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 ,

4.04 Radiation Effects in Nickel-Based Alloys R M Boothby National Nuclear Laboratory, Harwell, Oxfordshire, UK ß 2012 Elsevier Ltd All rights reserved 4.04.1 Introduction 123 4.04.2 4.04.2.1 4.04.2.2 4.04.2.3 4.04.3 4.04.4 4.04.4.1 4.04.4.2 4.04.5 4.04.5.1 4.04.5.2 4.04.6 References Void Swelling Compositional Dependence of Void Swelling Void-Swelling Models Swelling Behavior of Neutron-Irradiated Nimonic PE16 Irradiation Creep Microstructural Stability Dislocation Structures Precipitate Stability Irradiation Embrittlement Fast Neutron Irradiation Experiments Helium Implantation Experiments Concluding Remarks 124 124 129 133 136 138 138 139 140 140 145 147 148 Abbreviations AGR DFR EBR-II EDX HVEM N/2 NRT OA PFR PS SIPA ST STA TEM UTS VEC Advanced gas-cooled reactor Dounreay Fast Reactor Experimental Breeder Reactor-II Energy dispersive X-ray High-voltage electron microscope dpa calculated according to half-Nelson model dpa calculated according to Norget, Robinson, and Torrens model Overaged Prototype Fast Reactor Proof stress Stress-induced preferred absorption Solution treated Solution treated and aged Transmission electron microscope Ultimate tensile strength Variable energy cyclotron 4.04.1 Introduction Research into the effects of irradiation on nickelbased alloys peaked during the fast reactor development programs carried out in the 1970s and 1980s Interest in these materials focused on their high resistance to radiation-induced void swelling compared to austenitic steels, though a perceived susceptibility to irradiation embrittlement limited their application to some extent Nevertheless, the Nimonic alloy PE16 was successfully used for fuel element cladding and subassembly wrappers in the United Kingdom, and Inconel 706 was utilized for cladding in France Both of these materials are precipitation hardened and consequently have high creep strength, and much research and development of alternative alloys was directed toward maintaining swelling resistance and creep strength while aiming to alleviate, or at least understand, irradiation embrittlement effects There has been some revival of interest in nickel-based alloys for nuclear applications, and various aspects of radiation damage in such materials have recently been reviewed by Rowcliffe et al.1 in the context of Generation IV reactors, and by Angeliu et al.2 in consideration of their use for the pressure vessel of the Prometheus space reactor Nickel-based alloys are also candidate structural materials for molten salt reactors, for which resistance to corrosion by molten fluoride salts and high-temperature creep strength are prime requirements, though intergranular attack by the fission product tellurium and irradiation embrittlement due to helium production are potentially limiting factors for this application.3 This chapter focuses on the void swelling behavior, irradiation creep, microstructural stability, and irradiation embrittlement of precipitation-hardened nickel-based alloys Fundamental to all of these effects are the basic processes of damage production 123 124 Radiation Effects in Nickel-Based Alloys 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 discussions 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; Chapter 1.11, Primary Radiation Damage Formation; Chapter 1.12, Atomic-Level Level Dislocation Dynamics in Irradiated Metals and Chapter 1.18, Radiation-Induced Segregation Typical compositions of nickel-based alloys and some precipitation-hardened steels, which are considered in this chapter, are shown in Table 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 treated), STA (solution treated and aged), and OA (overaged) 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 displacement doses (dpa) are generally given in NRT (Fe) units, although the half-Nelson (N/2) model was Table widely used particularly in the United Kingdom in the 1970s6 The exact relationship between these units will vary depending on the neutron spectrum (which may differ, not only from one reactor to another, but also depending on location within a reactor), but approximate conversion factors for fast reactor core irradiations are 1026 n m2 E > 0:1MeVị ẳ 5dpa NRTFeị ẳ 6:25dpa N=2ị 4.04.2 Void Swelling 4.04.2.1 Compositional Dependence of Void Swelling Nimonic PE16 was first identified as a low-swelling alloy in the early 1970s Void swelling data derived from density measurements on fuel pin cladding materials from the Dounreay Fast Reactor (DFR) were reported by Bramman et al.7 and were complemented by electron microscope examinations described by Cawthorne et al.8 Swelling in STA PE16 was found to be lower than in heat-treated austenitic steels and comparable to cold-worked steels Comparison of data for PE16 and FV548 (a Nb-stabilized austenitic steel) irradiated under identical conditions in DFR to a peak neutron fluence of $6  1026 n mÀ2 indicated that the lower swelling of PE16 was due to smaller void concentrations at irradiation temperatures up to $550  C and reduced void sizes at higher Nominal compositions (wt%) of commercial and developmental nickel-based alloys Alloy Ni Cr Mo Ti Al Nb Mn Si C Nimonic PE16 Inconel 706 Inconel 718 Inconel 600 Inconel 625 Incoloy 800 Hastelloy X D21a D25a D66a D66b D68a D68b PE16 matrix Incoloy DS Alloy 7817 Alloy 7818 43 41.5 52.5 75 61 34 48 25 30 45 40 45 34 36 39 40 40 16.5 16 19 16 22 20.5 21 8.4 10.5 12 11 12 12.5 20 18 15 15 1.1 – 3.0 – 9.0 – 9.0 1.0 3.7 3.0 2.0 – – 4.0 – 3.2 3.0 1.2 1.8 0.9 0.3 0.3 0.4 – 3.3 1.8 2.5 3.0 1.8 1.6 – 0.04 2.0 0.3 1.2 0.2 0.5 0.2 0.3 0.4 – 1.7 1.3 2.5 1.5 0.4 0.25 – 0.02 0.9 – – 2.9 5.2 – 3.5 – – – – – – 3.6 2.8 – – – 3.0 0.1 0.2 0.2 0.2 0.2 0.9 0.5 1.0 1.0 – 0.2 0.3 0.2 0.1 1.0 0.2 0.2 0.2 0.2 0.2 0.2 0.2 0.5 0.5 1.0 1.0 0.5 0.5 0.4 0.4 0.2 2.0 0.5 0.5 0.05 0.03 0.04 0.08 0.05 0.07 0.10 0.04 0.04 0.03 0.04 0.03 0.02 0.07 0.08 0.02 0.02 a Composition indicated by Yang et al.4 Composition indicated by Toloczko et al.5 b Other 0.3Cu 0.5Cu 0.5W, 2.0Co Fe Bal Bal Bal Bal Bal Bal Bal Bal Bal Bal Bal Bal Bal Bal Bal Bal Bal Radiation Effects in Nickel-Based Alloys energy Ed ¼ 40 eV) In addition to precipitationhardened alloys, including PE16 and Inconel 706, this experiment included nonhardenable high-Ni alloys, such as Inconel 600 and Hastelloy X, a range of commercial steels, and Fe–Cr–Ni ternary alloys containing 15% Cr and 15–35% Ni The alloys were preimplanted with 15 appm helium prior to ion bombardment, and the irradiation temperature was chosen as being close to the peak swelling temperature for ionirradiated austenitic steels The extent of void swelling was determined by electron microscope examinations in low-swelling alloys, but was estimated from stepheight measurements (comparing the surfaces of irradiated and nonirradiated regions) in high-swelling materials As illustrated in Figure 1, the results showed negligible swelling (200 dpa (N/2) at 525  C Void swelling in 316 steel and nickel exceeded 10% at the highest doses examined, compared to $0.5% 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 precipitates, which were thought either to trap vacancies and interstitials at their surface, thereby enhancing pointdefect recombination, or to inhibit the climb of dislocations, 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 intermediate 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 10À2 dpa sÀ1 at 600  C, where the swelling rate was higher at small (3 nm) and large (70 nm) g0 particle 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.12 following bombardment with MeV Ni2+ ions at 625  C The damage rate in these experiments was 10À2 dpa sÀ1 and the displacement dose was originally estimated as 140 dpa but this was subsequently revised by Bates and Johnston13 to 116 dpa (based on displacement 125 30 20 10 10 10 20 30 11 40 50 60 Nickel (wt%) 12 13 70 80 90 Figure Swelling versus nickel content of commercial alloys and ternary Fe–15Cr–Ni alloys bombarded with Ni2ỵ ions to a damage level of 116 dpa at 625  C Reproduced from Johnston, W G.; Rosolowski, J H.; Turkalo, A M.; Lauritzen, T J Nucl