Comprehensive nuclear materials 4 02 radiation damage in austenitic steels

Comprehensive nuclear materials 4 02 radiation damage in austenitic steels Comprehensive nuclear materials 4 02 radiation damage in austenitic steels Comprehensive nuclear materials 4 02 radiation damage in austenitic steels Comprehensive nuclear materials 4 02 radiation damage in austenitic steels Comprehensive nuclear materials 4 02 radiation damage in austenitic steels Comprehensive nuclear materials 4 02 radiation damage in austenitic steels

4.02 Radiation Damage in Austenitic Steels F A Garner Radiation Effects Consulting, Richland, WA, USA ß 2012 Elsevier Ltd All rights reserved 4.02.1 4.02.2 4.02.2.1 4.02.2.2 4.02.3 4.02.4 4.02.5 4.02.6 4.02.7 4.02.8 4.02.8.1 4.02.8.2 4.02.8.3 4.02.8.3.1 4.02.8.3.2 4.02.8.3.3 4.02.8.3.4 4.02.8.3.5 4.02.9 4.02.9.1 4.02.9.2 4.02.9.3 4.02.9.4 4.02.9.5 4.02.9.5.1 4.02.9.5.2 4.02.9.5.3 4.02.9.6 4.02.9.7 4.02.9.8 4.02.10 References Introduction Basic Damage Processes Atomic Displacements Transmutation Differences in Neutron Spectra Transmutation Issues for Stainless Steels Evolution of Radiation-Induced Microchemistry and Microstructure A Cross-Over Issue Involving Radiation-Induced Microstructural Evolution and Transmutation Radiation-Induced Changes in Mechanical Properties Radiation-Induced Changes in Dimension Precipitation-Related Strains Void Swelling and Bubble Swelling Parametric Dependencies of Void Swelling Stress state Elemental composition Alloy starting state Irradiation temperature Influence of dpa rate on swelling Irradiation Creep Introduction Stages of Irradiation Creep Examples of Creep Behavior Creep Disappearance Recent Revisions in Understanding of Irradiation Creep Dependence of irradiation creep on dpa rate Dependence of creep and creep relaxation on neutron spectra Dependence of creep modulus on hydrostatic stress Stress Relaxation by Irradiation Creep Stress Rupture Fatigue Conclusions Abbreviations ATR BN-350 BN-600 BOR-60 Advanced Test Reactor in Idaho Falls, Idaho Russian acronym for Fast Neutron at 350 MW in Actau, Kazakhstan Russian acronym for Fast Neutron at 600 MW in Zarechney, Russia Russian acronym for Fast Experimental Reactor at 60 MW in Dimitrovgrad, Russia BR-2 BR-10 BWR CAGR CANDU DFR 34 35 35 37 37 40 44 49 50 61 62 65 67 67 68 69 69 70 74 74 78 79 79 83 83 84 85 86 88 89 90 91 Belgium Research Reactor-II in Mol, Belgium Russian acronym for Fast Reactor at 10 MW in Obninsk, Russia Boiling water reactor Commercial Advanced Gas Reactor Registered trademark for Canadian Deuterium Uranium Reactor Dounreay Fast Reactor in Dounreay, Scotland 33 34 Radiation Damage in Austenitic Steels DMTR EBR-II FFTF HFIR HFR IASCC IGSCC JMTR NRU ORNL ORR PWR T/F VVER Dounreay Materials Test Reactor in Dounreay, Scotland Experimental Breeder Reactor-II in Idaho Falls, Idaho Fast Flux Test Facility, fast reactor in Richland, WA High Flux Isotope Reactor at Oak Ridge National Laboratory High Flux Reactor in Petten, Netherlands Irradiation-assisted stress corrosion cracking Intergranular stress corrosion cracking Japan Material Testing Reactor in Oarai, Japan National Research Universal Reactor in Chalk River, Canada Oak Ridge National Laboratory: Oak Ridge Research Reactor in Oak Ridge, Tennessee Pressurized water reactor Thermal-to-fast neutron ratio Russian acronym for water-cooled, water moderated energetic reactor 4.02.1 Introduction Austenitic stainless steels are widely used as structural components in nuclear service in addition to being employed in many other nonnuclear engineering and technological applications The description of these steels and their as-fabricated properties is covered in Chapter 2.09, Properties of Austenitic Steels for Nuclear Reactor Applications This chapter describes the evolution of both microstructure and macroscopic property changes that occur when these steels are subjected not only to prolonged strenuous environments but also to the punishing effects of radiation While various nuclear environments involve mixtures of charged particles, high-energy photons and neutrons, it is the latter that usually exerts the strongest influence on the evolution of structural steels and thereby determines the lifetime and continued functionality of structural components To describe the response of austenitic stainless steels in all neutron environments is a challenging assignment, especially given the wide range of neutron spectra characteristic of various neutron devices This review of neutron-induced changes