Comprehensive nuclear materials 4 04 radiation effects in nickel based alloys ,

Comprehensive nuclear materials 4 04 radiation effects in nickel based alloys , Comprehensive nuclear materials 4 04 radiation effects in nickel based alloys , Comprehensive nuclear materials 4 04 radiation effects in nickel based alloys , Comprehensive nuclear materials 4 04 radiation effects in nickel based alloys , Comprehensive nuclear materials 4 04 radiation effects in nickel based alloys , Comprehensive nuclear materials 4 04 radiation effects in nickel based alloys ,

R M Boothby

National Nuclear Laboratory, Harwell, Oxfordshire, UK

ß 2012 Elsevier Ltd All rights reserved.

Abbreviations

AGR Advanced gas-cooled reactor

DFR Dounreay Fast Reactor

EBR-II Experimental Breeder Reactor-II

EDX Energy dispersive X-ray

HVEM High-voltage electron microscope

N/2 dpa calculated according to half-Nelson

model

NRT dpa calculated according to Norget,

Robinson, and Torrens model

STA Solution treated and aged

TEM Transmission electron microscope

UTS Ultimate tensile strength

VEC Variable energy cyclotron

4.04.1 Introduction

Research into the effects of irradiation on

nickel-based alloys peaked during the fast reactor

develop-ment programs carried out in the 1970s and 1980s

Interest in these materials focused on their high

resis-tance to radiation-induced void swelling compared to

austenitic steels, though a perceived susceptibility to

irradiation embrittlement limited their application tosome extent Nevertheless, the Nimonic alloy PE16was successfully used for fuel element cladding andsubassembly wrappers in the United Kingdom, andInconel 706 was utilized for cladding in France Both

of these materials are precipitation hardened andconsequently have high creep strength, and muchresearch and development of alternative alloys wasdirected toward maintaining swelling resistance andcreep strength while aiming to alleviate, or at leastunderstand, irradiation embrittlement effects Therehas been some revival of interest in nickel-basedalloys for nuclear applications, and various aspects ofradiation damage in such materials have recently beenreviewed by Rowcliffe et al.1in the context of Gener-ation IV reactors, and by Angeliu et al.2in consider-ation of their use for the pressure vessel of thePrometheus space reactor Nickel-based alloys arealso candidate structural materials for molten saltreactors, for which resistance to corrosion by moltenfluoride salts and high-temperature creep strengthare prime requirements, though intergranular attack

by the fission product tellurium and irradiationembrittlement due to helium production are poten-tially limiting factors for this application.3

This chapter focuses on the void swelling ior, irradiation creep, microstructural stability, andirradiation embrittlement of precipitation-hardenednickel-based alloys Fundamental to all of theseeffects are the basic processes of damage production

behav-123

by the creation of vacancies and interstitial atoms

in displacement cascades, and the ways in which

these point defects migrate and interact with, causing

the redistribution of, solute atoms Detailed

discus-sions of damage processes and radiation-induced

segregation are beyond the scope of this chapter

but these topics will be introduced where necessary,

particularly in relation to void swelling models

More detailed reviews are given in Chapter 1.01,

Fundamental Properties of Defects in Metals;

Chapter 1.03, Radiation-Induced Effects on

Microstructure;Chapter1.11, Primary Radiation

Damage Formation; Chapter 1.12, Atomic-Level

Level Dislocation Dynamics in Irradiated Metals

andChapter1.18, Radiation-Induced Segregation

Typical compositions of nickel-based alloys and

some precipitation-hardened steels, which are

con-sidered in this chapter, are shown inTable 1 Alloy

compositions are generally given in weight percent

throughout this chapter unless stated otherwise

Precipitation-hardened alloys may be utilized in a

number of different heat-treated conditions, which

are generally abbreviated here as ST (solution

trea-ted), STA (solution treated and aged), and OA

(over-aged) Further information on the material properties

of nickel alloys is given in Chapter 2.08, Nickel

Alloys: Properties and Characteristics

Neutron fluences are generally given for

E > 0.1 MeV unless indicated otherwise Atomic

dis-placement doses (dpa) are generally given in NRT

(Fe) units, although the half-Nelson (N/2) model was

widely used particularly in the United Kingdom inthe 1970s6 The exact relationship between theseunits will vary depending on the neutron spectrum(which may differ, not only from one reactor toanother, but also depending on location within areactor), but approximate conversion factors for fastreactor core irradiations are

1026n m
2ðE > 0:1MeVÞ

¼ 5dpa NRT Feð Þ ¼ 6:25dpa N=2ð Þ

4.04.2 Void Swelling

4.04.2.1 Compositional Dependence ofVoid Swelling

Nimonic PE16 was first identified as a low-swellingalloy in the early 1970s Void swelling data derivedfrom density measurements on fuel pin cladding mate-rials from the Dounreay Fast Reactor (DFR) werereported by Bramman et al.7and were complemented

by electron microscope examinations described byCawthorne et al.8Swelling in STA PE16 was found to

be lower than in heat-treated austenitic steels andcomparable to cold-worked steels Comparison ofdata for PE16 and FV548 (a Nb-stabilized austeniticsteel) irradiated under identical conditions in DFR to

a peak neutron fluence of6  1026

n m
2indicatedthat the lower swelling of PE16 was due to smallervoid concentrations at irradiation temperatures up

to 550C and reduced void sizes at higherTable 1 Nominal compositions (wt%) of commercial and developmental nickel-based alloys

aComposition indicated by Yanget al 4

bComposition indicated by Toloczkoet al 5

temperatures At around the same time, Hudson et al.9

compared the swelling behavior of PE16, type 316

steel, and pure nickel, using 20 MeV C2+ion

irradia-tions in the Harwell VEC (variable energy cyclotron)

The materials were implanted with 10 appm (atomic

parts per million) of helium prior to ion bombardment

to peak displacement doses>200 dpa (N/2) at 525C.