Mater.1974, 54, 24–40 126 Radiation Effects in Nickel-Based Alloys an increase in overall swelling from 0.1 MeV) at temperatures from 454 to 593  C Swelling rates during Ni ion irradiations at 675  C were higher by a factor of about five in reactor-conditioned material than in a nonconditioned sample The increased swelling rate was attributed to changes in the matrix composition resulting from an increased volume fraction of g0 in the reactor-conditioned material Early attempts to account for the effects of matrix composition on void swelling focused on the stability of the austenite phase Harries16 suggested that the swelling behavior of austenitic steels and nickelbased alloys could be rationalized in terms of their Ni and Cr equivalent contents (i.e., the relative austenite and ferrite stabilizing effects of their constituent elements), with the composition of highswelling alloys then falling into the (g ỵ s) phase field in the Fe–Cr–Ni ternary phase diagram Harries postulated that the partitioning of solute elements into the sigma phase would have a detrimental effect on the swelling resistance of austenite Watkin17 took a similar approach, but found that an improved correlation could be obtained using the concept of electron vacancy numbers rather than Ni and Cr equivalents The average electron vacancy number, Nv, of the matrix is calculated from the atomic fractions of its constituents, with allowance being made for the precipitation of carbides and g0 (or g00, etc.), and has been widely used to predict the susceptibility of nickel-based alloys to the formation of intermetallic phases.18Nv was calculated from: Nv ¼ 0:66Ni ỵ 1:70Co ỵ 2:66Fe ỵ 3:66Mn ỵ 4:66Cr ỵ MoÞ Watkin found that void swelling in a range of alloys with Ni contents up to $43%, which were irradiated in DFR to a peak dose of 30 dpa at 600  C, remained low for Nv below about 2.5 (corresponding to low susceptibility to s phase formation), but increased approximately linearly at higher Nv However, as was clearly argued by Bates and Johnston,13 correlations based on sigma-forming tendency could not account for the minimum in swelling observed at about 45% Ni, since higher Ni contents should continue to be beneficial Radiation Effects in Nickel-Based Alloys 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 temperatures in these experiments ranged from about 400 to 650  C Initial data for a range of commercial alloys, including ferritic and austenitic steels, as well as nickel-based alloys, were reported by Bates and Powell19 and Powell et al.,20 with higher dose data (up to a peak fluence (E > 0.1 MeV) of $25  1026 n mÀ2, corresponding to $125 dpa) being reported by Gelles21 and Garner and Gelles.22 Swelling data for Fe–Cr–Ni ternary alloys, irradiated in EBR-II to a peak fluence of 22  1026 n mÀ2 ($110 dpa), were presented 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% exhibited the lowest swelling Swelling in commercial nickel-based alloys was generally lower in ST than in aged conditions, this being attributed to the beneficial (though temporary) effect of minor elements remaining in solution and being able to interact with point defects19; subsequent precipitation during irradiation 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 mÀ2 and 25.3  1026 n mÀ2, with measurements for Inconel 600 and Inconel 625 reported at both fluence levels, data 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 Incoloy 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 reaching $80% at 593  C The reason for such high swelling in neutron-irradiated Hastelloy X (nickel content $48%) 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 composition of the matrix which may have affected the swelling behavior (Note that Hastelloy X was identified 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 127 Garner and Gelles,22 and are compared for irradiations at 538  C (more or less corresponding to the peak swelling temperature for PE16 in the AA-1 experiment) with lower fluence data from Bates and Powell19 in Figure The heat-treated conditions indicated in Figure are ST (ST h at 1080  C), A1 (ST and aged 16 h at 705  C), A2 (ST and aged h at 890  C plus h at 750  C), and OA (ST and aged 24 h at 840  C) Note that the silicon content of the PE16 used in these experiments was much lower at 0.01% than the level of $0.2% typically found in UK heats of the alloy Overall, the data appear to show little effect of initial heat treatment on the swelling of PE16, except that the OA condition exhibited the most swelling (5.2%) at the higher fluence Although it is clear that the swelling behavior of austenitic alloys is largely dependent on nickel content, there is ample evidence to show that minor solute additions can have significant effects Much of the work on minor solutes has focused on steels similar to type 316, but some data are available for higher nickel alloys For example, Mazey and Hanks24 used 46.5 MeV Ni6+ ion irradiations to examine the effects of Si, Ti, and Al additions on the swelling response of model alloys with base compositions approximating that of the matrix phase in PE16 Solute additions of $0.25% Si or 1.2% Ti reduced swelling, but the addition of $1.2% Al (in the absence of Si or Ti) markedly increased it The beneficial effect of Si was believed to arise from its high diffusivity in solution (this is discussed further in Section 4.04.2.2), whereas that of Ti appeared to be related to the formation of Z phase (hexagonal-structured Ni3Ti) The addition of Al resulted in an increase in the concentration of voids, the surfaces of which were coated in a thin layer of the g0 phase (Ni3Al) A beneficial effect of Si on the swelling response of modified Incoloy DS alloys under Ni6+ ion irradiation was also reported by Mazey et al.25 However, it should be noted that high Si contents can give rise to the formation of radiationinduced phases which are enriched with Ni and Si, such as the Ni3Si form of g0 and the silicide G-phase (M6Ni16Si7, where M is usually Ti, Nb, or Mn) G-phase particles are generally found in association with large voids and their formation may therefore give rise to an increase in the swelling rate.26,27 Swelling data derived from density measurements for neutron irradiated, modified Incoloy DS alloys, with Si contents ranging from 0.19 to 2.05% (compared to a specified level of 1.9–2.6% in the commercial alloy), are compared with data for a ‘PE16 matrix 128 Radiation Effects in Nickel-Based Alloys 5.0 4.0 Fluence 15 Swelling (%) 3.5 3.0 10 2.5 2.0 1.5 1.0 0.5 Fluence (1026 n m−2, E > 0.1 MeV) In 600 In 625 PE16 In 706 4.5 0.0 −0.5 350 400 450 500 550 600 650 700 Temperature (°C) 100 80 In 600 In 625 In 800 Hast X Fluence Swelling (%) 70 25 20 60 50 15 40 30 10 20 10 Fluence (1026 n m−2, E > 0.1 MeV) 90 −10 350 400 450 500 550 600 650 700 Temperature (°C) Figure Void swelling of nickel-based alloys irradiated in AA-1 rig in Experimental Breeder Reactor-II Based on data 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 alloy’ and Nimonic PE16 in Figure The materials were all in ST condition apart from PE16 which was in an STA condition (aged h at 750  C) The alloys were irradiated in the UK-1 rig in EBR-II to fluences in the range of 9–16  1026 n mÀ2 (E > 0.1 MeV) at temperatures of $390–640  C These data are previously 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 with indications of swelling peaks at about 440 and 640  C Increased Si levels tended to suppress the high temperature swelling peak and reduce the magnitude of swelling at lower temperatures The PE16 matrix alloy containing 0.24% Si exhibited a high temperature swelling peak but moderate swelling below $550  C, suggesting a beneficial effect of Mo (this being the main compositional difference between the PE16 matrix alloy and the modified Incoloy DS alloys) Radiation Effects in Nickel-Based Alloys 4.0 at lower temperatures However, swelling in the PE16 matrix alloy remained significantly higher at all temperatures than in STA Nimonic PE16 (containing 0.15% Si), indicating a significant benefit of the g0 forming elements Al and Ti 3.0 4.04.2.2 6.0 68 