in properties and dimensions of austenitic stainless steels in all spectral environments has therefore been compiled from a series of other, more focused reviews directed toward particular reactor types1–8 and then augmented with material from a recently published textbook9 and journal articles It should be noted, however, that many of the behavioral characteristics of iron-based stainless steels following neutron irradiation are also observed in nickelbased alloys Whenever appropriate, the similarities between the two face-centered-cubic alloy systems will be highlighted A more comprehensive treatment of radiation effects in nickel-base alloys is provided in Chapter 4.04, Radiation Effects in Nickel-Based Alloys This review is confined to the effects of neutron exposure only on the response of irradiated steels and does not address the influence of charged particle irradiation While most of the phenomena induced by neutrons and charged particles are identical, there are additional processes occurring in charged particle studies that can strongly influence the results Examples of processes characteristic of charged particle simulations are the injected interstitial effect,10,11 strong surface effects,12,13 dose gradients,14,15 and atypical stress states.16,17 Chapter 1.07, Radiation Damage Using Ion Beams addresses the use of charged particles for irradiation Austenitic stainless steels used as fuel cladding or structural components in various reactor types must often withstand an exceptionally strenuous and challenging environment, even in the absence of neutron irradiation Depending on the particular reactor type, the inlet temperature during reactor operation can range from $50 to $370  C The maximum temperature can range from as high as 650 to 700  C for structural components in some reactor types, although most nonfueled stainless steel components reach maximum temperatures in the range of 400–550  C During operation, the steel must also withstand the corrosive action of fission products on some surfaces and flowing coolant on other surfaces The coolant especially may be corrosive to the steel under operating conditions Some of these environmental phenomena are synergized or enhanced by the effect of neutron irradiation Dependent on the nature of the component and the length of its exposure, there may also be significant levels of stress acting on the component Stress not only influences cracking and corrosion (see Chapter 5.08, Irradiation Assisted Stress Corrosion Cracking) but can also impact the dimensional stability of stainless steel, primarily due to Radiation Damage in Austenitic Steels thermal creep and irradiation creep, and also from the influence of stress on precipitation, phase stability, and void growth, some of which will be discussed later However, it will be shown that neutron irradiation can strongly affect both the microstructure and microchemistry of stainless steels and high-nickel alloys, with strong consequences on physical properties, mechanical properties, dimensional stability, and structural integrity Stainless steels are currently being used or have been used as structural materials in a variety of nuclear environments, most particularly in sodiumcooled fast reactors, water-cooled and water-moderated test reactors, water-cooled and water-moderated power reactors, with the latter subdivided into light water and heavy water types Additionally, there are reactor types involving the use of other coolants (helium, lithium, NaK, lead, lead–bismuth eutectic, mercury, molten salt, organic liquids, etc.) and other moderators such as graphite or beryllium The preceding reactor types are based on the fission of uranium and/or plutonium, producing neutron energy distributions peaking at $2 MeV prior to moderation and leakage effects that produce the operating spectrum However, there are more energetic sources of neutrons in fusion-derived spectra, with the source peaking at $14 MeV and especially from spallation events occurring at energies of hundreds of MeV, although most spallation spectra are mixtures of high-energy protons and neutrons It is important to note that in each of these various reactors, there are not only significant differences in neutron flux-spectra but also significant differences in neutron fluence experienced by structural components These differences in fluence arise not only from differences in neutron flux characteristic of the different reactor types but also the location of the steel relative to the core