Void swelling in 316 steel and nickel exceeded 10% at

the highest doses examined, compared to0.5% in

PE16 Void nucleation appeared to occur earlier

in nickel (at 0.1 dpa) than in PE16 or type 316

(2 dpa), but the peak void concentration was higher

by a factor of about 10 in the austenitic steel than in

nickel or PE16

Hudson et al.9 originally attributed the swelling

resistance of PE16 to the presence of the coherent,

ordered face-centered cubic, Ni3(Al,Ti) g0

precipi-tates, which were thought either to trap vacancies and

interstitials at their surface, thereby enhancing

point-defect recombination, or to inhibit the climb of

dis-locations, thereby preventing them from acting as

preferential sinks for interstitial atoms In support of

the first of these two suggested mechanisms, Bullough

and Perrin10 argued that the surface of a coherent

precipitate would be a more effective trapping site

than an incoherent one where the identity of the point

defects would immediately be lost (and where, as a

consequence, void nucleation was likely to occur) The

efficiency of point defect trapping would be expected

to be greater the higher the total surface area of the g0

precipitates, that is, to be inversely proportional to

the precipitate size at constant volume fraction On

the other hand, the second mechanism proposed by

Hudson et al should be most effective at an

intermedi-ate particle size where dislocation pinning is strongest

Support for the latter process was provided by Williams

and Fisher11 from HVEM (high-voltage electron

microscope) irradiations of PE16 at a damage rate of

about 10
2dpa s
1at 600C, where the swelling rate

was higher at small (3 nm) and large (70 nm) g0particle

diameters than at intermediate sizes of about 20 nm

However, it is now considered that any effect

that the g0 precipitates may have on the swelling

resistance of Nimonic PE16 is secondary to that of

the matrix composition The generally low-swelling

behavior of Ni-based alloys compared to austenitic

steels was shown by Johnston et al.12following

bom-bardment with 5 MeV Ni2+ions at 625C The

dam-age rate in these experiments was 10
2dpa s
1and

the displacement dose was originally estimated as

140 dpa but this was subsequently revised by Bates

and Johnston13 to 116 dpa (based on displacement

energy Ed¼ 40 eV) In addition to hardened alloys, including PE16 and Inconel 706,this experiment included nonhardenable high-Nialloys, such as Inconel 600 and Hastelloy X, a range

precipitation-of commercial steels, and Fe–Cr–Ni ternary alloyscontaining 15% Cr and 15–35% Ni The alloys werepreimplanted with 15 appm helium prior to ion bom-bardment, and the irradiation temperature was chosen

as being close to the peak swelling temperature for irradiated austenitic steels The extent of void swellingwas determined by electron microscope examinations

ion-in low-swellion-ing alloys, but was estimated from height measurements (comparing the surfaces of irra-diated and nonirradiated regions) in high-swellingmaterials As illustrated inFigure 1, the results showednegligible swelling (<0.1%) in PE16, Inconel 706,Hastelloy X, and the Fe–15Cr–35Ni ternary alloy,low swelling (<1%) in other high-Ni alloys, but highswelling (generally>20%) in austenitic steels In com-mercial alloys containing 18% Cr, minimumswelling occurred at Ni contents of about 40–45%.Although void diameters generally appeared to besmaller in the Ni-based alloys than in austenitic steels,the main factor accounting for reduced swelling was amuch lower void concentration In the ternary alloys,reducing the Ni content from 35% to 30% resulted in

step-8 9

7 6

3

2 1

13 12 11 10 0

10

20 30 40 50 60

Nickel (wt%)

Commercial alloys Fe–15Cr–Ni alloys Commercial alloys:

an increase in overall swelling from<0.1% to 12%,

although it was noted that the 35% Ni alloy showed

a localized swelling of 5% in a region close to a

grain boundary Additional experiments reported by

Johnston et al.12indicated that the peak swelling

tem-perature for PE16 irradiated with 5 MeV Ni2+ions was

675C, but even then, swelling at 116 dpa remained

below 0.2%

Swelling data for a wider range of pure Fe–Cr–Ni

austenitic alloys, with Cr contents up to 30% and

Ni up to 100%, following Ni ion bombardment to

116 dpa at 675C, were reported by Bates and

Johnston.13 These results showed a strong

depen-dence on both Cr and Ni, with the swelling

increas-ing with increasincreas-ing levels of Cr but beincreas-ing minimized

at Ni contents of about 45–60% Examination of the

dose dependence of swelling in ternary alloys

con-taining 15% Cr and 20–45% Ni showed that the

incubation dose required for the onset of swelling

increased with increasing Ni content Furthermore,

although high-swelling rates of the order of 1% per

dpa were attained in 20–35% Ni alloys, the swelling

rate of the 45% Ni alloy remained low even at doses

above 250 dpa

Following their earlier C2+ion irradiation

experi-ments, Hudson and coworkers moved to the use of

46.5 MeV Ni6+ ions to investigate void swelling

behavior This was considered preferable because

the recoil spectra of high-energy Ni ions provided

a better simulation of fast neutron damage, and

because carbon implantation encouraged the

forma-tion of carbides which acted as void nucleaforma-tion sites

A summary of some of the Ni ion irradiation work

carried out by the Harwell group was given by Makin

et al.14 No significant differences in the swelling

behavior of Nimonic PE16 were evident between

ST or aged conditions Peak swelling in Ni6+

ion-irradiated PE16 (preimplanted with 10 appm He)

occurred at 625C, where a swelling of1.5% was

recorded at 120 dpa (N/2) Void concentrations in

PE16 were reported to be lower by a factor of about

5 than in similarly irradiated type 316 and 321

aus-tenitic steels

A drawback of charged particle irradiation

experi-ments for evaluating void swelling is that the

evolu-tion of other microstructural features may differ

significantly from that during neutron irradiation

(see also Chapter 1.07, Radiation Damage Using

Ion Beams) In the case of Nimonic PE16, for

exam-ple, the precipitation and/or redistribution of the

g0 phase during long-term neutron exposure might

be expected to influence swelling behavior In order

to simulate swelling in a more appropriate structure, Bajaj et al.15examined the effect of 4 MeV

micro-Ni2+ion irradiation on PE16, which had been conditioned by exposure to neutrons in ExperimentalBreeder Reactor-II (EBR-II) Reactor-conditionedsamples had been exposed to neutron fluences inthe range of 3–6 1026

pre-n m
2(E > 0.1 MeV) at peratures from 454 to 593C Swelling rates during

tem-Ni ion irradiations at 675C were higher by a factor

of about five in reactor-conditioned material than in anonconditioned sample The increased swelling ratewas attributed to changes in the matrix compositionresulting from an increased volume fraction of g0 inthe reactor-conditioned material

