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SELECTION OF MATERIALS FOR PROTOTYPE FAST BREEDERREACTOR

S.L.Mannan, S.C.Chetal, Baldev Raj and S.B.BhojeIndira Gandhi Centre for Atomic Research

ABSTRACT

This paper discusses the rationale behind the selection of materials for the differentcomponents of the 500 MWe sodium cooled Prototype Fast Breeder Reactor (PFBR).The major factors considered in the selection of materials include operating conditions,availability of design data in nuclear codes, ease of fabrication, international experienceand cost. Attempt has been made to minimise the number of materials and weldingconsumables in order to avoid mix up of materials during fabrication, and to reduce thecost of Research and Development on materials development and characterisation.From consideration of radiation damage, 20% cold worked Alloy D9 (15Cr-15Ni-Mo-Ti-Si) has been chosen for the initial core of the PFBR. Type 316L(N) stainless steel(SS) has been chosen for structural components of reactor assembly, other than the corecomponents, operating at temperatures above 700 K while 304L(N) SS is the choice forcomponents operating at lower temperatures. Modified 9Cr-1Mo steel is the choice forsteam generator while carbon steel has been chosen for top shield components of thereactor assembly. Stringent specifications for chemical compositions and othermechanical properties have been drawn for PFBR materials with the view to improvereliability of components.

1.0 INTRODUCTION

The design of Prototype Fast Breeder Reactor, a pool type 500 MWe sodium cooledreactor, has been completed and the construction of the reactor is to start soon. The coreof the reactor consists of fuel subassemblies containing (U,Pu) mixed oxide fuel, andthe subassemblies are immersed in a pool of liquid sodium. The heat transport systemconsists of primary sodium circuit, secondary sodium circuit and steam-water system.The flowsheet of the heat transport system is shown in Fig.1. Structural materialschosen for sodium circuit components must possess adequate high temperature lowcycle fatigue strength and creep strength, and should be compatible with liquid sodiumcoolant. Fuel clad and wrapper materials should be resistant to irradiation inducedswelling and embrittlement, sodium corrosion, and possess adequate end-of-life creepstrength and ductility. Steam generator materials must have sufficient high temperaturelow cycle fatigue and creep strengths, freedom from stress corrosion cracking (bothchloride and caustic environments) and resistance to sodium decarburisation. Through-thickness ductility is an important consideration in the choice of material for the topshield consisting of roof slab, large rotatable plug, small rotatable plug and control plug.Three classes of steels, namely, austenitic stainless steel, ferritic steel and carbon steelwill be used for the manufacture of different Nuclear Steam Supply System (NSSS)Components. The following sections discuss the selection of materials for coresubassembly, reactor assembly, steam generator and top shield of PFBR. The selectionof materials is based on the information available in the literature and the design codes,

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as well as the knowledge and understanding gained from R&D activities at IndiraGandhi Centre for Atomic Research (IGCAR).

Figure 1

PFBR heat transport flowsheet.

2.0 CORE SUBASSEMBLY

In a fast reactor, the active core in which most of the heat is generated by nuclearfission, consists of a large number of fuel subassemblies. Each subassembly consists ofa hexagonal wrapper tube which contains many smaller circular tubes (also called fuelpins) which are filled with the nuclear fuel pellets. A schematic of the fuel subassemblyand cut-out of fuel pins is shown in Fig.2. In PFBR, there are 181 fuel subassemblieswhich are arranged in a triangular array. Each fuel subassembly consists of 217 heliumbonded pins each of 6.6 mm outside diameter.

Successful operation of Fast Breeder Reactors (FBRs) is largely dependent onthe performance of core structural materials, i.e., clad and wrapper materials of the coresubassembly, which are subjected to intense neutron irradiation. The major criteria forselection of materials for clad and wrapper materials are given in Table 1. The neutronflux levels in FBRs are about two orders of magnitude higher (~1015 n/cm2s-1) ascompared to thermal reactors. This leads to unique materials problems like voidswelling, irradiation creep and helium embrittlement which determine the permissiblelife of fuel subassemblies in the core. Deformation of various components of thesubassemblies can occur owing to void swelling, thermal creep and irradiation creep.Swelling and irradiation creep are very sensitive to environmental variables. Differentialswelling can occur because of gradients in flux and temperature in the core. Wrapperdeformation should be limited, otherwise interaction between wrappers will lead to

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excessive loads for fuel handling. Irradiation-induced void swelling and creep canproduce different types of wrapper deformation. At the centre of the core, subassembliesare expected to remain straight with an elongation and an increase of distance acrossflats. But at the periphery, subassemblies may bow outwards owing to differential voidswelling on the opposite faces of the wrapper as a consequence of neutron flux gradient(Fig. 2). The combination of void swelling and creep induced by internal sodiumpressure produces dilation and rounding of wrapper faces. Besides, clad and wrappermaterials are required to be compatible with liquid sodium. The operating temperaturesare high and stresses of sufficient magnitude are also present so that creep strength,tensile strength and ductility are important requirements for the clad material.

Figure 2Schematic of fuel subassembly showing the cut out

of fuel pins, bulging and bowing.

For economic viability, the target burn-ups required for FBRs are large, more than 20atom % of heavy metal (200,000 MWd/t), and this can be achieved only by theavailability of materials resistant to void swelling, irradiation creep and irradiation

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embrittlement, as well as satisfying the high temperature mechanical properties. (Theuseful life time of fuel in a reactor is measured by the burnup, expressed as the totalamount of thermal energy generated per unit quantity of heavy element charged to thecore). The residence time of the fuel subassembly, and hence the achievable fuel burn-up, is limited by either the void swelling of the hexagonal sheath material or creepstrength of clad. Since fuel cycle cost is strongly linked with burn-up, selection ofmaterials resistant to void swelling and irradiation creep is very important.

2.1 Criteria for Selection of Materials2.1.1 Design Conditions

In PFBR, the fuel clad tubes experience temperatures in the range of 673-973 K understeady state operating conditions. Under transient conditions, (due to failure of pumps,rupture of pump to grid plate pipe, uncontrolled withdrawal of control rod etc.), thetemperatures can rise upto 1273 K. For target burn-up of 100,000 MWd/t, the maximumneutron dose is 85 dpa (displacements per atom, i.e., the number of times an atom isdisplaced from its lattice site). Major loads experienced by the fuel clad are the internalpressure due to accumulated fission gases released from fuel matrix (~5 MPa) andmoderate fuel-clad interaction1. The latter occurs especially during transient over power(TOP) incidents. Other loads are due to temperature gradients and irradiation inducedswelling gradients. Fretting and corrosion allowance have to be considered since cladtube wall thickness is low.

The hexagonal sheath of the core subassembly operates at relatively lowertemperatures than the fuel clad. The typical operating temperature range is 673-873 Kwhich incidentally falls within the peak swelling temperature range. During transients,the temperature may rise upto 1073 K. The peak neutron dose is about 85 dpa similar tothat for fuel clad. Major loads on the hexcan are the internal pressure due to sodiumcoolant (~0.6 MPa) and the interaction loads at the contact pads due to bowing of thesubassemblies under temperature and swelling gradients. Slight thermal stresses due totransients are also present because of the higher thickness of the wrapper. The majorissues in the choice of clad and wrapper materials are given2,3 in table 1.

ASME Pressure Vessel Code Section III, Division I, Subsection NH4 and RCC-MRcode5 (French code for fast reactors) which govern the design of out-of-core structuralcomponents for FBRs cannot be applied for the design of fuel subassemblies. Therefore,design has to be based on the rules available in the open literature and data available onin-reactor behaviour and out-of-pile properties.

2.1.2 International Experience

Table 2 gives the chemical composition of various alloys in use or under considerationfor clad and wrapper tubes in FBRs. Structural materials for fast reactor corecomponents have evolved continuously so as to improve fuel element performance. Thefirst generation materials belonged to 304 and 316 SS grades. These steels quicklyreached their limits because of unacceptable swelling at doses higher than 50 dpa.Many improvements were made by the addition of stabilising elements, by changes in

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Table 1Principal Selection Criteria for LMFBR Core Structural Materials.