dpa 116 dpa Swelling (%) 5.0 2.0 1.0 0.0 ST A1 A2 OA Figure Effect of heat treatment on void swelling of Nimonic PE16 irradiated in Experimental Breeder Reactor-II at 538  C 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 12.0 15 8.0 10 6.0 4.0 Fluence (1026 n m–2, E > 0.1 MeV) 10.0 Swelling (%) 129 2.0 0.0 300 350 400 450 500 550 Temperature (°C) PE16 STA PE16 Matrix DS (0.19 Si) 600 650 DS (0.56 Si) DS (1.05 Si) DS (2.05 Si) Figure 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 Void-Swelling Models Point defects created by atomic displacements are lost either through mutual recombination or by migration to sinks Void swelling requires a mobile population of excess vacancies and can only occur over a limited temperature range, typically $350– 700  C in neutron-irradiated steels and nickel-based alloys Rapid diffusion at higher temperatures reduces the concentration of radiation-induced vacancies to near thermal equilibrium levels Recombination dominates at lower temperatures, where reduced vacancy mobility prevents the formation of voids as the necessary counter-migration of matrix atoms cannot occur In the swelling regime, an increased bias for interstitials over vacancies at dislocation sinks gives rise to the surplus vacancies which agglomerate to form voids The flux of point defects to sinks, including void surfaces, dislocations, and grain boundaries, results in the segregation of particular solute atoms at the sinks and the depletion of others In austenitic steels and nickel-based alloys, it is generally found that nickel segregates at the point defect sinks This is generally attributed to the inverse Kirkendall effect described by Marwick,29 whereby faster diffusing solutes such as Cr move in the opposite direction to the vacancy flux and are depleted at the sink, and slower diffusing solutes such as Ni are enriched One of the earliest observations of nickel segregation at void surfaces due to the inverse Kirkendall effect was made by Marwick et al.30 in an alloy with a composition corresponding to that of the matrix phase in Nimonic PE16 (For more detailed discussions on radiationinduced segregation effects, see the reviews of Wiedersich and Lam,31 and Rehn and Okamoto.32) Venker and Ehrlich33 recognized that differences in the partial diffusion coefficients of alloy constituents might account for the effects of composition on swelling Any effect of this kind would generally be expected to be more significant the larger are the differences between the partial diffusion coefficients of the alloy components Garner and Wolfer34 examined Venker and Ehrlich’s conjecture and concluded that the addition of even small amounts of a fastdiffusing solute such as silicon to austenitic alloys would greatly increase the effective vacancy diffusion 130 Radiation Effects in Nickel-Based Alloys 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 mechanism 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 exchange, 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 diffusion data reported by Rothman et al.,36 indicated 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 calculated to increase at the higher nickel content Esmailzadeh 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 appropriate diffusion data As well as affecting the nucleation of voids, differences 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.37,38 The model examined the compositional dependence of the void bias and focused on the effects of nickel segregation at void surfaces Wolfer’s model indicated that the compositional gradients produced by radiation-induced segregation give rise to additional drift forces which affect the point defect fluxes and thereby modify the bias terms These additional drift forces arise from the effects of composition on point defect formation and migration energies, on the lattice parameter and the elastic moduli, and from the Kirkendall flux Wolfer’s calculations for binary Fe–Ni alloys indicated that the effect of the Kirkendall flux is small for interstitials but significant for vacancies Nevertheless, it was considered that the overall effect of compositional gradients on the bias terms is likely to be greater for interstitials than for vacancies due to other factors, particularly the effect of variations in the elastic moduli As noted by Garner and Wolfer,39 an increase in the shear modulus in the segregated regions around voids would reduce the bias for interstitials and therefore help to stabilize voids It is difficult to predict the significance of this effect in complex alloys, however, since depletion of Cr in the segregated region will tend to reduce the shear modulus, whereas enrichment of Ni in high-Ni alloys will tend to increase it.38 A more significant result of the model with regard to the effect of nickel on swelling is that there is a reversal in the sign of the Kirkendall force for vacancies in Fe–Ni alloys at $35% Ni Below this level, vacancies are predicted to be attracted into regions of higher Ni concentration, but above it, the opposite occurs Wolfer et al considered that this effect may account for the dependence of swelling on Ni content in austenitic alloys containing less than 35% Ni A generalized description of the swelling behavior of austenitic alloys, which was consistent with the model developed by Wolfer et al., was put forward by Garner40 (see also Chapter 4.02, Radiation Damage in Austenitic Steels) Garner’s ideas were largely based on the results of the EBR-II irradiation experiments and the earlier ion bombardment work of Johnston et al., both of which showed a strong dependence of swelling on nickel content It was considered that swelling was characterized by a transient period followed by a regime in which the swelling rate became constant In neutron-irradiated alloys, the swelling rate in the posttransient regime was generally found to be $1% per dpa In swellingresistant alloys, however, it was argued that such high swelling rates might not be observed owing to extended transient periods The duration of the Radiation Effects in Nickel-Based Alloys 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 included the effect on vacancy diffusivity described by Esmailzadeh and Kumar35; a possible correlation with the development of fine scale compositional fluctuations 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.42 Voids 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 highnickel alloys, particularly at relatively high temperatures 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 operation of the bubble to void transition was obtained by Mazey and Nelson,44 who implanted Nimonic PE16 (STA condition) and a PE16 matrix alloy (ST condition) 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 ( 0.1 MeV) (see Figure 6) Alloys containing 19% and 30% Ni exhibited high swelling rates at higher fluences, but swelling remained low in higher nickel alloys Similar effects were found in the ion-bombarded samples, where, for example, it was shown that there was no significant change in the void concentration in Fe–15Cr–45Ni at doses above 50 dpa in irradiations at 675  C, yet a marked increase in swelling rate occurred above 120 dpa Thus, contrary to earlier ideas, these investigations clearly demonstrated that the onset of a high swelling rate was not related to the cessation of void nucleation It follows that the transition to a high rate of swelling must be due to an increase in the growth rate of existing voids Muroga et al.45,46 observed that the total dislocation density in the irradiated Fe–15Cr–Ni alloys was only weakly dependent on nickel content This suggested that at the intermediate nickel levels, where the void concentration was low, dislocations were weak sinks (for both vacancies and interstitials) relative to voids In addition, it was observed that 132 Radiation Effects in Nickel-Based Alloys 20 400 300 16 0.2 dpa Bubbles 0.2 dpa 12 200 30 dpa 300 200 15 10 Bubbles Voids 100 0 Bubbles 300 60 dpa 15 Number of cavities in interval Nc ϫ 10–19 m–3 Number of voids in interval Nv ϫ 10–19 m–3 Number of bubbles in interval Nb ϫ 10–19 