Early attempts to account for the effects of matrixcomposition on void swelling focused on the stability

of the austenite phase Harries16 suggested that theswelling behavior of austenitic steels and nickel-based alloys could be rationalized in terms of their

Ni and Cr equivalent contents (i.e., the relativeaustenite and ferrite stabilizing effects of their con-stituent elements), with the composition of high-swelling alloys then falling into the (gþ s) phasefield in the Fe–Cr–Ni ternary phase diagram Harriespostulated that the partitioning of solute elementsinto the sigma phase would have a detrimental effect

on the swelling resistance of austenite Watkin17took

a similar approach, but found that an improvedcorrelation could be obtained using the concept ofelectron vacancy numbers rather than Ni and Crequivalents The average electron vacancy number,

Nv, of the matrix is calculated from the atomic tions of its constituents, with allowance being madefor the precipitation of carbides and g0 (or g00, etc.),and has been widely used to predict the susceptibility

frac-of nickel-based alloys to the formation frac-of lic phases.18Nvwas calculated from:

intermetal-Nv¼ 0:66Ni þ 1:70Co þ 2:66Fe þ 3:66Mn

þ 4:66 Cr þ Moð ÞWatkin found that void swelling in a range of alloyswith Ni contents up to43%, which were irradiated

in DFR to a peak dose of 30 dpa at 600C, remainedlow for Nv below about 2.5 (corresponding to lowsusceptibility to s phase formation), but increasedapproximately linearly at higher Nv However, as wasclearly argued by Bates and Johnston,13correlationsbased on sigma-forming tendency could not accountfor the minimum in swelling observed at about 45%

Ni, since higher Ni contents should continue to bebeneficial

A better understanding of the swelling behavior of

Fe- and Ni-based alloys resulted from a series of fast

neutron irradiation experiments which were carried

out in EBR-II in the early 1980s Irradiation

tem-peratures in these experiments ranged from about

400 to 650C Initial data for a range of commercial

alloys, including ferritic and austenitic steels, as well

as nickel-based alloys, were reported by Bates and

Powell19and Powell et al.,20with higher dose data (up

to a peak fluence (E > 0.1 MeV) of 25  1026n m
2,

corresponding to 125 dpa) being reported by

Gelles21 and Garner and Gelles.22Swelling data for

Fe–Cr–Ni ternary alloys, irradiated in EBR-II to a

peak fluence of 22 1026

n m
2 (110 dpa), werepresented by Garner and Brager.23 The extent of

void swelling in these experiments was determined

by density change measurements In general, alloys

with nickel contents in the range of 40–50%

exhib-ited the lowest swelling Swelling in commercial

nickel-based alloys was generally lower in ST than

in aged conditions, this being attributed to the

bene-ficial (though temporary) effect of minor elements

remaining in solution and being able to interact with

point defects19; subsequent precipitation during

irra-diation would be expected to reduce this benefit and

the resulting densification, though small, would also

effectively reduce the measured swelling Swelling

data for a number of ST alloys, which were irradiated

in the AA-1 rig in EBR-II, are shown in Figure 2;

data are shown for two withdrawals, at peak fluences

of 14.7 1026

n m
2and 25.3 1026

n m
2, with surements for Inconel 600 and Inconel 625 reported

mea-at both fluence levels, dmea-ata for Nimonic PE16 and

Inconel 706 at the lower level, and data for Incoloy

800 and Hastelloy X at the higher level The nickel

contents of the alloys range from about 34% in

Inco-loy 800 to 75% in Inconel 600 Swelling remained

relatively low in the three Inconel alloys and in PE16

However, both Incoloy 800 and Hastelloy X exhibited

high swelling at some temperatures, with swelling in

the latter reaching80% at 593C The reason for

such high swelling in neutron-irradiated Hastelloy

X (nickel content48%) is unclear, but it was noted

that densification up to 3% occurred at the lower

irradiation temperatures – indicating microstructural

instability and possibly signaling changes in the

com-position of the matrix which may have affected the

swelling behavior (Note that Hastelloy X was

identi-fied as a low-swelling alloy in the Ni2+ion irradiation

experiments described by Johnston et al.12)

Some data for different heat-treated conditions of

PE16 at the higher fluence level were reported by

Garner and Gelles,22and are compared for tions at 538C (more or less corresponding to thepeak swelling temperature for PE16 in the AA-1experiment) with lower fluence data from Batesand Powell19 in Figure 3 The heat-treated condi-tions indicated in Figure 3 are ST (ST 4 h at

irradia-1080C), A1 (ST and aged 16 h at 705C), A2 (STand aged 1 h at 890C plus 8 h at 750C), and OA(ST and aged 24 h at 840C) Note that the siliconcontent of the PE16 used in these experiments wasmuch lower at 0.01% than the level of0.2% typi-cally found in UK heats of the alloy Overall, the dataappear to show little effect of initial heat treatment

on the swelling of PE16, except that the OA tion exhibited the most swelling (5.2%) at the higherfluence

condi-Although it is clear that the swelling behavior ofaustenitic alloys is largely dependent on nickel con-tent, there is ample evidence to show that minor soluteadditions can have significant effects Much of thework on minor solutes has focused on steels similar

to type 316, but some data are available for highernickel alloys For example, Mazey and Hanks24 used46.5 MeV Ni6+ion irradiations to examine the effects

of Si, Ti, and Al additions on the swelling response ofmodel alloys with base compositions approximatingthat of the matrix phase in PE16 Solute additions

of 0.25% Si or 1.2% Ti reduced swelling, but theaddition of 1.2% Al (in the absence of Si or Ti)markedly increased it The beneficial effect of Si wasbelieved to arise from its high diffusivity in solution(this is discussed further inSection 4.04.2.2), whereasthat of Ti appeared to be related to the formation of Zphase (hexagonal-structured Ni3Ti) The addition of