Criterion Clad Tube Wrapper Tube

Irradiation effects Void swellingIrradiation creepIrradiation embrittlement

Void swellingIrradiation creepIrradiationembrittlement

Mechanicalproperties

Tensile strengthTensile ductilityCreep strengthCreep ductility

Tensile strengthTensile ductility

Corrosion Compatibility with sodiumCompatibility with fuelCompatibility with fissionproducts

Compatibility withsodium

Good workabilityInternational irradiation experienceas driver or experimental fuelsubassembly

Availability

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Table 2Chemical composition (in weight %) of various materials used world-over as in-core materials in fast breeder reactors.

Alloy Cr Ni Mo Mn Si C Ti Nb P S B N Co Fe

JPCA* 14.2 15.6 2.30 1.80 0.50 0.060 0.24 --- 0.030 balance

316Ti* 17.1 14.1 2.75 1.50 0.49 0.049 0.34 --- 0.015 balance

15-15Ti* 14.7 14.7 1.15 1.6 0.43 0.096 0.43 --- 0.007 balance

Si modified

15-15 Ti*14.9 14.8 1.46 1.5 0.95 0.085 0.5 --- 0.007 balance

PNC 1520* 15.0 20.0 2.5 1.90 0.80 0.060 0.25 0.11 0.025 balance

1.4970* 15.1 15.1 1.26 1.3 0.49 0.088 0.48 --- 0.004 balance

FV548* 16.5 11.8 1.44 1.14 0.35 0.11 --- 0.92 --- balance

ASTM A771(316 SS)

17.0-18.0

13.0-14.0

2.0-3.0

1.0-2.0

0.5-0.75

0.04-0.06

- 0.05 0.04 0.01 0.002 0.01 0.05 balance

ASTM A771(D9)

12.5-14.5

14.5-16.5

1.5-2.5

1.65-2.35

0.5-1.0

0.03-0.05

0.1-0.4

0.05 0.04 0.01 20 ppm 0.005 0.05 balance

PE16* 16.5 43.4 3.15 0.01 0.01 0.08 1.27 0.003 Al=1.20+ bal. Fe

PFBR D9 13.5-14.5

14.5-15.5

2.0-2.5

1.65-2.35

0.5-0.75

0.035-0.05

5-7.5x C

0.05 0.02 0.01 10-20ppm

0.005 0.05 balance

JPCA - Japanese Prime Candidate Alloy.* Typical composition

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the chemical composition of major and minor elements and by modifications of themetallurgical structure such as by cold working.

Peak swelling in austenitic stainless steels takes place generally in thetemperature range 673-973K during irradiation in FBRs6. The formation and growth ofvoids and consequently swelling is sensitive to nearly all the metallurgical variables likechemical composition and thermo-mechanical history, and irradiation parameters likefluence, dose rate and irradiation temperature. The fluence dependence of swelling canbe described as a long incubation dose, a transient period with low rate of swellingfollowed by an acceleration to a regime of near linear swelling rate. Improvements inswelling resistance in advanced alloys are by way of a longer incubation and transientregimes. The steady state swelling rate is constant for most austenitic SS alloys, being~1% per dpa, over a wide range of irradiation temperatures.

The trend in the development of radiation resistant 300 series austenitic stainlesssteels has been to increase nickel content and decrease chromium content in comparisonto the standard versions. Solute elements like titanium, silicon, phosphorous, niobium,boron and carbon play a dominant role in determining void swelling resistance. Thisobservation has led to the development of advanced core structural materials such as D9and D9I. Figure 3 compares the performances of a few clad materials and it clearlybrings out the increase in swelling resistance on moving from unstabilized 316 SS to316Ti, 15-15Ti (D9) and a silicon modified version of 15-15Ti steel7. The majordifference in these four alloys shown in Fig. 3 is the increase in the incubation dose forswelling. These doses are 45 dpa for cold worked 316 SS, 95 dpa for cold worked 316Ti and beyond 100 dpa for 15-15Ti and its Si modification. Cold worked 15-15Ti hasreached record dose of 140 dpa without excessive deformation. Ti/C ratio is known toplay an important role in determining irradiation behavior. Maximum swelling

Figure 3Variation with dose of the maximum diametral

deformation of fuel pins irradiated in Phenix for variouscladding materials7.

resistance in cold worked 15-15Ti of a high carbon grade (C 0.08 to 0.12 wt%) has beenobtained when Ti/C ratio is below the stoichiometric composition i.e. when the materialis understabilized (i.e., Ti content of less than four times the carbon content in weightpercentage)8. The reason for this behavior is the synergistic interrelation between freely

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migrating carbon and the formation of finely dispersed TiC particles. The fine TiCparticles are redissolved by recoil dissolution and then continuously contribute totrapping mechanism so long as the steel is understabilized.

Another class of alloys that has been studied extensively for their irradiationbehaviour is nickel base superalloys (PE16, IN706 etc.). Most of the studies have shownthat swelling is low in these alloys. But nickel base alloys suffer from serious irradiationembrittlement due to helium. 9-12% Cr ferritic-martensitic steels are considered as thelong-term solution for FBR core structural materials. Although these alloys [9Cr-1Mo(EM10), Modified 9Cr-1Mo (Gr. 91), 9Cr-2MoVNb (EM12), 12Cr-1MoVW (HT9)etc.] have excellent swelling resistance to doses even upto 200 dpa, (1% swellingreported in HT9 after irradiation at 693 K at 200 dpa), their creep resistance decreasesdrastically above 823 K. Therefore, they are not suitable for clad tubes. A high thermalcreep strength is not a primary requirement for the wrapper material since the operatingtemperatures are below or at the lower end of the creep range for these materials, andthe stresses are also low. A reduced creep strength is therefore acceptable. However, theincrease in ductile to brittle transition temperature (DBTT) due to irradiation is a causeof concern for ferritic steels. Consequently, extensive studies involving modification ofthe composition and initial heat treatments have been carried out to improve the fracturetoughness. The upper-shelf energy and shift in DBTT appear to saturate at highirradiation doses. Significant increase in toughness (i.e. low DBTT and high upper shelfenergy) have been realised by (a) avoiding the formation of delta-ferrite and ensuringfully martensitic structure in 12% Cr steels by close control of nickel and chromiumequivalent element concentration; (delta-ferrite regions exhibit greater void formationand swelling than adjacent regions) and (b) optimising the austenitising temperature torefine the prior austenite grain size, and tempering treatments to reduce the strength ofthe martensite in 9-12 % Cr steels9,10. 9Cr-1Mo grades of ferritic steels have beenreported to show the lowest increase in DBTT among the various grades of ferritic-martensitic steels11. These materials are very promising for wrapper applications and arelikely candidates for subsequent cores of PFBR. Table 3 lists the clad materials used inmajor FBRs in various countries12.

2. 2 Choice of Alloy D9 for Clad and Wrapper Tubes

Alloy D9 in 20% cold worked condition (20CW D9) has been chosen for clad andwrapper tubes for the initial core of PFBR. D9 is compatible with fuel and sodiumcoolant, and the international experience on its in-reactor behaviour is satisfactory.Chemical composition of alloy D9 is based on modifying that of 316 SS fromconsiderations of better swelling and irradiation creep behaviour. The desiredcomposition is achieved by controlled additions of silicon and titanium, increasingnickel content and lowering the chromium. Minor elements having strong neutronabsorption cross-section and impurities affecting weldability have been kept to aminimum. Small addition of boron is made to improve creep ductility. Permissibleinclusion contents are stringent so as to minimise radiation embrittlement and sodiumattack (as the cladding tube wall is extremely thin). Grain size is specified betweenASTM No. 7 to 9 for clad and 5 to 9 for wrapper tubes. Chemical composition of alloyD9 specified for PFBR is given in Table 2 along with other in-core materials usedworldwide. ASTM A771 specifications13 are also included in the Table for comparison.

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In the light of recent international alloy development efforts, and in-reactorexperience, modified D9 known as D9I, is considered for future cores of PFBR. Thebasic composition would be retained as that of D9. Phosphorous in the range of 0.025-0.04 wt%, silicon in the range of 0.7-0.9 wt%. and boron in the range of 0.004-0.006wt% are recommended.