For instance, boiling water reactors and pressurized water reactors have similar in-core spectra, but stainless steels in boiling water reactors are located much farther from the core, resulting in a factor of reduction of $20 in both neutron dose rate and accumulated dose compared to steels in pressurized water reactors 4.02.2 Basic Damage Processes 4.02.2.1 Atomic Displacements What are the nature and origins of neutron-induced phenomena in metals? The major underlying driving 35 force arises primarily from neutron collisions with atoms in a crystalline metal matrix When exposed to displacive irradiation by energetic neutrons, the atoms in a metal experience a transfer of energy, which if larger than several tens of eV, can lead to displacement of the atom from its crystalline position The displacements can be in the form of single displacements resulting from a low-energy neutron collision with a single atom or a glancing collision with a higher energy neutron More frequently, however, the ‘primary knock-on’ collision involves a larger energy transfer and there occurs a localized ‘cascade’ of defects that result from subsequent atom-to-atom collisions There are several other contributions to displacement of atoms from their lattice site, but these are usually of second-order importance The first of these processes involve production of energetic electrons produced by high-energy photons via the photoelectric effect, Compton Effect, or pair production.18 These electrons can then cause atomic displacements, but at a much lower efficiency than that associated with neutron-scattering events The second type of process involves neutron absorption by an atom, its subsequent transmutation or excitation, followed by gamma emission The emission-induced recoil of the resulting isotope often is sufficient to displace one or several atoms In general, however, such recoils add a maximum of only several percent to the displacement process and only then in highly thermalized neutron spectra.4 One very significant exception to this generalization involving nickel will be presented later For structural components of various types of nuclear reactors, it is the convention to express the accumulated damage exposure in terms of the calculated number of times, on the average, that each atom has been displaced from its lattice site Thus, 10 dpa (displacements per atom) means that each atom has been displaced an average of 10 times Doses in the order of 100–200 dpa can be accumulated over the lifetimes of some reactor components in various high-flux reactor types The dpa concept is very useful in that it divorces the damage process from the details of the neutron spectrum, allowing comparison of data generated in various spectra, providing that the damage mechanism arises primarily from displacements and not from transmutation The use of the dpa concept also relieves researchers from the use of relatively artificial and sometimes confusing threshold energies frequently used to describe the damage-causing portion of the neutron spectrum Neutrons with ‘energies greater than 36 Radiation Damage in Austenitic Steels X MeV,’ where X is most frequently 0.0, 0.1, 0.5, or 1.0 MeV, have been used for different reactor concepts at different times in history The threshold energy of 0.1 MeV is currently the most widely used value and is most applicable to fast reactors where large fractions of the spectra lay below 0.5 and 1.0 MeV Many older studies employed the total neutron flux (E > 0.0) but this is the least useful threshold for most correlation efforts Caution should be exercised when compiling data from many older studies where the neutron flux was not adequately identified in terms of the threshold energy employed There are rough conversion factors for ‘displacement effectiveness’ for 300 series austenitic steels that are useful for estimating dpa from >0.1 MeV fluences for both in-core or near-core spectra in most fission spectra Examples are $7 dpa per 1022 n cmÀ2 (E > 0.1) for most in-core light water spectra with lower in-core values of $5 dpa per 1022 n cmÀ2 (E > 0.1) for metal fueled fast reactors and $4 dpa per 1022 n cmÀ2 (E > 0.1) for oxide-fueled fast reactors.4 Such conversion factors should not be trusted within more than (10–15%), primarily due to spatial variations across the core resulting from neutron leakage For fast reactor spectra, E > 1.0 conversion factors are completely unreliable When E > 1.0 fluxes are employed in light water reactor studies, the conversion