Al resulted in an increase in the concentration ofvoids, the surfaces of which were coated in a thinlayer of the g0phase (Ni3Al) A beneficial effect of Si

on the swelling response of modified Incoloy DS alloysunder Ni6+ion irradiation was also reported by Mazey

et al.25However, it should be noted that high Si tents can give rise to the formation of radiation-induced phases which are enriched with Ni and Si,such as the Ni3Si form of g0and the silicide G-phase(M6Ni16Si7, where M is usually Ti, Nb, or Mn).G-phase particles are generally found in associationwith large voids and their formation may thereforegive rise to an increase in the swelling rate.26,27Swelling data derived from density measurementsfor neutron irradiated, modified Incoloy DS alloys,with Si contents ranging from 0.19 to 2.05% (com-pared to a specified level of 1.9–2.6% in the commer-cial alloy), are compared with data for a ‘PE16 matrix

con-alloy’ and Nimonic PE16 inFigure 4 The materials

were all in ST condition apart from PE16 which was

in an STA condition (aged 4 h at 750C) The alloys

were irradiated in the UK-1 rig in EBR-II to fluences

in the range of 9–16 1026

n m
2 (E > 0.1 MeV) attemperatures of 390–640C These data are pre-

viously unpublished except those for STA PE16

(heat DAA 766) which were reported by Boothby.28

Swelling in the modified Incoloy DS alloys generally

decreased with increasing Si content The 0.19% Si

alloy exhibited high swelling at all temperatures withindications of swelling peaks at about 440 and 640C.Increased Si levels tended to suppress the high tem-perature swelling peak and reduce the magnitude ofswelling at lower temperatures The PE16 matrixalloy containing 0.24% Si exhibited a high tempera-ture swelling peak but moderate swelling below

550C, suggesting a beneficial effect of Mo (thisbeing the main compositional difference between thePE16 matrix alloy and the modified Incoloy DS alloys)

−0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0

Temperature (°C)

0 5 10 15

Temperature (°C)

0 5 10 15 20

25

In 600

In 625

In 800 Hast X

at lower temperatures However, swelling in the PE16matrix alloy remained significantly higher at all tem-peratures than in STA Nimonic PE16 (containing0.15% Si), indicating a significant benefit of the g0forming elements Al and Ti.

4.04.2.2 Void-Swelling ModelsPoint defects created by atomic displacements arelost either through mutual recombination or bymigration to sinks Void swelling requires a mobilepopulation of excess vacancies and can only occurover a limited temperature range, typically 350–

700C in neutron-irradiated steels and nickel-basedalloys Rapid diffusion at higher temperatures reducesthe concentration of radiation-induced vacancies tonear thermal equilibrium levels Recombination dom-inates at lower temperatures, where reduced vacancymobility prevents the formation of voids as the neces-sary counter-migration of matrix atoms cannot occur

In the swelling regime, an increased bias for tials over vacancies at dislocation sinks gives rise to thesurplus vacancies which agglomerate to form voids.The flux of point defects to sinks, including voidsurfaces, dislocations, and grain boundaries, results inthe segregation of particular solute atoms at the sinksand the depletion of others In austenitic steels andnickel-based alloys, it is generally found that nickelsegregates at the point defect sinks This is generallyattributed to the inverse Kirkendall effect described

intersti-by Marwick,29 whereby faster diffusing solutes such

as Cr move in the opposite direction to the vacancyflux and are depleted at the sink, and slower diffusingsolutes such as Ni are enriched One of the earliestobservations of nickel segregation at void surfacesdue to the inverse Kirkendall effect was made byMarwick et al.30 in an alloy with a composition cor-responding to that of the matrix phase in NimonicPE16 (For more detailed discussions on radiation-induced segregation effects, see the reviews ofWiedersich and Lam,31and Rehn and Okamoto.32)Venker and Ehrlich33recognized that differences

in the partial diffusion coefficients of alloy ents might account for the effects of composition onswelling Any effect of this kind would generally beexpected to be more significant the larger are thedifferences between the partial diffusion coefficients

constitu-of the alloy components Garner and Wolfer34ined Venker and Ehrlich’s conjecture and concludedthat the addition of even small amounts of a fast-diffusing solute such as silicon to austenitic alloyswould greatly increase the effective vacancy diffusion

Figure 4 Void swelling data derived from density

measurements for Nimonic PE16, a PE16 matrix alloy,

and modified Incoloy DS alloys, irradiated in the UK-1 rig

in Experimental Breeder Reactor-II Unpublished data from

Boothby, R M.; Cattle, G C Void Swelling in EBR-2

Irradiated Nimonic PE16 and Incoloy DS; FPSG/P(90)10,

with permission from AEA Technology Plc.

Figure 3 Effect of heat treatment on void swelling of

Nimonic PE16 irradiated in Experimental Breeder Reactor-II

at 538C Adapted from Bates, J F.; Powell, R W J Nucl.

Mater.1981 , 102, 200–213; Garner, F A.; Gelles, D S.

In Effects of Radiation on Materials: 14th International

Symposium; Packan, N H., Stoller, R E., Kumar, A S.,

Eds.; American Society for Testing and Materials:

Philadelphia, PA, 1990; Vol II, pp 673–683, ASTM

STP 1046.