Table 3Materials selected for cladding in major FBRs

Reactor Country Fuel clad tube material

Rapsodie France 316 SS

Phenix France 316 SS

PFR U.K. M316 SS, PE 16

JOYO Japan 316 SS

BN-600 Russia 15-15Mo-Ti-Si

Super Phenix-1 France 15-15Mo-Ti-Si

FFTF U.S.A. 316 SS & HT9

MONJU Japan mod 316 SS

SNR-300 Germany X10 Cr Ni Mo Ti B1515(1.4970)

BN-800 Russia 15-15Mo-Ti-Si

CRBR U.S.A. 316 SS

DFBR Japan Advanced austenitic SS(PNC1520)

EFR Europe PE16 or 15-15-Mo-Ti-Si

FBTR India 316 SS

2.2.1 Mechanical Properties

Yield strength of CW15-15Ti irradiated to a dose of 83 dpa and tested at the irradiationtemperature is reported to be comparable to that of cold worked cold worked 316Ti7.Cold worked silicon-modified 15-15Ti possesses the best tensile properties; both yieldstrength and uniform elongation are higher than those of standard cold worked 316Ti atall the temperatures. Studies carried out at IGCAR on thermal creep properties of 20%cold worked 316 and 20% cold worked D9 have clearly established that D9 is strongerthan 316 SS14. As far as in-pile creep behaviour of D9 is concerned, very little

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information is available in the open literature. Tests conducted on 20% CW 316SS inEBR-II show that the in-pile rupture life is equal to or sometimes better than the out-of-pile rupture life15. Later tests conducted in FFTF Material Open Test assembly (MOTA)showed however that in-pile rupture life was slightly less than the out-of-pile rupturelife at higher temperatures16. This is shown in Fig.4, in which in-pile data at hightemperatures (LMP > 1.6x104) falls below the out of pile data for alloy D9. Aconservative design curve for D9 can be generated as one which passes through in-piledata at higher temperatures, but parallel to the out of pile curve. Out-of-pile creeprupture data generated on PFBR clad tubes are superimposed in Fig.4 for comparison.

Figure 4Stress versus Larson-Miller parameter plot16 for 20CW D9.

2.2.2 Sodium CorrosionWhereas corrosion effects of sodium are negligible for thick components, for thincomponents, effective loss of thickness due to leaching of solutes has to be accountedfor in the design. Clad tubes are the thinnest component in PFBR (~450 microns thick).Based on simulated laboratory experiments, and operating experience the world-over,corrosion allowance of 4 microns per year is considered sufficient. The tubes areexposed to a maximum period of about 2 years in the reactor. In addition, fuel sidecorrosion has to be included in the design burn-up analysis.

3.0 REACTOR ASSEMBLY3.1 Criteria for Selection of Materials

Major factors that are considered in the selection of materials for structural componentsof the reactor assembly are:

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(i) operating conditions and plant life(ii) availability of design data(iii) ease of fabrication(iv) international experience, and(v) cost.

The reactor assembly consists of core, grid plate, core support structure, main vessel,safety vessel, top shields and absorber rod drive mechanism (Fig.5). The sodiumtemperature in the primary cold pool during normal operastion is 670 K. The mean coreoutlet temperature will be about 820 K during operation, and 923 K under planttransient conditions. The environment of operation is liquid sodium or argon withsodium vapour or nitrogen gas depending upon the location of the component. Except

Figure 5

PFBR reactor assembly showing major components.

for the components near the core, like grid plate which may see an irradiation dose ofabout 1 dpa in 40 years of design life, for all other components such as inner vessel,main vessel, safety vessel, intermediate heat exchanger and pumps, irradiation is not aconsideration. Liquid sodium coolant with strict control of chemistry, particularly ofelements responsible for liquid metal embrittlement (As, Sb, Bi) and also of carbon andoxygen, does not pose any problem. In the primary sodium circuit, the primary stressesare low. However, secondary stresses of thermal origin are quite significant. Thesestresses are both steady and transient in nature. The secondary circuit will see largeprimary stresses of transient nature during any accidental big leak in the steamgenerator17,18. The design life of the structural components is 40 years. Creep, low cyclefatigue and creep-fatigue interactions are important considerations in the choice of

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materials. Sodium streams coming from the core channels are at different temperaturesand these, on impinging upon the component surface, produce local variations in thesurface temperature with associated stress fluctuations. This phenomenon called thermalstripping would lead to high cycle fatigue damage and this is another importantconsideration in choosing materials for structural components above the core. Theprincipal selection criteria for fast breeder reactor (FBR) structural materials aresummarised in Table 4.

Table 4Principal Selection Criteria for FBR Structural Materials

General Criterion Specific Criteria

Mechanical properties Tensile strenghthCreepLow cycle fatigueCreep-fatigue interactionHigh cycle fatigue

Design Inclusion in RCC-MR/ASME design codes.

Structural integrity Weldability

Workability

International experience

Easy availability

Lower cost

3.1.1 International Experience

Low alloy steels are not considered suitable for structural components of the primaryheat transport system since they do not possess adequate high temperature mechanicalproperties. Among stainless steels, ferritic stainless steels are not suitable because of (i)inadequate high temperature mechanical properties, (ii) 748 K (475 oC) embrittlement(iii) sigma phase embrittlement at high temperatures and (iv) difficulty in welding dueto grain coarsening. Martensitic stainless steels are prone to notch sensitivity, lowductility and suceptibility to embrittlement between 693 K and 823 K, and are thereforenot considered.

Austenitic stainless steels are chosen as the major structural materials in view oftheir adequate high temperature mechanical properties, compatibility with liquid sodiumcoolant, good weldability, availability of design data, good irradiation resistance andabove all the fairly vast and satisfactory experience in the use of these steels in sodiumcooled reactors. Designers of FBRs all over the world prefer monometallic constructionfor liquid sodium systems because of the concern of interstitial element transfer (carbonin particular) through liquid sodium due to the differences in thermodynamic activity in

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a bimetallic system. Hence, austenitic stainless steels are employed in the entire liquidsodium system even if the temperatures of some components are low enough to use lessexpensive ferritic steels. Table 5 lists the structural materials selected for majorcomponents such as reactor vessel, intermediate heat exchanger (IHX), and piping incurrently operating or designed FBRs all over the world12. The grades selected include304, 304L, 316, 316L, 321, 347 and their equivalents. Stabilised austenitic stainlesssteels 321 and 347 are less popular since their welds are prone to cracking duringwelding, during reheating and also in service. These steels have also poor creepductility. Table 6 gives the chemical composition of structural materials selected forEuropean Fast Reactor (EFR), Demonstration Fast Breeder Reactor (DFBR, Japan) andSuperphenix (France).

In high temperature sodium, austenitic stainless steels have good resistance togeneral corrosion and localised corrosion. Localised corrosion is absent since thesurface of steel is clean in sodium (no passive film) and electrochemical reaction is notpossible in non-aqueous medium. However, mass transfer of metallic elements in SScan take place under the influence of non-metallic impurities in liquid sodium such asoxygen and carbon. Oxygen leads to the formation of sodium chromite and carbonresults in carburisation or decarburisation which in turn influence the mechanicalproperties. Impurities other than carbon and oxygen such as chloride, calcium andpotassium are known to influence corrosion. Fast reactors and sodium loops have beensuccessfully run for many years testifying to the long-term compatibility of austeniticstainless steels with sodium at elevated temperatures.

3.2 Choice of 304L(N) and 316L(N) for Structural Components

For the structural components of Fast Breeder Test Reactor (FBTR) which attained firstcriticality in 1985 at Kalpakkam, austenitic stainless steel grade 316 is the principalgrade used. For PFBR low carbon austenitic stainless steel types 304 and 316, alloyedwith 0.06-0.08 wt% nitrogen, designated as 304L(N) and 316L(N) SS respectively havebeen selected for the structural components. Low carbon grades have been chosen toensure freedom from sensitisation during welding of the components to avoid risk ofchloride stress corrosion cracking during storage in coastal site. Since low carbongrades have lower strength than normal grades, nitrogen is specified as an alloyingelement to improve the mechanical properties so that the strength is comparable to 304and 316 SS. Although 304L(N) and 316L(N) are specified by ASME with nitrogen inthe range of 0.10 to 0.16 wt%, for PFBR, nitrogen content is limited to 0.08 wt% inview of improved weldability, code data availability and for minimising scatter inmechanical properties. For components operating at temperatures below 700 K, type304L(N) SS is preferred due to its lower cost while for high temperature componentsoperating in the creep range, 316L(N) SS has been favoured. Other major advantages ofaustenitic stainless steels 304L(N) and 316L(N) include existence of vast data base onmechanical properties including very long-term creep data, ease of availability andfabrication, and above all, the availability of design data in the RCC-MR code selectedfor PFBR design.