factor increases from $7 dpa per 1022 n cmÀ2 (E > 0.1) to $14 dpa per 1022 n cmÀ2 (E > 1.0) In Russia, a threshold energy of >0.5 MeV is popular for light water reactors with $9 dpa per 1022 n cmÀ2 (E > 0.5) All of these conversion factors assume that within several percent pure iron is a good surrogate for 300 series alloys Note that other metals such as Cu, Al, W, etc will have different conversion values arising from different displacement threshold energies and sometimes different displacement contributions A standard procedure for calculating dpa has been published,19 although other definitions of dpa were used prior to international acceptance of the ‘NRT model’ where the letters represent the first letter of the three author’s last name (see Garner1 for details on earlier models) Caution must be exercised when compiling doses from older studies where displacement doses were calculated using other models (Kinchin-Pease, Half-Nelson, French dpa, etc.) sometimes without clearly identifying the model employed Conversion factors between the NRT model and various older models of dpa are provided in Garner,1 but all models agree within $23% While sometimes controversial with respect to how far the dpa concept can be stretched to cover the full range of spectral differences for neutron and especially for charged particle environments, it appears that the dpa concept is very efficient to stretch over light water, heavy water, fusion, and spallation spectra, providing that all energy deposition and displacement processes are included Note in Figure how well the dpa concept collapses the data on neutron-induced strengthening of stainless steel into one response function for three very different spectra (light water fission, pure D–T fusion and ‘beam-stop’ spallation).20 300 LASREF, 40 ЊC RTNS-II, 90 ЊC OWR, 90 ЊC 250 Yield stress change (MPa) Yield stress change (MPa) 300 200 150 100 50 1017 1018 1019 Neutron fluence, E > 0.1 MeV 1020 250 LASREF, 40 ЊC RTNS-II, 90 ЊC OWR, 90 ЊC 200 150 100 50 10-3 10-2 dpa Figure Radiation-induced yield stress changes of 316 stainless steel versus (left) neutron fluence (n cmÀ2 E > 0.1 MeV), and (right) displacements per atom Reproduced from Heinisch, H L.; Hamilton, M L.; Sommer, W F.; Ferguson, P J Nucl Mater 1992, 191–194, 1177, as modified by Greenwood, L R J Nucl Mater 1994, 216, 29–44 Radiation Damage in Austenitic Steels 4.02.2.2 Transmutation It is important to note that material modification by radiation arises from two primary spectral-related processes In addition to the neutron-induced displacement of atoms there can be a chemical and/or isotopic alteration of the steel via transmutation With the exception of helium production, transmutation in general has been ignored as being a significant contributor to property changes of stainless steels and nickel-base alloys In this chapter, transmutation is shown to be sometimes much more important than previously assumed Both the displacement and transmutation processes are sensitive to the details of the neutron flux-spectra, and under some conditions each can synergistically and strongly impact the properties of the steel during irradiation In addition to the brief summary presented below on flux-spectra issues relevant to stainless steels, the reader is referred to various papers on transmutation and its consequences in different reactor spectra.5–8,18,21–23 Transmutation may be subdivided into four categories of transmutants Three of these are relevant to fission-derived or fusion-derived spectra, and the fourth is associated with spallation-derived spectra The first three are solid transmutants, gaseous transmutants, and ‘isotope shifts,’ the latter involving production of other isotopes of the same element While the latter does not change the chemical composition of stainless steels, it is an underappreciated effect that is particularly relevant to nickel-containing alloys such as stainless steels and nickel-base alloys when irradiated in highly thermalized neutron spectra Whereas the first three categories arise from discrete nuclear reactions to produce discrete isotopes of specific elements, the spallation-induced transmutation arising in accelerator-driven devices involves a continuous distribution of every conceivable fragment of the spalled atom, producing every element below that of the target atom across a wide range of isotopes for each element While individual solid transmutants in spallation spectra are usually produced at levels that not change the