and H.H. Smith, in "Effect of Radiation on Substructure and

Mechanical Properties of Metals and Alloys", ASTM STP 529, American Society for Testing and Materials, (1972) 493.

[62] L.A. James. "Fatigue Crack Propagation in Ausstenitic Stainless Steels", Atomic Energie Review, 14.1 (1976) 37.

[63] DJ. Michel, HE. Watson and H.H. Smith, in 'Structural Materials for Service at Elevated Temperatures in Nuclear Power Generation", A.O. Schaefer, Ed. MPC-1, ASME, New York, (1975)167.

[64] M.I. de Vries, in "Effects of Radiation on Materials", ASTM STP 782, American Society for Testing and Materials, (1982) 720.

[65] M.I. de Vries and DJ. Michel, in "Effect of Radiation on Materials", 12th Int. Conf. ASTM STP 870, F. A. Gamer and J.S. Perrin, Eds., Am. Soc. for Testing Materials, Philadelphia, 1985.

[66] G.I. Lloyd and C.G. Chipperficld, "Proposed Crack Growth and Toughness Values for 316 Steel and Weldments at Temperatures up to 430CC at IAEA/IWGFR Specialist Meeting on "Demontration of Structural Integrity Under Normal and Faulted Conditions", Chester, England, 3-5, June 1980.

10

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TABLE 1. NOMINAL CHEMICAL COMPOSITIONS OF MATERIALS

MAJOR ELEMENTS (wt,ft) STEEL C SI MN NI CR MO N, V Nb W

316L-SPH <.030 <.50 1.8 12.2 17.5 2.30 .07 -304L <.030 <1.0 <2.0 10.5 19.0 - -321 <.08 <1.0 <2.0 10.5 18 - -308L <.04 <1.0 2.5 10.0 19.0 .5 -19-12-2 .050 .55 1.5 11.5 18.5 2.05 - -16-8-2 .0375 <.50 2.15 8.0 16.0 2.0 -17-8-2 .080 <.8 1.5 8.75 17.5 2.0 -MANET .11 .25 1.75 .60 10.4 .575 - .20 HT9 .20 .30 .55 .55 11.75 1.0 .30 EM12 .15 .425 1.13 <.30 9.50 2.0 .30

In all cases restrictions are placed on S arid P and often on Co and Cu concentrations. Maximum boron content in 316L-

SPH is 0.020% and MANET has about 0.025% zirconium.

11

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PLASMA SIDE

STEEL STRUCTURE

COOLING PIPES

Fig 1. Schematic representation of cross section of a first-wall desig'.i with probable locations of welds.

12

I -

8 12 16 20 24 28 32

Cr EQUIVALENT (Cr*Mo+1.5Si+0.5Nb)

Fig. 2. Approximate positions of austenitic (base and weld metals) and ferritic steels on a Schaeffler type diagram.

800

600 -

0) 400

200]

- •

10

TYPE 316 STEEL IHRAMDTESTAT-400-C

B HAWTHORNE • EHRJCH • TAVASSOLI • AVERAGE

Ju 20

DOSE (dp.)

30 40

Fig. 3a. Evolution of yield stress with irradiation dose for type 316 steel.

ui

2 3

TYPE 3H STEEL IBR AND TEST A.T-4WC

a HAWTHORNE • EHRUCH • TAVASSOLI • AVERAGE

20 30

DOSE (dps)

Fig.3b. Evolution of uniform elongation with irradiation dose for type 316 steel.

I

1200

1000

800

^ 600

400

200

304 TYPE STEEL IRR AND TEST AT 400-427°C

• B • B

B

• i

B a B

• TAVASSOLI

• • • • « •

a -B • "

* * • •

• B • B

B UMRR • 1dpa

# • • • i l MICHEL AND Gf I .-, - -.-

H 9.5dpa # • • • i l MICHEL AND Gf I .-, - -.-

WY

2 4 Aa (mm)

Fig.4a. Single specimen J-R curves for irradiated 304 type steel showing a significant reduction in JJC after exposure to 9.5 dpa (from 200-300 KJ/m2 to « 100 KJ/rn2).

I - 23

300 j , 308 STAINLESS STEEL WELDMEMT

lRRATESTAT40fM27'C

200

100

B UNFR • 1dpa • 11.1dpa

TAVASSOU

e ti MICHEL AND GRAY _ J I

A* (mm)

Fig.4b. Single specimen J-R curves for irradiated weld metal type 308 showing less

significant reduction in J jç Nit attaining tow values (* 30 KJ/nr after 11.1 dpa), after

[14,51].

16

I-

10

IU <

TYPE 304 STEEL AND WELD -0.5 dpm. TESTED AT 3SQ-C

B UNRR-304L • UMRR-WELD • IRR-304L

CYCLES TO RUPTURE

Fig.5. Fatigue resistance of type 304 steel and its weld after low dose irradiation [SS], is similar to that of

the type 316L [ 14] zt low temperatures, see text for the effect of higher doses.

I - 23

10 100

STRESS INTENSITY FACTOR RANGE.AK (MPaVm)

Fig.6. Fatigue Crack Growth Rate of type 316L [65] irradiated at about 550°C and tested at different temperatures. Shown also is the upper limit proposed in [66] for weld and base metal irradiated and tested at low temperatures (380°C).

18