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Table 5Materials selected in FBRs for major components

Reactor Country ReactorVessel

IHX Primary circuitpiping

hot leg (cold leg)#

Secondarycircuit piping

hot leg(cold leg)

Rapsodie France 316 SS 316 SS 316 SS (316 SS) 316 SS (316 SS)

Phenix France 316L SS 316 SS (316 SS) 321 SS (304 SS)

PFR U.K. 321 SS 316 SS (321 SS) 321 SS (321 SS)

JOYO Japan 304 SS 304 SS 304 SS (304 SS) 2.25Cr-1Mo

(2.25Cr-1Mo)FBTR India 316 SS 316 SS 316 SS (316 SS) 316 SS (316 SS)

BN-600 Russia 304 SS 304 SS 304 SS 304 SS (304 SS)

SuperPhenix-1

France 316L(N) SS 316L(N) SS (304L(N) SS) 316L(N) SS

FFTF U.S.A. 304 SS 304 SS 316 SS (316 SS) 316 SS (304 SS)

MONJU Japan 304 SS 304 SS 304 SS (304 SS) 304 SS (304 SS)

SNR-300 Germany 304 SS 304 SS 304 SS (304 SS) 304 SS (304 SS)

BN-800 Russia 304 SS 304 SS 304 SS 304 SS (304 SS)

CRBRP U.S.A. 304 SS 304 and 316 SS 316 SS (304 SS) 316H (304H)

DFBR Japan 316FR SS 316 FR 316FR (304 SS) 304 SS (304 SS)

EFR Europe 316L(N) SS 316L(N) SS 316L(N) SS 316L(N) SS

# for pool-type reactor, there is no hot leg piping

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Table 6Chemical composition specified for 316L(N), 316FR and 316LN used/proposed in

EFR, DFBR and Superphenix, respectively. (single values denote maximum permissible, NS - not specified)

Element 316L(N) SS(EFR)

316FR(DFBR)

316L(N ) SS(Superphenix)

C 0.03 0.02 0.03

Cr 17-18 16-18 17-18

Ni 12-12.5 10-14 11.5-12.5

Mo 2.3-2.7 2-3 2.3-2.7

N 0.06-0.08 0.06-0.12 0.06-0.08

Mn 1.6-2.0 2.0 1.6-2.0

Si 0.5 1.0 0.5

P 0.025 0.015-0.04 0.035

S 0.005-.01 0.03 0.025

Ti NS NS 0.05

Nb NS NS 0.05

Cu .3 NS 1.0

Co .25 0.25 0.25

B .002 0.001 0.0015-0.0035

Nb+Ta+Ti 0.15

3.2.1 Chemical Composition

Table 7 gives the chemical compositions of types 304L(N) and 316L(N) SS specifiedfor PFBR alongwith the ASTM and RCC-MR specifications. The PFBR specificationsare more stringent than the ASTM specifications. The chemical composition rangeshave been narrowed down to reduce the scatter in mechanical properties. Thecomposition limits have been revised to meet specific property requirements.Chromium, molybdenum, nickel and carbon contents have been specified taking intoaccount intergranular corrosion resistance criteria developed from operating experiencewith nuclear reactors (both light water and fast reactors). Lower limits have beenspecified for carbon and nitrogen to ensure that the mechanical properties match thoseof 304 and 316 SS grades for which design curves are available in RCC-MR/ASME

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Table 7Comparison of PFBR specification for 304L(N) and 316L(N) SS with ASTM A240 and

RCC-MR RM-3331.(single values denote maximum permissible, NS - not specified)

Element ASTM304L(N)

PFBR304L(N)

ASTM-316L(N)

PFBR316L(N)

RCCMR316L(N)RM3331

C 0.03 0.024-0.03 0.03 0.024-0.03 .03

Cr 18-20 18.5-20 16-18 17-18 17-18

Ni 8-12 8-10 10-14 12-12.5 12-12.5

Mo NS 0.5 2-3 2.3-2.7 2.3-2.7

N 0.1-0.16 0.06-0.08 0.1-0.16 0.06-0.08 0.06-0.08

Mn 2.0 1.6-2.0 2.0 1.6-2.0 1.6-2.0

Si 1.0 0.5 1.0 0.5 0.5

P 0.045 0.03 0.045 0.03 0.035

S 0.03 0.01 0.03 0.01 0.025

Ti NS 0.05 NS 0.05 -

Nb NS 0.05 NS 0.05 -

Cu NS 1.0 NS 1.0 1.0

Co NS 0.25 NS 0.25 0.25

B NS 0.002 NS 0.002 0.002

Inclusion Contents (max.)

Type Thin Thick

Type A (sulphide) 1 0.5Type B (alumina) 2 1.5Type C (silicate) 2 1.5Type D (globular oxide) 3 2.0A+B+C+D 6 4.0

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codes. An upper limit has been specified for carbon to ensure freedom fromsensitisation during welding. The upper limit for nitrogen is lowered to 0.08 wt%compared to 0.16 wt% in ASTM and ASME specifications mainly on consideration ofimproving weldability and minimising scatter in mechanical properties. Phosphorus,sulphur and silicon are treated as impurities, as they have adverse effects on weldability.Therefore, acceptable maximum limits are reduced to values that can be achieved insteel making practice. Considering the adverse effects of titanium, niobium, copper andboron on weldability, maximum permissible limits have been imposed, although nosuch limits exist in ASTM specifications. A minimum level has been specified formanganese to improve weldability. Upper limit has been specified for cobalt to reduceCo60 activity induced by neutron irradiation, so as to facilitate ease of eventualmaintenance of the components of primary sodium system. In addition to more stringentcomposition limits, a specification for inclusions has been added keeping in view thatsulphide inclusions are most detrimental especially from welding considerations, andglobular oxides are least harmful. A grain size finer than ASTM No. 2 is specified so asto achieve optimum high temperature mechanical properties.

3.2.2 Mechanical Properties

316L(N) SS will be used for components experiencing relatively higher temperatures(above 770 K) while 304L(N) SS has been selected for the rest of the structuralcomponents since cost of 304L(N) SS is less by 20% in terms of material cost ( about15% if extra thickness required is also taken into account). During prolonged operationat elevated temperatures, stainless steels undergo microstructural changes such asprecipitation of carbides and brittle intermetallic phases. Embrittlement will not be aproblem for components operating below 700 K since precipitation is extremelysluggish at these temperatures. For Grid Plate, though temperatures are not in the creepregion, 316L(N) SS is preferred over 304L(N) SS in view of better ductility afterirradiation. Studies carried out at IGCAR (upto 10,000 hours) which are consistent withthe international experience have shown that creep rupture strength of 316L(N) grade issuperior to 316 SS; it has generally lower creep rates than type 316 SS. Theimprovement in properties is attributed to solid solution strengthening by nitrogen andprecipitation strengthening by fine carbides. Figure 6 shows a comparison of the creeprupture strength of 316L(N) with 316 SS. The 316L(N) data relate to long term creeptests (upto 60,000 hours) from breeder programmes in Germany19, Japan20, France20 andIndia. The reference 316 SS data are taken from the long term creep programme ofORNL, USA20. RCC-MR design curve for expected minimum stress to rupture issuperimposed in Fig.6. Creep rupture strength of 316L(N) SS is better especially forlonger rupture times.