alloy composition significantly, the very wide range of elements produced allows the possibility that deleterious impurities not normally found in the original steel may impact its continued viability This possibility has not received sufficient attention and should be examined further if spallation devices continue to be developed Another consequence of spallation-relevant transmutation is that the induced radioactivity per unit 37 mass is correspondingly much higher than that produced per dpa in other spectra The majority of the spalled fragments and their daughters/granddaughters are radioactive with relatively short half-lives, leading to materials that are often much more difficult to examine than materials irradiated in fission spectra Most importantly, there is a very strong production of hydrogen and helium in spallation spectra at levels that are one or two orders of magnitude greater than produced in most fission or fusion spectra.5,6,21 While there is a tendency to view displacement and transmutation processes as separate processes, it will be shown later that under some circumstances the two processes are strongly linked and therefore inseparable in their action to change alloy behavior 4.02.3 Differences in Neutron Spectra There are significant differences in neutron spectra for water-cooled, sodium-cooled, and other types of fission-based reactors It should be noted that there is a conventional but slightly misleading practice to differentiate between ‘fast’ and ‘thermal’ reactors Thermal reactors have a significant portion of their spectra composed of thermal neutrons Thermalized neutrons have suffered enough collisions with the moderator material that they are in thermal equilibrium with the vibrations of the surrounding atoms Efficient thermalization requires low-Z materials such as H, D, and C in the form of water, graphite, or hydrocarbons At room temperature the mean energy of thermalized neutrons is 0.023 eV The designation ‘fast’ reactor, as compared to ‘thermal’ reactor, refers to the portion of the neutron spectrum used to control the kinetics of ascent to full power for each type of reactor As shown later, this practice incorrectly implies to many that fast reactors have ‘harder’ neutron spectra than ‘softer’ thermal reactors Actually, the opposite is true Examples of typical flux-spectral differences in fission-based reactors are shown in Figures 2–5 The local spectrum at any position is determined primarily by the fuel (U, Pu) and fuel type (metal, oxide, carbide, etc.), the coolant identity and density, the local balance of fuel/coolant/metal as well as the proximity to control rods, water traps, or core boundaries Additionally, it is possible to modify the neutron spectra in a given irradiation capsule by including in it 38 Radiation Damage in Austenitic Steels 1.E + 16 1015 HFIR Flux/lethargy Flux per unit lethargy HFIR-PTP 1.E + 14 1014 ORR HFIR-RB* 1.E + 12 ATR-ITV 1.E + 10 1013 EBRII FFTF 1.E + 08 1.E - EBR II 1012 10-8 10-7 10-6 10-5 10-4 10-3 10-2 10-1 100 Neutron energy (MeV) 101 102 Figure Difference in neutron flux-spectra of two water-cooled test reactors (high-flux HFIR and lower-flux ORR) and one high-flux sodium-cooled fast reactor (EBR-II) 1014 T/F ~0.15 1.E - 1.E - 1.E - 1.E - Neutron energy (MeV) 1.E + Figure Comparison of flux-spectra in various test reactors Note that FFTF is softer in spectrum compared to EBR-II due to the use of oxide fuel rather than metal fuel Neither fast reactor has measurable fluxes of thermal neutrons In the PTP position of HFIR a water trap strongly contributes to a high thermal-to-fast ratio, while in the RB* (removable beryllium) position the predominance of Be over water reduces the thermal population In the ATR position where the ITV assembly was located, the use of strong absorber sleeves strongly depressed the thermal flux Flux per unit lethargy 1013 Baffle bolt Top of bolt head 12 10 1011 Upper core plate 10 10 109 10-8 10-7 10-6 10-5 10-4 10-3 10-2 10-1 100 Neutron energy (MeV) 101 102 Figure Typical neutron flux-spectra of internal components of a pressurized water reactor, having a thermal-to-fast neutron ratio smaller by factors of 10–20 than that of typical light water test reactors Reproduced from Garner, F A.; Greenwood, L R In 11th International Conference on Environmental Degradation of Materials in Nuclear Power Systems – Water Reactors; 2003; pp 887–909 or enclosing it with a moderator or absorber Metal hydrides are used in fast reactors to soften the spectrum, while