m–3 100 Voids 16 30 dpa 12 16 60 dpa 12 200 10 100 0 0 (a) 10 15 20 25 30 35 40 45 50 55 60 Cavity diameter d nm (b) 10 20 30 40 50 60 70 80 90 100 110 120 130 Cavity diameter, d (nm) Figure Histograms showing size distributions of bubbles/voids in (a) solution treated and aged Nimonic PE16 and (b) solution treated PE16 matrix alloy, irradiated with Ni6ỵ ions at 625  C 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 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 intermediate 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 Microchemical 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 arising from the inverse Kirkendall effect A correlation between the magnitude of void swelling and radiation-induced segregation was shown for Fe–Cr–Ni ternary alloys by Allen et al.48 The compositional dependence of radiation-induced segregation was determined using a model based on the earlier work of Marwick,29 which incorporates both the vacancy flux to the voids and the back-diffusion of vacancies due to the solute gradients set up by the inverse Kirkendall effect Vacancy diffusivities for various alloy compositions were determined by the measurements of grain boundary segregation in protonirradiated samples Swelling data for ion and neutron-irradiated alloys were then compared with the expected swelling propensity defined by the ratio of the forward to back diffusion terms calculated at the appropriate irradiation temperature The materials for which vacancy diffusivity data were determined included Fe-based alloys containing 16–24% Cr and 9–24% Ni, and Ni-based alloys containing 18% Cr and either zero or 9% Fe This work did not specifically examine 40–50% Ni alloys corresponding to the highest swelling resistance, though the results indicated that swelling generally decreased with increasing levels of nickel enrichment and chromium depletion at void surfaces 136 Radiation Effects in Nickel-Based Alloys regarded as benign and permitted continued operation, 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 mechanisms is beyond the scope of this chapter, which will instead concentrate on experimental data which enable comparisons to be made between nickelbased alloys and austenitic steels However, some insight into irradiation creep mechanisms is given in Section 4.04.4.1, where the effect of stress on the evolution of dislocation structures is described Irradiation creep mechanisms are discussed more fully in Chapter 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 pressurized tubes and helical springs, have been used to measure irradiation creep strains The data are therefore generally converted to effective strain e  values, using the Soderberg and effective stress s 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 irradiation temperatures for PE16 specimens ranged from about 280 to 340  C, with displacement doses up to a maximum of $13 dpa (N/2) For austenitic steels, the irradiation creep strain was found to be linearly dependent on the applied stress and the displacement dose, comprising transient and steady-state components as follows: g ẳ At ỵ Bd t where d is the displacement dose and A and B are material-dependent creep coefficients For PE16 in a STA condition (1 h at 1080  C plus 16 h at 700  C), creep at dose rates of $5  10À7 dpa (N/2) sÀ1 was characterized by an initial period of low strain and an increased creep rate at higher displacement doses Mosedale et al.62 described the g=t versus dpa creep curve for STA PE16 as parabolic, though the maximum observed creep rate was similar to that in austenitic steels and Harries59 represented the creep strain above a threshold dose of dpa (N/2) by g ¼ 4:3  10À6 tðd À 8Þ where t is in MPa; converting to effective strain/ stress values and to NRT units of displacement dose (assuming dpa (N/2) ¼ 0.8 dpa (NRT-Fe)) would reduce the creep coefficient by a factor of 2.4 Data presented by Lewthwaite and Mosedale63 showed that ST PE16 behaved similarly to the STA condition, though OA conditions exhibited higher creep strains due to a combination of increased creep rates and low threshold doses (around dpa) An apparent dose-rate dependency was observed, with steadystate creep coefficients for STA and OA PE16 increased by factors of $2 at lower damage rates of $0.5–1.5  10À7 dpa (N/2) sÀ1 and threshold doses reduced to $0.5 dpa or less A similar effect of dose rate on the creep strain per dpa was also reported for austenitic steels.64 Steady-state creep coefficients (MPaÀ1 dpaÀ1) and creep strain rates (MPaÀ1 sÀ1) for PE16 as a function of dose rate are compared with data for cold-worked steels M316 and FV548 in Figure 10 The data plotted in Figure 10 are derived from the results of Lewthwaite and Mosedale63,64 but are converted to effective strain/stress values and NRT(Fe) dpa units to enable comparison with other published data It is evident that the irradiation creep behavior of STA and OA (24 h at 800  C) PE16 is similar to that of the austenitic steels Creep rates at higher dose rates are generally lower than would be indicated from the linear extrapolation of low dose rate data Lewthwaite and Mosedale63 Radiation Effects in Nickel-Based Alloys Creep coefficient, B (10–6 MPa–1 dpa–1) 6.0 M316 FV548 PE16 OA PE16 STA 5.0 4.0 3.0 2.0 1.0 0.0 0.0 1.0 2.0 3.0 4.0 5.0 Displacement rate (10–7 dpa s–1) 10.0 Creep strain rate (10–13 MPa–1 s–1) 9.0 8.0 M316 FV548 PE16 OA PE16 STA 7.0 6.0 5.0 4.0 3.0 2.0 1.0 0.0 0.0 1.0 2.0 Displacement rate 3.0 4.0 5.0 (10–7 dpa s–1) Figure 10 Steady-state creep coefficients and creep strain rates for Nimonic PE16 and austenitic steels, derived from the measurements of Lewthwaite and Mosedale Adapted from Lewthwaite, G W.; Mosedale, D In Proceedings of International Conference on Irradiation Behaviour of Metallic Materials for Fast Reactor Core Components, Ajaccio, Corsica, June 4–8, 1979; Poirier, J., Dupouy, J M., Eds.; Le Commissariat a l’Energie Atomique (CEA): Saclay, France, 1979; pp 399–405; Lewthwaite, G W.; Mosedale, D J Nucl Mater 1980, 90, 205–215 considered that the measured irradiation creep rates for PE16 at low dose rates were in close agreement with the expected rates for SIPA-(stress-induced preferred absorption of interstitials at dislocations) controlled creep It was suggested by Mosedale et al.62 137 that reduced creep rates at higher dose rates might be attributable to increased recombination rates for vacancies and interstitial atoms, although a more detailed assessment of this effect by Lewthwaite and Mosedale64 proved inconclusive and a dose-rate dependency has not generally been observed in other experiments.60 Garner and coworkers65,66 considered that the higher creep rates measured by Lewthwaite and Mosedale at lower displacement rates were an aberration due to transient effects at low dpa levels Nevertheless, this does not alter the finding that the irradiation creep behavior of PE16 is comparable to that of austenitic steels Paxton et al.67 examined the in-reactor creep behavior of a number of alloys, including Nimonic PE16, Inconel 706, and Inconel 718, as well as austenitic and ferritic steels, in pressurized tube experiments carried out in EBR-II at 540  C to fluences up to  1026 n mÀ2 (E > 0.1 MeV) Diametral strains measured in pressurized tubes (with hoop stresses in the approximate range of 25–175 MPa) were corrected for void swelling and/or densification observed in unstressed specimens (though this does not allow for any effects of stress on swelling or precipitation processes) Precipitation-hardened alloys exhibited lower creep strains than solid solution strengthened steels, with the Inconel alloys superior to PE16 at 540  C The creep resistance of the precipitationhardened materials was also dependent on heat treatment, with ST conditions generally superior to aged conditions However, it was noted that ST conditions also exhibited greater densification – giving rise to the possibility of increased fuel–clad interactions in fuel elements In-reactor creep strains were discussed in terms of a widely used model which includes a term for creep enhancement due to swelling The total