Reasonable amount of data are available in the open literature on the LCF andcreep-fatigue interaction behaviour of nitrogen alloyed 316 SS intended for fast reactorapplications. Comparative evaluation of strain-controlled LCF behaviour of nitrogenalloyed 316FR (which is the Japanese fast reactor grade steel) and 316 S shows thatcontinuous cycling LCF life is not significantly influenced by minor variations incarbon and nitrogen contents20 (Fig.7). However, the minor changes in carbon andnitrogen have been found to have substantial effects on creep-fatigue interactionbehaviour. Hold-time tests conducted at tensile peak strain at elevated temperatures

18

clearly indicated that while increasing dwell time reduced fatigue lives in both 316 SSand its nitrogen-alloyed version, the extent of such reduction is smaller for nitrogen-alloyed 316 SS20 (Fig.8); the creep-fatigue interaction tests on 316FR SS had beenperformed concurrently in Japan and USA. Results of studies carried out at IGCAR areconsistent with the behaviour reported by other countries.

101 102 103 104 105100

200

300

400

500

Temperature: 873 K

316L(N) - Superphenix, France316FR - DFBR, Japan316L(N) - Germany316L(N) - PFBR, India316 - ORNL, USA316L(N) - RCC-MR design curve

Stre

ss, M

Pa

Rupture time, h

Figure 6Comparison of creep rupture strengths of 316 and 316L(N)

SSs from various countries19,20.

102 103 104 105 106

100

101

Japan Type 316FR, 1 x 10-3 s-1, 50 mm Plate ORNL Type 316FR, 1 x 10-3 s-1 or HIgher, 50 mm Plate U.S. Type 316, 4 x 10-3 s-1, 538-566 oC 316L(N)-PFBR, India

Tota

l Stra

in R

ange

(%)

Cycles to Failure

Figure 7Comparison of fatigue behaviour of 316 and 316FR SSs20.

19

10-3 10-2 10-1 100 101101

102

103

104

866-873 KTotal Strain Range 1.0%

316FR, Plate ORNL 316FR, Plate JAPC 316 Stainless Steel, Heat 81583 316 Stainless Steel, Heat 65808 316Stainless Steel, Heat 65808 Aged 5,040 h at 593 oC 316L(N), PFBR, India, Strain Range 1.2% 316L(N), PFBR, India, Strain Range 0.8% 316L(N), ICL (Levalillant et. al)

Zero Hold

Cycl

es to

Fai

lure

Tensile Hold Time (h)

Figure 8Effect of hold time on fatigue lives of 316 and 316L(N) SSs20.

3.2.3 Welding Consumables and Properties of Welds

Welding is extensively employed in the fabrication of FBR components. Weldmetal cracking and heat affected zone (HAZ) cracking are major areas of concern inwelding austenitic stainless steels. Weld metal cracking can be controlled by optimisingthe chemical composition of the welding consumables. The optimised chemicalcomposition for PFBR for 316(N) SS welding electrodes is given in Table 8 alongwiththe ASME specifications for E-316 SS. Carbon in the range of 0.045-0.055 wt% andnitrogen in the range of 0.06-0.1 wt% are specified to provide weld joints withimproved creep strength and freedom from sensitisation in the as-welded state. Inaddition, ferrite in the weld metal is specified to be between 3-7 ferrite number (FN) topromote ferritic solidification mode. A minimum of 3 FN is specified to ensure freedomfrom hot cracking in the weld metal. Because delta-ferrite undergoes phase changes tocarbides and brittle intermetallic phases at high temperatures, an upper limit of 7 FN hasben specified. Nitrogen in the specified range has no detrimental effect on weldability of316L(N) stainless steel. Heat-affected zone (HAZ) cracking is avoided by specifyinglower permissible limits for P,S and Si and also by specifying limits on B, Ti and Nbwhich are not specified in the ASTM standards for the base metal. 316(N) SS electrodeswould be utilised for welding of both 316L(N) SS and 304L(N) SS base materials. Thiswould avoid any mix-up of electrodes in welding if a different electrode is selected for304L(N) SS. 16-8-2 filler wire will be used for TIG welding since this composition hasbetter microstructural stability, creep strength and toughness. 16-8-2 electrodes forMMA welding are not easily available.

Evaluation of creep properties of 316(N) weld metal has shown that its creeprupture strength and ductility are better than those of 316 SS weld metal. Figure 9 showsa comparison of creep rupture strengths of 316 and 316(N) weld metals at 923 K; about30% increase in rupture strength was observed by alloying with nitrogen21. There is noinformation available in the open literature on the LCF behaviour of 316(N) welds andtheir joints. Detailed investigations conducted at IGCAR at 873 K revealed the LCF life

20

in the order; base metal> weld metal> weld joint. The rank order of lives exhibited by316(N) and its weld joint is very much similar to those shown by stainless steels 304,316 and their weld joints. The poor strain-controlled fatigue resistance of weldments isattributed to the presence of coarse grains in the HAZ which act as a metallurgical notchleading to shortening of the crack initiation phase.

Table 8PFBR specifications for Modified 316 electrodes and permissible limits for

delta-ferrite as per WRC-92 FN diagram. (single values specified aremaximum permissible, NS- not specified).

Element ASME SFA 5.4E 316

PFBR316(N)

C 0.08 0.045-0.055

Cr 17-20 18-19

Ni 11-14 11-12

Mo 2-3 1.9-2.2

N NS 0.06-0.10

Mn 0.5-2.5 1.2-1.8

Si 0.9 0.4-0.7

P 0.04 0.025

S 0.03 0.02

Ti+Nb+Ta NS 0.1

Cu 0.75 0.5

Co NS 0.2

B NS 20 ppm

δ-Ferrite 3-10 FN 3-7 FN

21

10 100 1000 100008090

100

200

300923 K

316(N) SS (all-weld)316 SS (all-weld)

Stre

ss, M

Pa

Rupture Life, h

Figure 9Comparison of creep rupture strengths of 316 and 316(N) SS weld metals21.

4.0 STEAM GENERATORIn PFBR, the Steam generator is a vertical, countercurrent shell-and-tube type heatexchanger with sodium on shell side, flowing from top to bottom, and water/steam ontube side. The flow to the tube bundle entry is made uniform by the flow distributiondevice (annular perforated plates) located in the annular region before sodium entry tothe tube bundle. The tubes are placed in triangular pitch. Each tube is provided withexpansion bend in sodium flow path. The tubes are supported at various locationsincluding at the middle of tube expansion bend.

The very high reactivity of sodium with water makes the steam generator a keycomponent in determining the efficient running of the plant and demand high integrityof steam generator components. The high integrity of SG components can be achievedby choosing a proper material followed by an optimised design and fabrication. It isdecided to manufacture the steam generator in mono-metallic material (tube, shell andthick section tube-sheet/plate), since employing a single structural material enhances thereliability of the critical tube to tube sheet welds. Modified 9Cr-1Mo ferritic steel hasbeen selected for all the steam generator components. The selection of Modified 9Cr-1Mo steel is based on several important considerations and these are described in thefollowing sections.

4.1 Criteria for Selection of MaterialsThe principal selection criteria of materials for steam generator are given in Table 917.These include the general criteria as well as the criteria directly related to the use ofmaterials in sodium-heated steam generator. Materials selected for LMFBR steamgenerator application should meet requirements of high temperature service such as

22

high temperature mechanical properties including creep and low cycle fatigue,resistance to loss of carbon to liquid sodium which leads to reduction in strength,resistance to wastage in case of small leaks leading to sodium-water reaction andresistance to stress corrosion cracking in sodium and water media.

Table 9Principal Selection Criteria for LMFBR Steam Generator Material.

General Criteria Criteria related to use in sodium

Mechanical properties- Tensile strength- Creep strength- Low cycle fatigue and High cycle fatigue- Creep-fatigue- Ductility- Ageing effects

Mechanical properties in sodiumSusceptibility to decarburisation

Inclusion in pressure vessel codes oravailability of adequate data

Corrosion under normal sodiumchemistry condition, fretting and wear

Corrosion resistance under storage(pitting) normal and off-normalchemistry conditions

Corrosion resistance in the case ofsodium water reaction (Stress corrosioncracking, self enlargement of leak andimpingement wastage)

Workability

Weldability

Availability

Cost

4.1.1 International Experience

The choice of materials for FBR steam generators in operation or under design worldwide is presented in Table 10. The sodium inlet and steam outlet temperatures for thesesteam generators are also included in this table. In the case of PFBR steam generator,the sodium inlet temperature is 798 K whereas the steam outlet temperature is 766 K. Itcan be seen from this Table that 2.25Cr-1Mo steel, either plain or stabilised grade, hasbeen used in evaporators, whereas in superheaters austenitic stainless steels have alsobeen used. However, the recent trend favours the use of Modified 9Cr-1Mo steel forLMFBR steam generator applications.