in mixed-spectrum reactors the thermalto-fast ratio can be strongly reduced by incorporating elements such as B, Hf, Gd, and Eu The most pronounced influence on neutron spectra in fission reactors arises from the choices of coolant and moderator, which are often the same material (e.g., water) Moving from heavy liquid metals such as lead or lead–bismuth to lighter metals such as sodium leads to less energetic or ‘softer’ spectra Use of light water for cooling serves as a much more effective moderator Counterintuitively, however, this leads to both more energetic and less energetic spectra at the same time, producing a twopeaked ‘fast’ and ‘thermal’ distribution separated by a wide energy gulf at lower fluxes Such two-peaked spectra are frequently called ‘mixed spectra.’ The ratio of the thermal and fast neutron fluxes in and near such reactors can vary significantly with position and also with time.4 Using heavy water, we obtain a somewhat less efficient moderator that does not absorb neutrons as easily as light water, but one that produces an even more pronounced two-peak spectral distribution where the thermal-to-fast neutron ratio can be very large These spectral differences lead to strong variations between various reactors in the neutron’s ability to displace atoms and to cause transmutation Depending on the reactor size and its construction details there can also be significant variations in neutron spectra and ‘displacement effectiveness’ within a given reactor and its environs, especially where more energetic neutrons can leak out of the core Examples of these variations of displacement effectiveness for fast reactors are shown in Figures and Compared to fission-derived spectra, there are even larger spectral differences in various fusion or spallation neutron devices The reader should note the emphasis placed here on flux-spectra rather than simply spectra If we focus only on light water-cooled reactors for example, there are in general three regimes of neutron flux of relevance to this review First, there are the relatively low fluxes typical of many experimental reactors that Radiation Damage in Austenitic Steels 1016 Thermal flux, clean core Thermal flux, 21-day core Total nonthermal flux >0.111 MeV >0.821 MeV 1015 1013 12 H2O Permanent beryllium Removable beryllium Control region Outer fuel annulus H2O annulus Inner fuel annulus H2O outer annulus 1014 300-g Pu target Neutronflux (neutrons per cm2 s–1) 39 16 20 24 28 32 36 40 44 Radial distance from core center (cm) 48 52 56 60 Figure Variation in fast and thermal fluxes in HFIR as a function of radial position at mid-core at 85 MW, also showing change in thermal population with burn-up (Source: ORNL website) can produce doses of 10 dpa or less over a decade Second, there are moderate flux reactors that are used to produce power that can introduce doses as high as 60–100 dpa maximum over a 30–40 year lifetime and finally, some high-flux thermal reactors that can produce 10–15 dpa yearÀ1 in stainless steels Most importantly, fast reactors also operate in the high-flux regime, producing 10–40 dpa yearÀ1 Therefore, the largest amount of published highdpa data on stainless steels has been generated in fast reactors Some phenomena observed at high exposure, such as void swelling, have been found to be exceptionally sensitive to the dpa rate, while others are less sensitive (change in yield strength) or essentially insensitive (irradiation creep) These sensitivities will be covered in later sections For light water-cooled reactors, the various flux regimes need not necessarily involve large differences in neutron spectra, but only in flux However, the very large dpa rates characteristic of fast reactors are associated with a significant difference in spectrum This difference is a direct consequence of the fact that fast reactors were originally designed to breed the fissionable isotope 239Pu from the relatively nonfissile isotope 238U, which comprises 99.3% of natural uranium In order to maximize the breeding of 239Pu, it is necessary to minimize the unproductive capture of neutrons by elements other than uranium One 5.5 Row 5.0 dpa 1022 (E > 0.1) Row 4.5 4.0 -20 -10 10 20 Axial position (cm) Figure Displacement effectiveness values of dpa per 1022 n cmÀ2 (E > 0.1 MeV) across the small core (30 cm tall and $30 cm diameter) of the EBR-II fast reactor, showing effects of neutron leakage to soften the spectrum near the core axial boundaries Near core center (Row 2) the spectrum and displacement effectiveness