effective creep strain e is given by e ẳ B0 ft s  ỵ DS s where B0 is the creep compliance, ft is the neutron fluence, D is the creep–swelling coupling coefficient, and S is the fractional swelling A contribution from thermal creep may be expected at 540  C, but data to correct for this component were not available and hence the creep coefficients could not be determined precisely The stress dependence of the measured creep strain was approximately linear in the low swelling precipitation-hardened alloys, though nonlinearity attributable to the effects of stress on swelling was observed in the solid solution alloys An approximate value of B0 of 1.5  10À28 MPaÀ1 (n cmÀ2)À1, which is equivalent to  10À7 MPaÀ1 138 Radiation Effects in Nickel-Based Alloys dpaÀ1, was derived by Paxton et al for the Inconel alloys Ehrlich60 subsequently made estimates of B0 for the other materials included in this study, which ranged from $1.4  10À6 MPaÀ1 dpaÀ1 for ST PE16 to $10À5 MPaÀ1 dpaÀ1 for cold-worked 316 steel Paxton et al noted that values of the creep–swelling coefficient D appeared to be much larger for the solid solution strengthened steels than for the precipitationhardened alloys, with the higher values being attributable to increased thermal creep components and/or the effects of stress on swelling Gilbert and Chin68 examined the nonisothermal creep behavior of EBR-II-irradiated PE16 and Inconel 706 Both materials were in ST conditions Pressurized tubes, with nominal hoop stresses of 100 MPa for PE16 and 200 MPa for Inconel 706, were irradiated at 425, 540, and 590  C, both isothermally and with temperature steps Diametral strains for isothermally irradiated PE16 increased with increasing fluence and temperature as expected Following temperature changes from 540 to 590  C or 425  C, the creep rate for PE16 adjusted to the isothermal rate at the new temperature For Inconel 706, however, the isothermal creep rate was highest at 425  C, and an upward step to 540  C resulted in a reduced creep rate; a downward step from 540 to 425  C gave rise to an increased creep rate that exceeded the isothermal rate at 425  C; and an upward step from 540 to 590  C reduced the creep rate, even though the isothermal creep rate was higher at 590 than 540  C The complex in-reactor creep behavior of Inconel 706 appeared to be related to the stability of the ordered body-centered tetragonal, Ni3Nb g00 phase and its effect on thermal creep resistance Gilbert and Chin considered that the inreactor deformation of Inconel 706 was primarily controlled by thermal rather than irradiation creep processes, since similar creep rates were reported to occur in thermal control tests Microstructural examinations made by Thomas69 indicated that g00 precipitated during irradiation above $500  C but dissolved at lower temperatures, thereby reducing the creep strength of the material Gelles70 subsequently reported that the dissolution of g00 at low irradiation temperatures appeared to be promoted by the application of stress since more of this phase was retained in unstressed material Toloczko et al.5 investigated the swelling and creep behavior of five austenitic alloys which were irradiated in PFR in a pressurized tube experiment at $420  C The materials examined included the solid solution strengthened steels 316 and D9, and the higher-Ni precipitation-hardened alloys D21, D68, and D66 Dose rate variations were examined by positioning specimens at different axial locations within the reactor core The tubes were removed periodically for diameter measurements, with peak doses of $50 dpa being attained at the highest flux level Hoop stresses ranged from to 150 MPa, and swelling as a function of dose was estimated from measurements on unstressed tubes assuming that densification effects were completed during the first irradiation cycle There was some scatter in the results but the creep coefficient B0 was found to be relatively independent of alloy composition and dose rate, with typical values of $1.0–1.4  10À6 MPaÀ1 dpaÀ1 (though higher values were determined for type 316 steel) The creep–swelling coupling coefficient D was also independent of dose rate but appeared to be material dependent (with values in the approximate range of 0.4–1.6  10À2 MPaÀ1), though this variability could not be associated with any particular compositional factor Similar results for two precipitation-hardened high-nickel alloys (with similar compositions to Nimonic PE16, but with additions of $0.5% Nb), which were irradiated in a pressurized tube experiment in the Russian fast reactor BN-350 to $90 dpa at 400  C, were also reported by Porollo et al.71 4.04.4 Microstructural Stability 4.04.4.1 Dislocation Structures Dislocation structures in irradiated pressurized tube samples were examined by Gelles et al.72 The materials which were examined included stressed and unstressed samples of ST PE16, and stressed samples of ST and STA Inconel 706 A subsequent paper by Gelles73 extended these investigations to the stressed samples of PE16 in STA and OA conditions Further details of this work were also provided by Garner and Gelles74, and by Gelles.70 Examination of ST PE16, which was irradiated at 550  C to  1026 n mÀ2 (E > 0.1 MeV) at hoop stresses of and 167 MPa, revealed that the distribution of Frank dislocation loops was similar on all the four {111} planes in the unstressed sample but was anisotropic in the stressed material In the stressed sample, the loop density on any particular {111} plane increased with increasing magnitude of the normal stress component on that plane A near-linear relationship between the loop density and the normal Radiation Effects in Nickel-Based Alloys component of the deviatoric stress tensor, sDN (¼ sN À sH , where sN is the normal component of the applied stress on a particular plane and sH is the hydrostatic stress), was found for PE16 This result is in line with the SIPA loop growth model described by Garner et al.75 No such correlation was found in the similarly irradiated and stressed Inconel 706 samples, however, possibly because the low creep rate of this material at 550  C did not allow the relaxation of internal stresses Unfaulting of Frank dislocation loops with a/3 {111} Burgers vectors proceeds via interaction with a/6{112} Shockley partials to produce perfect a/2 {110} line dislocations Gelles70 described how this occurs via a two-step process, with the necessary partial dislocations (two per interstitial loop) first being nucleated by an interaction of the faulted loop with a suitable perfect dislocation and then sweeping across the loop to reestablish the perfect dislocation Gelles73 examined the distribution of Burgers vectors among the six possible a/2{110} perfect dislocation types in irradiated pressurized tube samples of PE16 The samples examined included the stressed ST condition irradiated at 550  C, and STA and OA conditions which were both irradiated at 480  C to a fluence of  1026 n mÀ2 at a hoop stress of 331 MPa The results showed highly anisotropic distributions in the Burgers vectors of perfect dislocations in all the three heat-treated conditions, with dislocation densities of the various types differing by factors of up to 10–40 in each sample The level of anisotropy produced in the population of perfect dislocations was significantly greater than in the dispersion of Frank loops This is a feasible outcome since, in principle, all loops may be unfaulted by just two variants of the six a/2{110} perfect dislocation types In effect, the development of anisotropic dislocation structures is a response of the material to produce the strain which is required to accommodate the applied stress Furthermore, it was found that the perfect dislocations in the irradiation creep samples of PE16 were primarily of edge type lying on {100} planes rather than {111} slip planes, indicating that they could only contribute to the creep strain via climb (i.e., by the SIPA mechanism) and not by processes involving dislocation glide 4.04.4.2 Precipitate Stability Early models of precipitate stability under irradiation were based on the ideas of Nelson et