coefficient (i.e., would enhance the diffusion rate for

all matrix elements) The overall effect is analogous to

an increase in temperature – resulting in an effective

decrease in the vacancy supersaturation and hence a

reduction in the void nucleation rate This

mecha-nism is generally accepted as the explanation for the

beneficial effect of silicon in reducing swelling in

austenitic steels and nickel-based alloys Although

this relies on the diffusion of silicon via vacancy

ex-change, silicon is also generally observed to segregate

to point defect sinks and since it is an undersized

solute, this is believed to occur by the migration of

interstitial–solute complexes There is, however, no

reason to suppose that both diffusion mechanisms

cannot operate simultaneously

Garner and Wolfer34 originally considered that

since nickel diffuses relatively slowly in austenitic

alloys, an increase in nickel content would have the

opposite effect to silicon However, a later assessment

made by Esmailzadeh and Kumar,35 based on

diffu-sion data reported by Rothman et al.,36indicated that

the void nucleation rate in Fe–15Cr–Ni alloys would

decrease with an increase in nickel content from 20 to

45% This result is obtained because, although nickel

remains the slowest diffusing species, the effective

vacancy diffusion coefficient of the system is

calcu-lated to increase at the higher nickel content

Esmail-zadeh and Kumar’s calculations also confirmed

the beneficial effect of silicon, with the addition of

1% Si predicted to be as effective in suppressing

void nucleation as increasing the nickel content

from 20 to 45% Effects at nickel contents above

45% could not be examined due to a lack of

appro-priate diffusion data

As well as affecting the nucleation of voids,

differ-ences in the diffusion rates of the various solutes

might also be expected to influence void growth

Simplistically, this can be thought of as being partly

due to the segregation of slower diffusing solutes

reducing the rate of vacancy migration in the vicinity

of the voids However, a further consequence of such

nonequilibrium solute segregation was identified by

Marwick,29 who realized that it would give rise to

an additional vacancy flux which would oppose the

radiation-induced flux to the sink As discussed by

Marwick, this additional flux (the Kirkendall flux)

may itself be an important factor in limiting void

growth, since it will reduce the probability of vacancy

annihilation at sinks and increase the likelihood of

point defect recombination

The effect of nickel content on void swelling was

considered further in a model developed by Wolfer

and coworkers.37,38The model examined the sitional dependence of the void bias and focused onthe effects of nickel segregation at void surfaces.Wolfer’s model indicated that the compositional gra-dients produced by radiation-induced segregationgive rise to additional drift forces which affect thepoint defect fluxes and thereby modify the bias terms.These additional drift forces arise from the effects ofcomposition on point defect formation and migrationenergies, on the lattice parameter and the elastic mod-uli, and from the Kirkendall flux Wolfer’s calculationsfor binary Fe–Ni alloys indicated that the effect of theKirkendall flux is small for interstitials but significantfor vacancies Nevertheless, it was considered that theoverall effect of compositional gradients on the biasterms is likely to be greater for interstitials than forvacancies due to other factors, particularly the effect ofvariations in the elastic moduli As noted by Garnerand Wolfer,39an increase in the shear modulus in thesegregated regions around voids would reduce the biasfor interstitials and therefore help to stabilize voids It

compo-is difficult to predict the significance of thcompo-is effect incomplex alloys, however, since depletion of Cr in thesegregated region will tend to reduce the shear modu-lus, whereas enrichment of Ni in high-Ni alloys willtend to increase it.38 A more significant result of themodel with regard to the effect of nickel on swelling isthat there is a reversal in the sign of the Kirkendallforce for vacancies in Fe–Ni alloys at35% Ni Belowthis level, vacancies are predicted to be attracted intoregions of higher Ni concentration, but above it,the opposite occurs Wolfer et al considered that thiseffect may account for the dependence of swelling on

Ni content in austenitic alloys containing less than35% Ni

A generalized description of the swelling behavior

of austenitic alloys, which was consistent with themodel developed by Wolfer et al., was put forward

by Garner40 (see also Chapter 4.02, RadiationDamage in Austenitic Steels) Garner’s ideas werelargely based on the results of the EBR-II irradiationexperiments and the earlier ion bombardment work

of Johnston et al., both of which showed a strongdependence of swelling on nickel content It wasconsidered that swelling was characterized by a tran-sient period followed by a regime in which theswelling rate became constant In neutron-irradiatedalloys, the swelling rate in the posttransient regimewas generally found to be1% per dpa In swelling-resistant alloys, however, it was argued that such highswelling rates might not be observed owing to ex-tended transient periods The duration of the

transient regime was shown to be dependent on alloy

composition and could extend for many tens of dpa in

low-swelling materials The duration of the transient

regime was implicitly linked to the completion of

void nucleation but, at the time these ideas were

put forward, relatively few measurements of void

concentrations were available, as swelling data were

mainly derived from dimensional or density changes

Factors that were proposed to account for the

influence of nickel on the void nucleation rate

in-cluded the effect on vacancy diffusivity described by

Esmailzadeh and Kumar35; a possible correlation with

the development of fine scale compositional

fluctua-tions by a spinodal-like decomposition process

(observed by Dodd et al.41 in ion-irradiated ternary

Fe–Cr–Ni alloys); and an effect of nickel on the

minimum critical radius for the formation of stable

voids.42Voids are unstable below a critical size, and

will generally shrink unless stabilized by gas atoms;

the minimum stable void radius is dependent on a

number of factors, including temperature and defect

bias, and Coghlan and Garner suggested that the

compositional dependence of the vacancy diffusivity

would also affect this critical size In other words, it

was considered that the transition from gas bubble to

void would require a larger bubble size in

high-nickel alloys, particularly at relatively high

tempera-tures in the swelling regime where void nucleation

becomes increasingly difficult Hoyt and Garner43

subsequently argued that the minimum critical void

radius concept might account for the minimum in

swelling found at the intermediate nickel contents,

provided that a compositional-dependent bias factor

for dislocations was also incorporated into the model

The compositional dependence of the bias factor

arises from solute segregation, which reduces the

strain energy of dislocations and decreases the ratio

of the bias for interstitials compared to vacancies

It is of interest that early evidence for the

opera-tion of the bubble to void transiopera-tion was obtained by

Mazey and Nelson,44who implanted Nimonic PE16

(STA condition) and a PE16 matrix alloy (ST

condi-tion) with 1000 appm He to produce a high density

of gas bubbles before subsequent irradiation with

46.5 MeV Ni6+ ions The PE16 matrix alloy used

in this particular experiment was a low Si variant

(<0.02 wt%) which was known to exhibit relatively

high swelling The mean bubble size following

helium implantation at 625C was higher by a factor

of about two in the matrix alloy (11 nm diameter)

than in the commercial PE16 alloy (5 nm diameter)