are dictated primarily by the use of metal fuel, producing a maximum of $5.2 dpa per 1022 n cmÀ2 (E > 0.1 MeV) In mid-core Row the radial leakage is just becoming significant strategy used to accomplish this goal is to avoid thermalization of the reactor neutrons, which requires that no low atomic weight materials such as H2O, D2O, Be, or graphite be used as coolants or as moderators For this purpose, sodium is an excellent coolant with a moderate atomic weight The use of sodium results 40 Radiation Damage in Austenitic Steels 6.0 BC FFTF core 5.0 Above core FFTF cycles and dpa 1022 (E > 0.1) FFTF cycle 10 4.0 3.0 −100 −75 −50 −25 25 50 75 100 125 150 Distance from core midplane (cm) Figure Values of dpa per 1022 n cmÀ2 (E > 0.1 MeV) across the much larger core of FFTF for two different fuel/experiment loadings, showing a lesser effect of neutron leakage in larger cores Note, however, that the in-core values are less than the in-core values of EBR-II, reflecting the softer spectra arising from the use of oxide fuel Far from the core the displacement effectiveness values are lower, determined primarily by the absence of fuel and the balance of sodium and steel in a neutron spectrum that is nominally single-peaked rather than the typical double-peaked (thermal and fast) neutron spectrum found in light water or heavy water reactors The single-peaked fast reactor spectrum is significantly less energetic or softer, however, than that found in the fast peak of light water reactors Depending on the fuel type (metal vs oxide) the mean energy of fast reactor spectra varies from $0.8 to $0.5–0.4 MeV while light water-cooled reactors have a fast neutron peak near $1.2 MeV One consequence of attaining successful breeding conditions is that the spectrum-averaged crosssection for fission is reduced by a factor of 300–400 relative to that found in light water spectra To reach a power density comparable to that of a light water power-producing reactor, the fast reactor utilizes two concurrent strategies: increases in fissile enrichment to levels in the order of 20% or more, and most importantly, an increase in neutron flux by one or two orders of magnitude Thus, for a given power density, the fast reactor will subject its structural materials to the punishing effects of neutron bombardment at a rate that is several orders of magnitude greater than that in light water reactors At the same time, however, the softer ‘fast’ spectrum without thermalized neutrons leads to a significant reduction in transmutation compared to typical light water spectra, at least for stainless steels and nickel-base steels 4.02.4 Transmutation Issues for Stainless Steels For most, but not all fission-derived spectra, stainless steels are relatively immune to transmutation, especially when compared to other elements such as aluminum, copper, silver, gold, vanadium, tungsten, and rhenium,5,21,24–27 each of which can rapidly become two or three component alloys via transmutation arising from thermal or epithermal neutrons Whereas the properties of these metals are particularly sensitive to formation of solid transmutation products, stainless steels in general not change their composition by significant amounts compared to preexisting levels of impurities, but significant amounts of helium and hydrogen can be produced in fission-derived spectra, however In stainless steels the primary transmutant changes that arise in various fission and fusion reactor spectra involve the loss of manganese to form iron, loss of chromium to form vanadium, conversion of boron to lithium and helium, and formation of helium and hydrogen gas.4,28 While each of these changes in solid or gaseous elements are produced at relatively small concentrations, they can impact the evolution of alloy properties and behavior For instance, vanadium is not a starting component of most 300 series stainless steels, but when included it participates in the formation of carbide Radiation Damage in Austenitic Steels with the major alloy components This type of reaction occurs only above high neutron threshold energies (>6 MeV) Figure shows that nickel is the major contributor to helium production by (n, a) reactions,36 and thus the helium generation rate scales almost directly with nickel content for a large number of commercial steels A similar behavior occurs for production of hydrogen by transmutation via high-energy neutrons, where nickel is also the major source of hydrogen compared to other elements in the steel.4,7 In this