al.,76 who 139 suggested that precipitates would evolve to an equilibrium size determined by competing processes affecting their growth, via the radiation enhanced and/or thermal diffusion of solutes, and their simultaneous dissolution due to damage arising in collision cascades Two dissolution mechanisms were suggested: recoil dissolution due to the displacement of atoms from the precipitate into the matrix, and disordering dissolution of ordered phases such as g0 , with the latter predicted to be the more effective The model predicted that fine precipitates would continue to grow to some equilibrium size (dependent on temperature, dose rate, and solute levels), but that precipitates greater than this size would shrink Experimental evidence for the dissolution of large preexisting Ni3Al g0 precipitates in heavy-ionirradiated Ni–Al alloys was shown by Nelson et al.76 These ideas were developed further and applied to g0 precipitates in ion and neutron-irradiated alloys by Baron et al.77 The model developed by Baron et al indicated that, at a given particle size, a higher solute supersaturation was required under irradiation than in a purely thermal environment The model appeared to be consistent with the observed coarsening behavior of g0 precipitates during irradiation, though no evidence for the shrinkage of large particles was presented For example, data for PE16 irradiated at fluences up to 7.5  1026 n mÀ2 at 560  C, which were reported by Chang and Baron,78 only examined the growth of g0 particles up to a maximum radius of $15 nm under conditions where the predicted maximum equilibrium radius was $35 nm However, detailed examinations of g0 structures in neutron-irradiated Nimonic PE16 which were made by Gelles79 found no evidence to indicate that irradiation-induced dissolution mechanisms limited the particle size Microstructural examination of PE16, originally in ST, STA, and OA conditions, irradiated in EBR-II to $27 dpa (5.4  1026 n mÀ2, E > 0.1 MeV) at 600  C, revealed that preexisting g0 dispersions in aged material were maintained but continued to coarsen even in the OA condition, and that a fine dispersion formed in ST material Coarsening of the g0 particles in the OA material was accompanied by the formation of fine background precipitates in some regions Further in-reactor precipitation of g0 also occurred at point defect sinks, including void surfaces and dislocations, in all the heat-treated conditions Additional examinations by Gelles80 of ST PE16, irradiated to $30–50 dpa at temperatures in the range of 430–650  C, indicated that g0 coarsening was controlled by radiation-enhanced diffusion 140 Radiation Effects in Nickel-Based Alloys below 600  C with an activation energy that (in agreement with theoretical predictions for a process governed by point defect recombination) was approximately a quarter of that for thermal diffusion As described in Section 4.04.5.1 in relation to irradiation embrittlement effects, Yang81 examined an identically irradiated set of ST PE16 samples as Gelles, focusing on the precipitation of g0 at grain boundaries Similar g0 structures to those described by Gelles and Yang were also observed by Boothby28 in the aged conditions of EBR-II-irradiated PE16, though at higher irradiation temperatures (!540  C for the STA condition, and !600  C for the OA condition), where doses were in the range 66–74 dpa, the spherical g0 precipitates which formed during thermal aging were almost entirely replaced by ‘skeletal’ forms nucleated at point defect sinks Figure 11 shows an example of the g0 distribution, imaged in dark field, in STA PE16 irradiated to 69 dpa at 570  C; although small spherical precipitates were retained in a narrow region adjacent to the grain boundary, a much coarser dispersion is evident at the boundary itself and within the bulk of the grain 4.04.5 Irradiation Embrittlement The effects of fast neutron irradiation on the tensile properties of several precipitation-hardened nickelbased alloys were investigated in the 1970s and 1980s The materials examined included a number of g0 /g00 hardened alloys, such as the Inconel alloys 706 and 718 and the developmental alloys D68 and 7818, as well as g0 -hardened alloys similar to Nimonic PE16 Earlier work by Broomfield et al.82 on thermal reactor irradiated materials indicated that PE16 was more susceptible to irradiation embrittlement at elevated test temperatures than austenitic steels Broomfield83 found that thermal neutron irradiated PE16 was most severely embrittled in low strain tests at $550–650  C, and attributed this to an increased tendency for intergranular failure arising from the effects of helium generated from the 10B(n,a)7Li reaction Boron itself is considered to have a beneficial effect on (unirradiated) creep rupture life, as it segregates to grain boundaries and inhibits intergranular cracking, and additions of a few 10s of ppm are therefore, generally made to nickel-based alloys, including PE16.84 Nickel is also a major source of helium in neutron-irradiated alloys, with the twostage 58Ni(n,g)59Ni(n,a)56Fe reaction becoming the dominant source at high thermal neutron fluences, and nickel threshold reactions accounting for the greater part of helium production in fast neutron spectra.85 For example, the rate of helium generation in fast reactor irradiated PE16 was estimated by Boothby28 to be $1.2 appm per dpa, with about 85% of the helium being generated from nickel threshold reactions (see also Chapter 1.06, The Effects of Helium in Irradiated Structural Alloys) Nevertheless, other factors, including irradiationinduced strengthening and grain boundary segregation and precipitation effects, have been implicated in the embrittlement of fast neutron irradiated nickel-based alloys 4.04.5.1 Fast Neutron Irradiation Experiments 200 nm Figure 11 Dark field, transmission electron micrograph, illustrating the distribution of g0 precipitates in solution treated and aged Nimonic PE16 irradiated in Experimental Breeder Reactor-II to 69 dpa at 570  C Unpublished data from Boothby, R M The Microstructure of EBR-II Irradiated Nimonic PE16; AEA TRS 2002 (FPSG/P(90)23), with permission from AEA Technology Plc Rowcliffe and Horak86 investigated the tensile properties of Inconel 706 (in a multistep ‘fully aged’ condition) and Inconel 718 (ST condition) following irradiation in EBR-II to fluences of 4–5  1026 n mÀ2 (E > 0.1 MeV) Irradiation temperatures (Ti) ranged from 450 to 735  C, with tensile tests being performed at a strain rate of  10À4 sÀ1 at temperatures corresponding to Ti and to Ti ỵ 110  C Yield stresses and total elongation data for Inconel 706 are shown in Figure 12 and for Inconel 718 in Figure 13 Data for Inconel 706 showed very high (>1000 MPa) yield stresses and ultimate tensile strengths (UTS) in Radiation Effects in Nickel-Based Alloys specimens irradiated at temperatures up to and including 500  C This high tensile strength was maintained in a specimen irradiated at 500  C but tested at 610  C Although there was some reduction in strength in specimens irradiated at 560  C and above, the UTS remained above 650 MPa in specimens irradiated at 625  C The very high tensile strengths exhibited at the lower irradiation temperatures were attributed to the instability of the (ordered body-centered tetragonal) g00 phase below 525  C and its consequent dissolution, leading to the reprecipitation of nickel and niobium as (ordered face-centered cubic) g0 on dislocation loops At higher irradiation temperatures, both g0 and g00 were stable, but 1400 20 Yield at Ti + 110 °C Elong at Ti + 110 °C 1200 Yield at Ti 18 16 14 12 800 10 600 400 Total elongation (%) Elong at Ti 1000 Yield stress (MPa) 141 200 400 450 500 550 600 650 700 750 800 Temperature (°C) Figure 12 Yield stress and total elongation values at the irradiation temperature (Ti) and at Ti þ 110  C for Experimental Breeder Reactor-II-irradiated Inconel 706 Based on data from Rowcliffe, A F.; Horak, J A Am Nucl Soc Trans 1981, 38, 266–267 20 1400 Yield at Ti Elong at Ti 1200 Yield at Ti + 110 °C Elong at Ti + 110 °C 18 Yield stress (MPa) 1000 14 12 800 10 