Examination of the alloys following subsequent

irradiation also at 625C revealed high swelling(12% at 60 dpa) with a uniform distribution of largevoids but no remaining helium bubbles in the matrixalloy, and low swelling (1% at 60 dpa) with abimodal distribution of bubbles plus voids in thestandard PE16 alloy (see Figure 5) These resultswere interpreted as providing evidence for the con-cept of a critical stable void size, with only a smallfraction of bubbles in the commercial PE16 alloy,but all of the bubbles in the matrix alloy, beingsufficiently large to grow as voids Although notspecifically discussed by Mazey and Nelson, thecompositional differences between the two alloyssuggest that the presence of Si and/or the g0formingsolutes Al plus Ti may help to reduce void nucleation

in PE16

The belief advanced by Garner,40that sluggish voidnucleation generally accounted for low swelling innickel-based alloys, persisted for some time However,data reported by Muroga et al.45,46largely overturnedthis view Muroga et al carried out microstructuralexaminations of a series of EBR-II-irradiatedFe–15Cr–Ni ternary alloys with Ni contents rangingfrom 15 to 75 wt%, and of archived samples of similaralloys from the heavy-ion bombardment experiments

of Johnston et al.12Examination of alloys irradiated inEBR-II at 510C showed that the saturation void con-centration was dependent on nickel content and wasminimized at35–45% Ni, but revealed that there was

no increase in void numbers in any of the materialsabove a fluence of 2.6 1026

n m
2(E > 0.1 MeV) (seeFigure 6) Alloys containing 19% and 30% Ni exhib-ited high swelling rates at higher fluences, but swellingremained low in higher nickel alloys Similar effectswere found in the ion-bombarded samples, where, forexample, it was shown that there was no significantchange in the void concentration in Fe–15Cr–45Ni

at doses above 50 dpa in irradiations at 675C, yet amarked increase in swelling rate occurred above

120 dpa Thus, contrary to earlier ideas, these gations clearly demonstrated that the onset of a highswelling rate was not related to the cessation of voidnucleation It follows that the transition to a high rate ofswelling must be due to an increase in the growth rate

investi-of existing voids

Muroga et al.45,46observed that the total tion density in the irradiated Fe–15Cr–Ni alloys wasonly weakly dependent on nickel content This sug-gested that at the intermediate nickel levels, wherethe void concentration was low, dislocations wereweak sinks (for both vacancies and interstitials)relative to voids In addition, it was observed that

disloca-dislocation loops persisted to higher doses at the

intermediate nickel contents, indicating a lower

growth rate for the loops – again implying an effect

of nickel on dislocation sink strength Based on these

observations, Muroga et al suggested that a reduced

dislocation bias for interstitials at the

intermedi-ate nickel contents might explain the influence of

nickel on the early stages of void development An

additional factor was required to account for the

eventual transition to a high swelling rate

Microche-mical data presented by Muroga et al.46 suggested

that this transition was related to the depletion of

nickel in the matrix owing to its enrichment at void

surfaces

A complete description which incorporates all of

the composition-dependent factors which affect the

nucleation and growth of voids is lacking However,

there is a general consensus that the major influence

of alloy composition arises through its effects on the

effective vacancy diffusivity and on segregation

aris-ing from the inverse Kirkendall effect A correlation

between the magnitude of void swelling and

radia-tion-induced segregation was shown for Fe–Cr–Ni

ternary alloys by Allen et al.48 The compositionaldependence of radiation-induced segregation wasdetermined using a model based on the earlier work

of Marwick,29 which incorporates both the vacancyflux to the voids and the back-diffusion of vacanciesdue to the solute gradients set up by the inverseKirkendall effect Vacancy diffusivities for variousalloy compositions were determined by the measure-ments of grain boundary segregation in proton-irradiated samples Swelling data for ion andneutron-irradiated alloys were then compared withthe expected swelling propensity defined by the ratio

of the forward to back diffusion terms calculated at theappropriate irradiation temperature The materialsfor which vacancy diffusivity data were determinedincluded Fe-based alloys containing 16–24% Cr and9–24% Ni, and Ni-based alloys containing 18% Crand either zero or 9% Fe This work did not specifi-cally examine 40–50% Ni alloys corresponding to thehighest swelling resistance, though the results indi-cated that swelling generally decreased with increas-ing levels of nickel enrichment and chromiumdepletion at void surfaces

20 16 12 8 4 0

0

0 10 20 30 40

Cavity diameter, d (nm)

50 60 70 80 90 100 110 120 130 4

8 12 16 0 4 8 12 16

Figure 5 Histograms showing size distributions of bubbles/voids in (a) solution treated and aged Nimonic PE16 and (b) solution treated PE16 matrix alloy, irradiated with Ni6þions at 625C to damage levels of 30 and 60 dpa following implantation with 1000 appm He (producing 0.2 dpa) at the same temperature Reproduced from Mazey, D J.;

Nelson, R S J Nucl Mater.1979, 85–86, 671–675.

4.04.2.3 Swelling Behavior of

Neutron-Irradiated Nimonic PE16

Brown et al.49 compared the swelling behavior of

STA Nimonic PE16 and two cold-worked austenitic

steels (M316 and Nb-stabilized FV548) which were

irradiated in DFR as fuel pin cladding Two PE16

clad pins were examined, which were irradiated to

burn-ups of 6.1% and 21.6% of heavy atoms,

corresponding to peak damage levels of about 17

and 80 dpa, respectively Void concentrations and

swelling were lower in PE16 than in the austenitic

steels Swelling data, void concentrations, and void

diameters for the two PE16 pins examined by Brown

et al are shown inFigure 7 Note that Brown et al.49

only showed trend lines for void concentration and

void size in the less highly irradiated pin and

com-pared the swelling tendencies of the two pins; the

individual data points were not plotted and those

shown inFigure 7are previously unpublished data

obtained by Sharpe Brown et al stated that the void

concentration in PE16 decreased with increasing

irradiation temperature but did not alter greatly

with an increasing dose above17 dpa It should be

noted, however, that swelling measurements for the

higher burn-up pin were restricted to temperatures

below 525C, so that a direct comparison of voidconcentrations in the two pins cannot be made athigher temperatures Although there were fewervoids in PE16 than in the two steels, the voids

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Temperature (°C)

DFR ~ 17 dpa DFR ~ 80 dpa

DFR ~ 17 dpa DFR ~ 80 dpa

0 10 20 30 40 50 60 70 80 90 100

in Fast Reactor Irradiated Commercial High Nickel Alloy; DFMC/P(82)27, with permission from AEA Technology Plc.