4.2 Choice of Modified 9Cr-1Mo Steel

For PFBR steam generator, a range of materials starting from ferritic steels (2.25Cr-1Mo, Nb stabilised 2.25Cr-1Mo, 9Cr-1Mo (grade 9), Modified 9Cr-1Mo (grade 91)),

23

austenitic stainless steels (AISI 304/316/321) and alloy 800 were examined22. In view ofthe poor resistance to aqueous stress corrosion cracking (SCC), austenitic stainlesssteels of 300 series were not considered for the steam generator. Alloy 800 shows betterresistance to SCC than austenitic steels, but it is not immune to stress corrosion crackingin chloride and caustic environments. Therefore, ferritic steels are the most preferred forsteam generator applications (Table 10). Among the ferritic steels, 2.25Cr-1Mo and9Cr-1Mo steels and their variants were considered for the steam generator. The basis forchoosing Modified 9Cr-1Mo (grade 91) and its relative properties with other candidatematerials are described in the following sections.

Table 10

Materials selected for steam generator in major FBRs

Tubing materialReactor Sodiuminlet(K)

Steamoutlet

(K)Evaporator Superheater

Phenix 823 785 2.25Cr-1Mo2.25Cr-1Mo stabilised 321 SS

PFR 813 786 2.25Cr-1Mo stabilisedReplacement unit in 2.25Cr-1Mo

316 SSReplacementunit in 9Cr-1Mo

FBTR 783 753 2.25Cr-1Mo stabilised

BN-600 793 778 2.25Cr-1Mo 304 SS

Super

Phenix-1

798 763 Alloy 800(once through integrated)

MONJU 778 760 2.25Cr-1Mo 304 SS

SNR-300 793 773 2.25Cr-1Mo stabilised 2.25Cr-1Mostabilised

BN-800 778 763 2.25Cr-1Mo 2.25Cr-1Mo

CRBR 767 755 2.25Cr-1Mo 2.25Cr-1Mo

DFBR 793 768 Modified 9Cr-1Mo (grade 91)(once through integrated)

EFR 798 763 Modified 9Cr-1Mo (grade 91)(once through integrated)

Before describing the various properties of grade 91 steel which led to itsselection for PFBR steam generator, it is important to mention the material specificationfor steam generator, which is close to that specified in ASTM. The chemicalcomposition (Table 11) is controlled within close limits to avoid scatter in themechanical properties. Lower limits are specified for residual elements, like sulphur,phosphorous and silicon to improve weldability and reduce the inclusion content toensure a high degree of cleanliness.

24

4.2.1 Mechanical PropertiesElevated temperatures tensile properties data on grade 91 steel are available in theliterature23 as well as in design code (RCC-MR)24. The strength of this steel is found tobe higher than those of 2.25Cr-1Mo and plain 9Cr-1Mo steels. Our own efforts todevelop the steel indigenously and the preliminary evaluation of its tensile propertieshave led to expected results25. The yield and ultimate tensile strength values of Modified9Cr-1Mo steel (12 mm plates developed and supplied by MIDHANI Ltd., Hyderabad)in normalised and tempered condition are found to be higher than the average strengthvalues reported in the literature, as well as those specified in RCC-MR code.

Table 11PFBR specifications for Modified 9Cr-1Mo Steel tubes and welding consumables.

Element Base Metal(tube)

Filler wire(weld deposit)

Electrode

C 0.08 – 0.12 .08-.12 0.08-0.12

Cr 8.00 – 9.00 8-9.5 8.0-9.5

Mo 0.85 – 1.05 0.85-1.05 0.85-1.05

Si 0.20 – 0. 50 0.2-0.4 0.2-0.3

Mn 0.30 – 0.50 0.5-1.2* 0.5-1.2*

V 0.18 – 0.25 0.15-0.22 0.15-0.22

Nb 0.06 – 0.10 0.04-0.07 0.04-0.07

N 0.03 – 0.07 0.03-0.07 0.03-0.07

S 0.01 max. 0.01 max. 0.01 max.

P 0.02 max. 0.015 max. 0.01 max.

Cu 0.10 max. 0.25 max.

Ni 0.20 max. 0.6-1.0* 0.6-1.0*

Al 0.04 max. 0.04 max

Sn 0.02 max. --

Sb 0.01 max. --

Ti 0.01 max. --

Note: Ni + Mn less than or equal to 1.5

The creep strength of Modified 9Cr-1Mo steel is significantly higher than thoseof the conventional 2.25Cr-1Mo and plain 9Cr-1Mo steels over a wide temperaturerange and is greater or equal to that of AISI type 304 austenitic stainless steel up to 873K (Fig. 10)23. Comparison of creep strengths of various grades of ferritic steels widelyused for high temperature applications suggests that Modified 9Cr-1Mo steel exhibits

25

higher creep strength than most of the other materials26 (Fig. 11). (In Fig.11, open circleis used as the symbol to represent nine different grades of ferritic steels which convergeat long test durations). Though for shorter creep lives, 12Cr-1Mo-1W-0.3V exhibitshigher creep strengths, for larger creep lives, its creep strength approaches that of lowaloy feritic steels because of microstructural degradation. Modified 9Cr-1Mo steel doesnot exhibit such a drastic reduction in creep strength at longer durations due to thestability of the microstructure; its creep strength remains higher than that of severalferritic steels at longer test durations. It is confirmed by the data generated at IGCARwhich is also included in Fig. 11. This is the most important aspect favouring theselection of Modified 9Cr-1Mo steel for steam generator. Further, the higher creepstrength of Modified 9Cr-1Mo steel allows the use of comparatively thinner tubes. Thethinner tube and higher thermal conductivity of this material reduce the heat transferarea requirements of the steam generator.

Figure 10Comparison of 105 h creep rupture strengths of several materials22.

Comparison of low cycle fatigue (LCF) data of plain 9Cr-1Mo and Modified9Cr-1Mo steel (both hot rolled and hot forged) indicate that Modified 9Cr-1Mo steelexhibits higher continuous cycling LCF resistance compared to plain 9Cr-1Mo steel27.However, hot forged Modified 9Cr-1Mo steel shows lower fatigue life compared to the

26

hot rolled material. Detailed investigations performed on LCF behaviour of forged thicksection tube plate (1000 mm dia. and 300 mm thickness) of plain 9Cr-1Mo steel atIGCAR has revealed inferior fatigue life of the forged product compared to hot rolledthin section. The reduction in LCF resistance of forged alloys has been attributed to theshortening of crack initiation phase associated with relatively coarse grain size.

Figure 11Creep-rupture strength of eleven types of ferritic heat resistant steels. Open

circle is used as the symbol to represent nine different grades of ferriticsteels which converge at long test durations26.

With the introduction of tensile hold, fatigue life of Modified 9Cr-1Mo steeldecreases rapidly with increasing hold time up to 1 h28. Tensile hold of 1 h reduces lifeto ~ 1/4 of the continuous cycling life. Beyond 1 h, there is insignificant decrease in lifewith increasing hold time in the range 1-3 h. This implies that hold times of longerduration would not degrade creep-fatigue performance of the material. However, a fewstudies of larger duration holds (~ 10 h) are being planned to confirm this behaviour.Reduction in life under creep-fatigue condition is attributed primarily to the reduction instrength of the steel due to microstructural degradation associated with the coarsening ofprecipitates and dislocation substructure.

4.2.2 Welding Consumables

For the welding of Modified 9Cr-1Mo steel, consumables having composition closelymatching with that of the base metal are normally employed. However, achievingrequired toughness in the weld metal after PWHT has been a problem in this steel,especially in the case of shielded metal arc (SMA) welds. Therefore, in theAWS/ASME specification for the consumables, minor modifications have been madefor Ni, Mn, Nb, V and N. In addition, the specification calls for determination of

27

RTNDT, and use of only alloyed core wire for making SMA welding electrodes. Thespecified composition is given in Table 11.