600 400 Total elongation (%) 16 200 400 450 500 550 600 650 Temperature (°C) 700 750 800 Figure 13 Yield stress and total elongation values at the irradiation temperature (Ti) and at Ti ỵ 110  C for Experimental Breeder Reactor-II-irradiated Inconel 718 Based on data from Rowcliffe, A F.; Horak, J A Am Nucl Soc Trans.1981, 38, 266–267 142 Radiation Effects in Nickel-Based Alloys precipitate coarsening resulted in lower tensile strength Elongations to failure for tests carried out at the irradiation temperature were between 1.5% and 3% up to 625  C, compared to >8% in unirradiated material Irradiation embrittlement was generally more severe in tests at Ti ỵ 110  C, particularly at 610–735  C where the lowest recorded ductility was 0.2% Fractures in irradiated Inconel 706 were predominantly intergranular, with failure believed to be facilitated by the decohesion of Z phase (hexagonal Ni3(Ti,Nb)) platelets which were formed at grain boundaries during the initial heat treatment Rowcliffe and Horak’s data for ST Inconel 718 showed similar trends to Inconel 706 Precipitation of the g0 and g00 phases occurred during the irradiation of Inconel 718, resulting in yield strengths in excess of 1000 MPa at irradiation temperatures up to 560  C and above 800 MPa at 625  C The ductility of Inconel 718 was reduced from more than 30% in the unirradiated condition to 0.2% or less in specimens which were irradiated at 500560  C and tested at Ti ỵ 110  C In contrast to Inconel 706, failures in irradiated Inconel 718 were reported to be predominantly transgranular Crack propagation in Inconel 718 appeared to have been via a ‘channel’ fracture mechanism, that is, with deformation occurring by highly localized planar slip and consequent linkage of radiation-induced voids Bajaj et al.87 examined the tensile properties of Nimonic PE16 irradiated in EBR-II to neutron fluences up to a maximum of $7  1026 n mÀ2 (E > 0.1 MeV), at temperatures in the range of 450– 735  C The alloy was in a STA (1 h at 900  C plus h at 750  C) condition, and appears to have been the same low-Si heat of PE16 that was subsequently used in the AA-1 swelling experiment described by Garner and Gelles.22 Tensile tests were carried out at 232  C (to simulate refueling conditions), at the irradiation temperature Ti and at Ti ỵ 110  C (to simulate reactor transients), at a strain rate of  10À4 sÀ1, and with a small number of tests at  10À3 sÀ1 Irradiated specimens tested at 232  C generally showed a substantial increase in yield stress and a small increase in UTS over the unirradiated values (although samples irradiated at the highest temperature of 735  C exhibited some softening), and retained good levels of ductility with total elongation values above 10% Yield stress and total elongation data for PE16 at higher test temperatures are shown in Figure 14 for specimens irradiated to a fast neutron fluence of 4.3  1026 n mÀ2 (enabling direct comparison with the data for the similarly irradiated Inconel alloys shown in Figures 12 and 13) Specimens tested at the irradiation temperature again showed strengthening at temperatures in the range of 450–625  C and softening at 735  C, with good 20 1000 Yield at Ti + 110 °C Elong at Ti + 110 °C 800 Elong at Ti 18 16 Yield stress (MPa) 14 12 600 10 400 Total elongation (%) Yield at Ti 200 400 450 500 550 600 650 Temperature (°C) 700 750 800 Figure 14 Yield stress and total elongation values at the irradiation temperature (Ti) and at Ti ỵ 110  C for Experimental Breeder Reactor-II-irradiated Nimonic PE16 Based on data from Bajaj, R.; Shogan, R P.; DeFlitch, C.; et al In Effects of Radiation on Materials: 10th Conference; Kramer, D., Brager, H S., Perrrin, J S., Eds.; American Society for Testing and Materials: Philadelphia, PA, 1981; pp 326–351, ASTM STP 725 Reprinted, with permission, from ASTM STP725-Effects of Radiation on Materials, copyright ASTM International, 100 Barr Harbor Drive, West Conshohocken, PA 19428 Radiation Effects in Nickel-Based Alloys ductility at 450  C but total elongations reduced to $3% at 560625  C Tests at Ti ỵ 110  C showed further increases in tensile strength (consistent with the greater hardening expected from irradiation at a lower temperature) and more severe embrittlement with ductility levels at 670–735  C reduced to 0.3% at a fluence of 4.3  1026 n mÀ2 and to zero (i.e., failure before yield) in higher dose samples (7.1  1026 n mÀ2) Tests at Ti at the higher strain rate resulted in an improvement in ductility by a factor of two or three Examination of fracture surfaces showed that failures were predominantly intergranular in irradiated samples tested above $550  C, transgranular at 232  C, and mixed mode at 450–550  C Bajaj et al considered that the irradiation embrittlement of PE16 evident at high temperatures could simply be explained by matrix hardening with little or no change in the grain boundary fracture strength – evidenced by increases in yield strength but no significant changes in true (as opposed to engineering) UTS values – so that mechanisms relying on the weakening of grain boundaries could be discounted for the test conditions studied Sklad et al.50 reported tensile data for two aged conditions of Nimonic PE16 which were irradiated in EBR-II to 1.2  1026 n mÀ2 (E > 0.1 MeV) at 500  C and tested at strain rates from $3  10À5 to  10À3 sÀ1 There was no significant difference in the postirradiation properties of the two differently aged conditions, although one aging treatment (2 h at 800  C plus 16 h at 700  C) resulted in an unirradiated yield stress $25% higher than the other condition (1 h at 900  C plus h at 750  C) No effect of strain rate on tensile properties was evident in tests at the irradiation temperature, where total elongations remained above 10% Tests at higher temperatures were made only at the lowest strain rate, with failure elongations being reduced to 1.6% at 600  C and 0.5% at 700  C The low ductility failures were associated with an increased tendency toward intergranular fracture, and additional tests, in which samples irradiated to  1026 n mÀ2 at 500  C were fractured in situ in an Auger spectrometer, revealed helium release from samples which fractured intergranularly as well as the segregation of Ni, P, and S to grain boundaries Helium release was estimated at $0.03 He atoms per grain boundary atom No grain boundary helium bubbles were observable by TEM, and it was therefore considered that helium either remained in solution as a partial monolayer or was present in unresolved bubbles less than 1–2 nm in diameter The presence of grain boundary helium bubbles in Nimonic PE16 was reported by Fisher et al.88 in 143 sections of AGR (advanced gas-cooled reactor) tie bars irradiated at 512  C and above AGR tie bars, which are approximately 10 m long and are under load only during charging and discharging of the fuel element stringers, operate at temperatures from 325 to 650  C from bottom to top, with peak doses of $3 dpa occurring at around the m position Stress-rupture testing at 600  C at an applied stress of 500 MPa showed a trough in properties (i.e., a minimum in failure times) and intergranular failures in sections of some tie bars which were irradiated at temperatures in the range of 350–400  C where grain boundary helium bubbles were not generally observed Even so, grain boundary cavitation was observed in a fractured tie bar section which was irradiated at 360  C, with the cavities appearing to be nucleated (possibly at submicroscopic helium bubbles) at the intersection of slip bands with the boundary The trough in stress-rupture properties occurred in tie bar sections which exhibited both high yield strengths (attributable to high concentrations of dislocation loops and small voids) and high levels of grain boundary segregation EDX (energy dispersive X-ray) analyses showed a significant enrichment of Ni and Si, and a depletion of Fe, Cr, and Mo, at the grain boundaries of sections irradiated at 335–585  C In addition, high levels of Si were detected in sections irradiated at 335–512  C in the g0 phase that precipitated at the surface of voids, with