Figure 6 Fluence dependence of swelling and void

density of Fe–15Cr–Ni alloys irradiated in Experimental

Breeder Reactor-II at 510C Swelling data obtained by

immersion density measurements by Garner and Kumar47

are also shown Reproduced from Muroga, T.; Garner, F A.;

Ohnuki, S J Nucl Mater.1991, 179–181, 546–549.

appeared to be homogeneously distributed and to have

developed during the early stages of irradiation; once

nucleated, the growth rate of voids in PE16 remained

low These observations are clearly contrary to early

models which suggested that low swelling rates result

from incomplete void nucleation and extended

tran-sient regimes Rather, in agreement with the more

recent observations of Muroga et al.,45,46 it appears

that the swelling resistance of PE16 is due to a

combi-nation of a comparatively low saturation void

concen-tration, which is reached at a relatively low

displacement dose, and a low void growth rate There

does not appear to be any evidence of an accelerated

swelling rate in PE16 once void nucleation is complete

Additional data on void concentrations in

neutron-irradiated PE16 are available from Cawthorne et al.,8

Sklad et al.,50 and Boothby.28 The results presented

by Cawthorne et al for PE16 fuel pin cladding

irra-diated in DFR to a peak fluence of 5.6 1026

n m
2(28 dpa) differ from those shown in Figure 7 in

that, although void number densities are similar for

irradiations at 380–520C, void concentrations are

about an order of magnitude higher at 350C and

600–630C Such discrepancies might arise from

uncertainty and/or variability in irradiation

tempera-tures Another possibility is that void nucleation was

incomplete at the higher irradiation temperatures in

the lower burn-up pin examined by Brown et al Data

from Sklad et al show an increase in void numbers in

unstressed PE16 specimens irradiated in EBR-II at

500C from an average (for two differently heat

trea-ted conditions) of about 4 1019

to 1.2 1020

m
3with increasing fluence from 1.2 1026

to4.0 1026

n m
2 (E > 0.1 MeV), that is, from 6 to

20 dpa In this case, the void concentration and overall

swelling of0.2% at 20 dpa remain below the levels

shown in Figure 7 for the DFR-irradiated pin

at 17 dpa; this may reflect the effect of stress on

swelling for fuel pin cladding

Void swelling data determined from Transmission

electron microscope (TEM) examinations of three

heats of PE16 which were irradiated in the UK-1

rig in EBR-II are shown inFigure 8, which includes

previously unpublished results for the low boron

(4 ppm) heat Z184 as well as data for heats DAA766

and Z260D (with 18 and 70 ppm boron, respectively)

which were reported by Boothby.28 Data are shown

for all three heats in the STA condition (ST 1020C

and aged 4 h at 750C) and for DAA766 in the OA

condition (a multistage heat treatment that included

aging at 900C, slow cooling to 750C, and then

aging for 16 h at that temperature, resulting in the

precipitation of TiC and an overaged g0 structure).Swelling data derived from the density measure-ments of STA PE16 heat DAA 766 from the sameexperiment are shown inFigure 4 An example of the

0.0 2.0 4.0 6.0 8.0 10.0

Temperature (°C)

0 5 10 15

STA DAA766

OA DAA766 STA Z260D STA Z184

STA DAA766

OA DAA766 STA Z260D STA Z184

in Experimental Breeder Reactor-II Adapted from Boothby,

R M J Nucl Mater.1996, 230, 148–157; Unpublished data for Boothby, R M The Microstructure of EBR-II Irradiated Nimonic PE16; AEA TRS 2002 (FPSG/P(90)23), with permission from AEA Technology Plc.

void distribution in the OA condition is shown in

Figure 9 Note that the voids in neutron-irradiated

PE16 tend to be cuboidal and that enhanced growth

of voids attached to TiC precipitates (located at the

site of a prior grain boundary) has occurred

Neutron fluences and irradiation temperatures

in the UK-1 experiment were similar to those for the

first withdrawal of the AA-1 rig for which data is

shown in Figure 2 Void concentrations for heats

DAA766 and Z260D shown inFigure 8appear to be

less temperature-dependent than for the fuel pin

clad-ding data shown inFigure 7 Void numbers are

gener-ally lower than in the cladding at temperatures

up to 550C, but are intermediate between the

results of Brown et al.49and Cawthorne et al.8for

irra-diations at 600C Void concentrations for PE16

irradiated to fast neutron fluences (E > 0.1 MeV) of

9.4–12.3 1026

n m
2at 477–513C in the UK-1

ex-periment were very similar to those determined by

Sklad et al.50 for 4.0 1026

n m
2at 500C The lowboron heat Z184 showed atypical behavior, with a very

high concentration of small voids and low swelling at

438C, but high swelling owing to increased void sizes

at normal void concentrations at temperatures above

513C It is probable that the effect of boron on swelling

is related to the formation of boron–vacancy complexes,

which can give rise to the nonequilibrium segregation

of boron in the presence of quenched-in thermal

vacan-cies as well as to radiation-induced effects.51

Some variability in the swelling response of

Nimonic PE16 in PFR (Prototype Fast Reactor)

components was reported by Brown and Linekar.52

Increased swelling in PE16 subassembly and guidetube wrappers in PFR compared to expectationsbased on the performance of DFR pin claddingappeared to be related to temperature fluctuations,particularly at temperatures below 400C during theearly operation of PFR Void concentrations werereported to be higher in the PFR components, and

it was suggested (by Cawthorne, unpublished data)that this may have been due to the release of vacan-cies from vacancy loops which had formed duringlower temperature excursions In fact, the void con-centration reported by Cawthorne et al.8for DFR pincladding irradiated at 350C was higher than thehighest value reported for the PFR components by

a factor of about 3, but this comparison was not made

by Brown and Linekar There were also indications ofheat-to-heat variability and effects of the fabricationroute on the swelling of PE16 wrappers in PFR.Nevertheless, swelling of PE16 wrappers, althoughhigher than expected, remained low in absolute termsand did not give rise to any operational problems.Although PE16 was originally selected as the ref-erence wrapper material for PFR and as an alterna-tive to cold-worked M316 steel for fuel pin cladding,PE16 was favored as a cladding material with 12%Crferritic–martensitic steel wrappers in subsequentsubassembly designs.53The 12%Cr steel was chosen

as a wrapper material because of its superior swellingresistance, but its use was limited to relatively lowtemperatures owing to inadequate strength at thehigher operating temperatures experienced by pincladding Design calculations for PE16 fuel pin clad-ding made by Cole54 indicated that cladding hoopstresses, which arise from the internal pressure fromthe gaseous fission products released from the fuel,were much lower than the yield stress of the materialand were generally expected to remain below about