4.2.3 Use of Modified 9Cr-1Mo Steel for Thick Sections

Apart from superior elevated temperature mechanical properties of 9Cr-1Mo grades ofsteels as compared to 2.25Cr-1Mo steel, the ability and tolerance for providing nearly aconstant microstructure over a large section size is another aspect in the selection ofModified 9Cr-1Mo steel for steam generator. This in turn ensures only a small variationin mechanical properties with increasing thickness or between the surface and the centreof thick section products. The microstructural stability and the mechanical properties of9Cr-1Mo steel and its variants for large section sizes have been examined in detaileither by simulated heat treatments on bar material or by producing the actual thicksections27,29-33. These investigations show encouraging results for this steel to be usedfor thick sections.

4.2.4 Trimetallic Transition Joint

As the main structural and piping material is austenitic 316LN SS, and the steamgenerator material is Modified 9Cr-1Mo steel, a dissimilar weld involving these twomaterials is inevitable in the fabrication of steam generator. A large number ofpremature failures of the direct joint involving austenitic stainless steel and ferritic steeloperating at high temperature have been reported in the past, mainly from fossil powerplants. The failures are mainly attributed to (i) large differences in the thermalexpansion coefficients of these two steels, which lead to generation of thermal stressesduring start-up and shut-down, (ii) difference in the creep strength of these materials,and (iii) carbon migration from ferritic steels to austenitic steels leading to formation ofsoft zone near the interface. It has been shown that introduction of an intermediate pieceof material, having thermal expansion coefficient value between austenitic SS andferritic steel, can significantly reduce the thermal stresses generated. (Coefficient ofthermal expansion for 316L(N) SS and 9Cr-1Mo steel are 18.5 µm/m/K and 12.6µm/m/K respectively). Accordingly, a trimetallic joint configuration in which an Alloy800 spool piece welded to 316L(N) SS pipe on the one side, and Modified9Cr-1Mosteel pipe on the other side, is chosen for this dissimilar joint. Although alloy 600 isanother material being used for transition joints, alloy 800 has been preferred over alloy600 since the material is included in ASME code and also was the choice of transitionjoint (2.25Cr-1Mo -SS) for CRBRP steam generator. For welding Alloy 800 toModified 9Cr-1Mo, Inconel 82/182, welding consumable is recommended. For weldingAlloy 800 to 316L(N) SS, 16-8-2, filler wire is selected.

5.0 TOP SHIELD (ROOF SLAB)

Roof slab, along with rotating plugs and control plug forms the top shield, which is thetop cover for the main vessel (Fig. 5). It provides biological and thermal shielding in thetop axial direction of the reactor. It also acts as support for various components such asmain vessel, pumps, intermediate heat exchangers, decay heat exchangers, control plug,in-vessel transfer machine etc. Roof slab is essentially a box structure with top andbottom plates interconnected by vertical cylindrical shells and radial stiffeners weldedto them. The gap between top and bottom plates excluding the space occupied by its

28

cooling system is filled with concrete. Diameter of the roof slab is 12.9 m and its heightis 1.8 m.

5.1 Criteria for Selection of Material

5.1.1 Design Conditions

Mechanical load coming on the roof slab is quite high. Self weight of the roof slabalong with that of filled-in concrete would be around 650 t. As it supports main vessel,pumps, heat exchangers, rotatable plugs etc, weight of these components are alsotransferred to the roof slab. Thermal loads coming on the roof slab is relatively less.During the normal operation of the reactor, the temperature experienced by the roof slabwould be of the order of 373–393 K. In the event of loss of cooling, the temperaturecould rise up to a maximum of 473 K. The minimum operating temperature is theambient temperature (298 K).

Radiation field close to the roof slab is relatively low. Neutron flux at the bottomof the roof slab is 105 n/m2s-1 while gamma flux is of the order of 3x104 R/h34. Theenvironment to which the roof slab would be exposed is also not very severe. Its bottomsurface would be exposed to argon gas containing sodium vapour (reactor cover gas)while the top surface would be exposed to air in the reactor containment building at 308K. For design, safety and fabrication, roof slab is classified as a Class I component andit is designed as per RCC-MR code.

The main property that the roof slab materials should possess is good mechanicalstrength in the temperature range 298–493 K. As the structure is massive, fabricationwould involve extensive welding and hence the weldability of the material should begood. Stress relieving after welding would be difficult and hence choice of material andthickness should be such that post weld heat treatment is not called for. Further,reference nil ductility transition temperature (RTNDT) after irradiation should be 33 Kbelow the minimum operating temperature. Material should also be compatible withsodium mist laden argon gas. It should be possible to weld the roof slab material withaustenitic stainless steel, the material chosen for main vessel and other primarystructural components. Finally material should be of low cost and easily available.

5.2 International Experience

Materials chosen for roof slab fabrication in other fast breeder reactors is shown inTable. 12. Steels chosen for Phenix and Superphenix, A42P2 and A48P2 respectivelyare carbon steels specified in French code RCC MR35,36. Similarly, A516 Gr 65 is alsocarbon steel specified in ASTM standard for moderate and low-temperature service.Thus, choice of material for all the reactors has been ferritic steels, with carbon steelbeing the most common choice. This is because properties of these steels are goodenough to meet the design, fabrication and service requirement for the roof slab.

29

Table 12Materials selected for Top Shield for various Fast Breeder Reactors.

Reactor Material

Phenix Carbon steel (A42P2)

Superphenix-1 Carbon steel (A48P2)

Superphenix-2 Carbon steel

PFR Carbon steel

FFTF Carbon Steel

CRBR Low Alloy Steel

EFR Carbon steel (A48P2)

5.3 Choice of A48P2 Grade Carbon Steel for Top Shield (Roof Slab)

Carbon steels meant for pressure vessel applications at moderate and low temperatureswith very good notch toughness meet all the above requirements including those of costand availability. Austenitic stainless steels would also meet all these requirements,though cost is much higher than that of the carbon steel. Advantage of austeniticstainless steel material would be that there is no mandatory in-service ultrasonicinspection required as in the case of a bimetallic weld joint. However, as the completedmain vessel is planned to be ultrasonic inspected to be inspected as part of in-serviceinspection, introduction of bimetallic weld is not a real disadvantage. Therefore, specialgrade (low carbon) A48P2 steel specified in French code RCC-MR is chosen as thematerial of construction for the roof slab. ASTM A516, Grad 65 is another material thatsatisfied all the requirements like A48P2. With the decision to use RCC-MR codes fordesign and fabrication, the French specification was adopted since it takes care ofspecific requirements for roof slab.

5.3.1 Lamellar Tearing

For fabrication involving steel plates of large thickness (typically > 20 mm), animportant factor to be considered is the resistance to lamellar tearing. It is a form ofcracking which occurs in the base metal during welding, if the volume fraction ofelongated sulphide or silicate inclusions in the rolling direction of the steel is high37.Stresses developed during welding, can lead to debonding at the interface between theseinclusions and the matrix. Cracks thus formed link together forming characteristic steplike appearance. It is more common in T-type joints involving thick sections in whichweld fusion line is parallel to the surface, and welding stresses are high in the throughthickness direction. A number of T-welds between radial stiffeners and top/bottomplates are present in the roof slab. Lamellar tearing was one of the major problemsencountered during manufacturing of the roof slab for the Prototype Fast Reactor

30

(PFR)38. Cracks were observed in the material below the cruciform joints and repeatedrepair of the joints was necessary before the required standards were achieved.

Resistance to lamellar tearing of the steel is improved by reducing sulphur in thesteel as it would bring down the inclusion content. Refining precesses like vacuumdegassing employed during steel making process brings down sulphur level in the steel.Susceptibility of a steel product to lamellar tearing can be assessed from its ductility inthe short transverse direction. A minimum reduction in area of 20% in a tension testusing a specimen with loading axis as the short transverse direction of the steel productis required to ensure that it is resistant to lamellar tearing39. Thus by specifying lowsulphur content in the steel, vacuum degassing during steel making and minimumductility in short transverse direction, it is possible to ensure that the lamellar tearingdoes not occur during roof slab fabrication. Chemical composition of this steel is givenin Table 13. Maximum limit for sulphur in A48P2 is 0.012 wt.%. Manufacturingprocess of this steel includes degassing under vacuum, which will lead to very lowinclusion content. Low sulphur content and vacuum degassing would drastically reducethe susceptibility to lamellar tearing. Maximum limit for carbon content is 0.22 wt %and Mn content is specified in the range of 0.8-1.5 wt.% to improve weldability and toprovide the tensile strength and toughness. The minimum short transverse ductilityspecified for A48 P2 steel is 35%. As already mentioned short transverse ductility is anindex of resistance of the steel to lamellar tearing and a high value of short transverseductility ensures that the steel is free from lamellar tearing. Stress relieving heattreatment is not mandatory for A48P2 steel up to a thickness of 35 mm40.