the Si content increasing with decreasing irradiation temperature Although the presence of Si-enriched g0 phase at grain boundaries could not be confirmed, it was suggested that its formation may have contributed to the minimum in stress-rupture life, which was thought to result from the weakening of the boundaries relative to the matrix Grain boundary helium bubbles were also observed by Boothby and Harries89 and Boothby28 in PE16 irradiated in DFR and EBR-II at 535  C and above Tensile testing of DFR-irradiated PE16, exposed to $20 dpa at 465–635  C, and strained at a rate of 2.5  10À6 sÀ1 at temperatures approximating those of irradiation, revealed severe embrittlement with minimum elongations of $0.2% at 550  C; TEM examination of strained specimens provided evidence of intergranular cavitation, and the ductility data were interpreted using a model for the diffusion-induced growth of cavities nucleated at grain boundary helium bubbles.89 The postirradiation tensile properties and microstructure of developmental g0 (D21, D25, and D66) and g0 /g00 (D68) strengthened alloys were discussed 144 Radiation Effects in Nickel-Based Alloys by Yang et al.4 The alloys were all irradiated in a ST condition; additionally, D25 was irradiated in an aged (24 h at 700  C) condition (STA), and D66 in a 30% cold-worked plus aged (11 h at 800  C plus h at 700  C) condition (CWA) Specimens were irradiated at 450–735  C to a fast neutron fluence of  1026 n mÀ2 (E > 0.1 MeV) in EBR-II, and were tested at Ti, Ti ỵ 110  C and 232  C Severe irradiation embrittlement was evident in the ST alloys and STA D25, particularly in tests at Ti ỵ 110  C Zero ductility was recorded in the lower-Ni alloy D21 (25Ni–8Cr) irradiated and tested at 550 and 600  C Severe ductility losses were associated with intergranular failures, which were attributed to irradiation-induced solute segregation and consequent precipitation of brittle g0 layers at grain boundaries However, reasonable levels of ductility, ranging from to 6%, coupled with transgranular failures, were obtained at all temperatures in irradiated CWA D66 (45Ni–12Cr) The preirradiation grain boundary structure of this material, comprising a ‘necklace’ of small recrystallized subgrains plus large g0 particles and discrete Laves particles, remained stable with no indication of irradiation-induced g0 layers Yang et al considered that the radiation-induced segregation of g0 forming solutes to grain boundaries was inhibited by the introduction of a high density of dislocation sinks by cold working Vaidyanathan et al.90 and Huang and Fish91 examined the embrittlement of EBR-II-irradiated, precipitation-hardened alloys, using ring ductility tests In this test, small sections of tubing are compressed and the ductility, defined as the strain at the initiation of cracking, is deduced from the change in the sample radius of curvature at maximum load Both experiments included Inconel 706 and Nimonic PE16 in ST conditions, while Vaidyanathan et al also examined the developmental alloys D25 and D68 in ST and STA conditions Peak fluences in these experiments were around 6–7  1026 n mÀ2 (E > 0.1 MeV) and irradiation temperatures were in the range 460–616  C All the materials exhibited low ductility failures at high test temperatures, particularly in tests at about Ti þ 110  C where ductilities were generally below 0.1%, though Vaidyanathan et al found that postirradiation heat treatments (typically of h at 785  C) produced a moderate recovery in ductility Based largely on observations reported by Yang81 for irradiated ST PE16, Vaidyanathan et al and Huang and Fish considered that the irradiationinduced embrittlement of precipitation-hardened alloys could generally be attributed to the formation of brittle g0 layers at grain boundaries However, the arguments presented were far from conclusive – microstructural examinations of the developmental alloys which were reported by Vaidyanathan et al showed only weak indications of g0 precipitation in D25 even within the grains, and evidence for g0 precipitation at grain boundaries in D68 was not found in the low ductility tested samples but only in material irradiated to a higher fluence Yang81 examined the microstructure of a low Si (0.01%) heat of ST PE16, which was irradiated in EBR-II to doses of about 30 and 50 dpa at temperatures from 425 to 650  C Grain boundary g0 layers were observed in ST PE16 samples which were irradiated at 510  C or above but not at 425  C, and helium bubbles were detected at boundaries in samples irradiated at 600–650  C It was considered by Yang that the irradiation-induced embrittlement of ST PE16 was mainly attributable to the cleavage fracture of grain boundary g0 layers and that any effects of helium were of secondary importance However, although grain boundary precipitation of g0 was observed by Boothby28 in PE16 irradiated to relatively high doses in EBR-II, there was no evidence for the formation of intergranular g0 layers in the aged conditions of PE16 which exhibited low ductility failures following irradiation in DFR to $20 dpa.89 Thus, although it remains possible that the formation of grain boundary g0 layers may aggravate the embrittlement, it was considered by Boothby28 that the irradiation embrittlement of PE16 is primarily due to helium A breach in solution-annealed Inconel 706 fuel pin cladding, irradiated to 5% burn-up in EBR-II, was reported by Yang and Makenas.92 The rupture occurred from 12.7 to 18.4 cm from the bottom of the pin, corresponding to irradiation at 447–526  C at a fluence of  1026 n mÀ2 (E > 0.1 MeV) Microstructural examinations revealed a brittle intergranular fracture, with failure being attributed to a combination of matrix hardening due to g0 precipitation and grain boundary weakening due to the formation of interconnected Ni3(Ti,Nb) Z phase particles In contrast to the work of Rowcliffe and Horak86 where grain boundary Z phase was precipitated during a preirradiation aging treatment, this phase formed during the irradiation period in the solutionannealed cladding Precipitation of Z was considered to be irradiation enhanced because it was not formed in long-term thermal annealing experiments at 480– 540  C Grain boundary precipitation of Z phase was also observed at the hot (650  C) end of the fuel pin cladding, with both g0 and g00 in the matrix Cauvin et al.93 and Le Naour et al.94 also attributed irradiation embrittlement effects in Inconel 706 Radiation Effects in Nickel-Based Alloys cladding to the combined effects of matrix hardening and the precipitation of Z at grain boundaries Inconel 706 fuel pin cladding, fabricated from four heats with Nb contents varying from to 3% and in two heat-treated conditions (solution annealed or aged), was irradiated in the Phenix fast reactor up to a maximum of 100 dpa Tensile tests on cladding sections were carried out at a strain rate of  10À4 sÀ1 Tensile tests performed at ambient temperature showed high UTS (>1000 MPa) along the full length of the pins with peak values of $1500 MPa in sections irradiated near 500  C; ductility values (uniform elongations only were given) remained low ( MeV) of 1.7  1024 n mÀ2 and a thermal fluence of 5.9  1024 n mÀ2 The helium content of the reactor-irradiated specimens was estimated to be $45 appm, produced mainly from the thermal neutron reaction with 10B Tensile tests were carried 145 out at the implantation/irradiation temperature at a strain rate of $5  10À4 sÀ1 The results showed similar trends in helium-implanted and neutronirradiated specimens, with the total elongation values tending to decrease with increasing tensile strength Variations in tensile strength for each alloy were largely attributable to variations in the initial heat treatment and working schedules However, there were some indications of softening and reduced ductility in the neutron-irradiated specimens compared to those injected with helium Overall, the g0 hardened alloy 7817 exhibited relatively high tensile strength (typically >700 MPa) but low ductility following helium implantation or neutron irradiation (with total elongation values generally