70 MPa In addition, the void swelling and irradiationcreep behavior of PE16 were considered to be wellmatched to the fuel swelling, so that fuel–clad inter-action stresses also remain low Fuel pins with PE16cladding successfully attained high burn-ups in PFR,with some 3500 pins exceeding dose levels of 100 dpaand 265 pins reaching maximum doses of 155 dpa.55Very few failures of PE16 clad pins were recorded –three failures occurred in pins which had reachedburn-ups over 17 at.%, with one failure at 11.3 at.%burn-up which was believed to have resulted from

a fabrication defect.56 In addition to the four PE16cladding failures in PFR, Plitz et al.57 recorded 14failures in austenitic steel cladding, all at lower burn-ups than in PE16 The failures in PE16 cladding were

200 nm Figure 9 Void structure in PE16 (OA condition)

irradiated in Experimental Breeder Reactor-II to 58 dpa at

513C Reproduced from Boothby, R M J Nucl.

Mater.1996, 230, 148–157.

regarded as benign and permitted continued

opera-tion, with no significant loss of fuel into the primary

circuit coolant A peak burn-up of 23.2 at.%,

corresponding to a peak dose in the PE16 cladding

of 144 dpa, was achieved in PFR in an experimental

fuel cluster Postirradiation examinations of pins

from this cluster and a high burn-up subassembly

(18.9 at.%, with a peak cladding dose of 148 dpa)

were carried out by Naganuma et al.58 Maximum

diametral strains of less than 1% were measured,

attributable to the combined effects of void swelling,

creep deformation arising from internal gas pressure

in the pins, and small contributions from mechanical

interactions between the fuel and cladding in the

lower part of the pins

4.04.3 Irradiation Creep

A detailed discussion of irradiation creep

mechan-isms is beyond the scope of this chapter, which will

instead concentrate on experimental data which

enable comparisons to be made between

nickel-based alloys and austenitic steels However, some

insight into irradiation creep mechanisms is given

inSection 4.04.4.1, where the effect of stress on the

evolution of dislocation structures is described

Irra-diation creep mechanisms are discussed more fully

inChapter 1.04, Effect of Radiation on Strength

and Ductility of Metals and Alloys Several reviews

of irradiation creep data are available in the literature,

for example, by Harries,59 Ehrlich,60 and Garner,61

and although these have tended to focus on austenitic

steels, the behavior of nickel-based alloys generally

appears to be similar

Different types of test specimen, including

pres-surized tubes and helical springs, have been used to

measure irradiation creep strains The data are

therefore generally converted to effective straine

and effective stress s values, using the Soderberg

formalism60:

e=s ¼ e=s ¼ g=3t ¼ 4eH=3sH

where e, g, and eH are tensile, surface shear and

hoop strains; and s, t, and sH are tensile, surface

shear and hoop stresses, respectively

Irradiation creep experiments carried out in DFR

used helical spring specimens, which were loaded in

tension and periodically removed for measurements

DFR data for austenitic steels and Nimonic PE16

were reviewed by Mosedale et al.62 and Harries,59

and results for PE16 were reported in full by

Lewthwaite and Mosedale.63 Average irradiationtemperatures for PE16 specimens ranged fromabout 280 to 340C, with displacement doses up to

a maximum of 13 dpa (N/2) For austenitic steels,the irradiation creep strain was found to be linearlydependent on the applied stress and the displacementdose, comprising transient and steady-state compo-nents as follows:

g ¼ At þ Bdtwhere d is the displacement dose and A and B arematerial-dependent creep coefficients For PE16 in

a STA condition (1 h at 1080C plus 16 h at 700C),creep at dose rates of 5  10
7dpa (N/2) s
1wascharacterized by an initial period of low strain and anincreased creep rate at higher displacement doses.Mosedale et al.62described the g=t versus dpa creepcurve for STA PE16 as parabolic, though the maxi-mum observed creep rate was similar to that in aus-tenitic steels and Harries59 represented the creepstrain above a threshold dose of 8 dpa (N/2) by

g ¼ 4:3  10
6tðd
8Þwhere t is in MPa; converting to effective strain/stress values and to NRT units of displacement dose(assuming 1 dpa (N/2)¼ 0.8 dpa (NRT-Fe)) wouldreduce the creep coefficient by a factor of 2.4 Datapresented by Lewthwaite and Mosedale63 showedthat ST PE16 behaved similarly to the STA condi-tion, though OA conditions exhibited higher creepstrains due to a combination of increased creep ratesand low threshold doses (around 1 dpa) An apparentdose-rate dependency was observed, with steady-state creep coefficients for STA and OA PE16increased by factors of 2 at lower damage rates of

0.5–1.5  10
7dpa (N/2) s
1 and threshold dosesreduced to 0.5 dpa or less A similar effect ofdose rate on the creep strain per dpa was alsoreported for austenitic steels.64 Steady-state creepcoefficients (MPa
1dpa
1) and creep strain rates(MPa
1s
1) for PE16 as a function of dose rate arecompared with data for cold-worked steels M316 andFV548 inFigure 10 The data plotted inFigure 10are derived from the results of Lewthwaite and Mose-dale63,64 but are converted to effective strain/stressvalues and NRT(Fe) dpa units to enable comparisonwith other published data It is evident that the irra-diation creep behavior of STA and OA (24 h at

800C) PE16 is similar to that of the austenitic steels.Creep rates at higher dose rates are generally lowerthan would be indicated from the linear extrapolation

of low dose rate data Lewthwaite and Mosedale63