Table 13PFBR specification for A48P2 carbon steel plate for roof slab.

Composition* (wt.%) Mechanical Properties#Steel

C Mn Si P S Cr Ni Mo YS

MPa

UTSMPa

El.,%

STD,%

A48P2 0.22 0.8-1.5

0.4 0.035 0.012 0.2 0.3 0.1 285 470-550

25 35

* Where range is not specified it is the maximum permissible limit# Where range is not specified, it is the minimum value requiredSTD-short tranverse ductility measured as reduction in area

The fabrication experience of the roof slab structure of Superphenix-1 hasconfirmed that no preheating/post-weld heat treatment is required for thickness up to 35mm during welding. Further, it has been decided to introduce an additional requirementof nil ductility transition temperature (RTNDT) as less than or equal to 253 K duringprocurement to ensure good toughness.

5.3.2 Radiation Embrittlement

In principle, effect of radiation on roof slab material should also be considered. If thematerial chosen is austenitic stainless steel, this is not a serious concern. However, in

31

the case of carbon steels, ductile brittle transition temperature increases with neutronfluence. The neutron flux at the bottom of the roof slab is 105n/cm2s-1. The fast neutronflux (E > 0.1 eV) below the roof slab is 10 n/cm2s-1 which gives a fluence of2.19x1011n/cm2over a period of 40 years. This fluence is negligible and is not expectedto produce any significant change in DBTT. Hence, effect of radiation can be neglected.

5.3.3 Welding Consumable

Most of the welding involved in the roof slab is similar welding between carbonsteel and carbon steel, except in the bottom plate side where there is a dissimilar jointinvolving carbon steel (A48P2) and austenitic stainless steel (316L(N)). Electrodes forcarbon steel welding conform to SFA 5.1 of ASME section II-C with minormodifications slated further. The specification for chemical composition of electrodes isgiven in the Table 14. The ASME specification has been made tighter in the case ofelements C, S and P as per RCC-MR specifications. RTNDT values also have beenspecified. For dissimilar weld involving carbon steel to stainless steel, E309 weldingelectrode conforming to SFA-5.4 of ASME section II-C is selected. This is a highly-alloyed (22-25wt.% Cr and 12-14 wt.% Ni) austenitic stainless steel weldingconsumable which ensures that no hard phases, like martensite, are formed in the weldmetal due to dilution from the carbon steel side, and the microstructure is austenitic withsufficient volume fraction of δ-ferrite, to prevent hot cracking.

Table 14PFBR specification for carbon steel welding consumable chosen for roof slab

fabrication (wt.%). The values indicate the maximum limit.

Element C Mn Si Ni Cr Mo V P S

Electrode 0.08 1.2 0.6 0.3 0.2 0.3 0.04 0.02 0.020

6.0 SAFETY GRADE DECAY HEAT REMOVAL (SGDHR) SYSTEM

During reactor shut down, the decay heat generated in the core is removed through thenormal heat transport path consisting of primary sodium, secondary sodium and watersteam circuits. System for decay heat removal (DHR) through this normal heat transportpath is known as Operation Grade Decay Heat Removal (OGDHR) system. However,the normal heat transport path will not be available during the shut down of the reactorunder the following conditions.

1. Steam/ Water side is not available2. Off site power failure (grid failure)3. Both secondary sodium loops are not available

Under such situations, decay heat from the reactor core is removed through safety gradedecay heat removal (SGDHR) system. Figure 12 shows the SGDR system that has beendesigned for PFBR. This system consists of (i) Decay Heat Exchanger (DHX) (ii)Sodium to Air Heat Exchanger (AHX) and (iii) piping connecting DHX and AHX.

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DHX is immersed in primary sodium hot pool and it transfers heat from primary sodiumto intermediate sodium. AHX transfers the heat from intermediate sodium to outside air.In SGDHR, sodium flow between DHX and AHX is maintained by temperaturedifference of the intermediate sodium and the distance between the thermal centres ofDHX and AHX. AHX is provided with a stack on the air side that causes the airflow bynatural convection for cooling the intermediate sodium. During normal operation of thereactor, SGDHR is in the poised condition, maintaining only limited flow ofintermediate sodium and the air dampers in AHX are kept partially open to facilitatesmooth transition to decay heat removal state. SGDHR is designed with a capacity of

removing 8MW of power per loop and 4 independent circuits are provided. The designlife of the system is 40 years. Under poised condition, AHX temperature will be nearlysame as the hot pool sodium temperature (820 K).

Material chosen for DHX, and piping between DHX and AHX is 316L(N)stainless steel, same as the main structural material for PFBR. The operating conditionsexperienced by DHX and piping would not be much different from that of the IHX.AHX has finned tubes, the fins of which are helically wound and welded to the tubesusing high frequency resistance welding process. The heat transfer in AHX is from hotsodium to air. Thus, tubes in AHX are directly in contact with air which would containmoisture and salt content. Any leak in the tube would expose sodium to air leading tosodium fire. Hence, tube material shall be resistant to various forms of corrosion likeuniform corrosion, pitting corrosion etc. A systematic study on the susceptibility of

Reactor

M

Dump Tank

Air Stack

Sodium - Air Heat Exchanger

Argon

Expansion Tank

Air

Nitrogen

Air

Dampers

Air

T

Argon

DHX

Casing

T

MHot pool

Figure 12Safety grade decay heat removal system (SGDHR) designed for PFBR.

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316L(N), Alloy 800 and 9Cr-1Mo steels to various forms of corrosion has concludedthat 9Cr-1Mo steel can be chosen for this application from corrosion point of view.9Cr-1Mo is also the cheapest among these three materials. Since Modified 9Cr-1Mosteel and not plain 9Cr-1Mo steel has been chosen for steam generator, the samematerial may be selected for AHX tubes also. In view of non-availability of finned tubesin Gr.91, with fin material same as tube material, AISI 409 SS has been selected as thefin material based on industrial response..

7.0 SUMMARY

Selection of materials for different components of PFBR, a 500MWe sodium cooledFast Breeder Reactor, has been carried out considering various factors such as operatingconditions, weldability fabricability, international experience and cost. For corecomponents of the reactor radiation damage leading to void swelling, irradiation creepand embrittlement is the major determinant of the choice of materials. Alloy D9stainless steel confirming to ASTM A771 (20% CW 15Cr-15Ni-Mo-Ti-Si) has beenrecommended for the clad and wrapper tubes for the initial cores. Subsequent core sub-assemblies would have improved version of D9 as clad and 9Cr 1Mo as the wrapper.For structural components forming part of reactor assembly, 304L(N) and 316L(N)stainless steels are recommended. 304L(N) is to be used for components, which willhave lower temperature where creep is not important. Low carbon (less than 0.03%) hasbeen specified to avoid sensitisation during welding and nitrogen additions are specifiedto increase the strength of the material comparable to normal carbon versions of thestainless steels. Steam generator is a very critical component of the plant as sodium andwater/steam are separated by a single wall. Stress corrosion cracking is an importantconsideration for material selection. Modified 9Cr1Mo steel has been recommended forthe steam generator components. For top shield components including the roof slab,through thickness ductility and impact toughness are important considerations formaterial selection. Carbon steel grade A48P2 (given in RCC-MR) has beenrecommended. For the Na-air heat exchanger modified, 9Cr-1Mo steel for tubes andAISI 409SS as fin material has been recommended. These materials have very closecontrol on composition and reduced levels of residual elements to improve reliability ofreactor components.

ACKNOWLEDGMENT

The authors gratefully acknowledge discussions with many colleagues in Materials andMetallurgy Group and Reactor Group of Indira Gandhi Centre for Atomic Research.

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