case, the threshold energy is around MeV with 58 Ni being the major contributor This generality concerning nickel as the major source of He and H is preserved in more energetic fusion-derived spectra, although the He/dpa and H/dpa generation rates in fusion spectra are much larger than those of fast reactor spectra When moving to very energetic spallation-derived neutron and proton spectra, however, the observation that nickel accounts for most of the helium and hydrogen is no longer correct Iron, nickel, chromium, cobalt, and copper produce essentially the same amounts of helium and hydrogen for energies above $100 MeV as shown in Figure 9.6 Another very important helium-generation process also involves nickel Helium is produced via the two-step 58Ni(n, g)59Ni(n, a)56Fe reaction sequence.37,38 This sequence operates very strongly in mixed-spectrum reactors 59Ni is not a naturally occurring isotope and is produced from 58Ni Thus, this helium contribution involves a delay relative to 0.14 Cross-section (barns) precipitates that change the distribution and chemical activity of carbon in the alloy matrix Carbon plays a number of important roles in the evolution of microstructure1 and especially in grain boundary composition The latter consideration is very important in determining the grain boundary cracking behavior, designated irradiation-assisted stress corrosion cracking (IASCC), especially with respect to the sensitization process.29 The strong loss of manganese in highly thermalized neutron spectra has been suggested to degrade the stability of insoluble MnS precipitates that tie up S, Cl, and F, all of which are elements implicated in grain boundary cracking.30 Late-term radiationinduced release of these impurities to grain boundaries may participate in cracking, but this possibility has not yet been conclusively demonstrated In some high-manganese alloys such as XM-19 manganese serves to enhance the solubility of nitrogen which serves as a very efficient matrix strengthener In highly thermalized spectra the loss of manganese via transmutation has been proposed to possibly lead to a decrease in the strength of the alloy and perhaps to induce a release of nitrogen from solution to form bubbles.31 The overwhelming majority of published transmutation studies for stainless steels and high-nickel alloys steels have addressed the effects of He/dpa ratio on mechanical properties and dimensional instabilities Much less attention has been paid to the effect of H/dpa ratio based on the long-standing perception that hydrogen is very mobile in metals and therefore is not easily retained in steels at reactor-relevant temperatures As presented later, this perception is now known to be incorrect, especially for water-cooled reactors The focus of most published studies concerned the much higher helium generation rates anticipated in fusion spectra ($3–10 appm He/dpa) compared to the lower rates found in fast reactors ($0.1–0.3 appm He/dpa).32 It was later realized that in some highly thermalized test reactors, such as HFIR, very large generation rates could be reached ($100 appm He/dpa), and even in pressurized water reactors the rate could be very high ($15 appm He/dpa).33 In heavy water reactors the rate can be much larger, especially in out-of-core regions.34,35 While some helium arises from (n, a) reactions with thermal and epithermal neutrons interacting with the small amounts of boron found in most stainless steels, the major contribution comes initially from high-energy threshold-type (n, a) reactions 41 0.12 Ni 0.10 0.08 0.06 Cr Ti 0.04 Fe 0.02 Energy (MeV) 10 20 Figure Cross-sections for (n, a) reactions as a function of neutron energy for common elements used in stainless steels Reproduced from Mansur, L K.; Grossbeck, M L J Nucl Mater 1988, 155–157, 130–147 Nickel dominates the production of helium at higher neutron energies 42 Radiation Damage in Austenitic Steels 2500 Inconel 304L 316L 9Cr–1Mo Fe Co Ni Cu 1500 1000 1.6 500 0 10 15 Ratio to initial value He (appm) 2000 60 Ni Natural nickel 58Ni 67.85% 60Ni 26.2% 1.2 61Ni 58 Ni 0.8 62Ni 64Ni 6.1% total dpa Figure Measured amount of helium in alloys and pure metals that were irradiated by a mixed spectrum of high energy neutrons and protons produced by 800 MeV proton irradiation of tungsten rods There is some significant uncertainty in the dpa assignment for Inconel 718 at the highest dose Otherwise the He/dpa ratio appears to be independent of composition Reproduced from Garner, F A.; Oliver, B M.; Greenwood, L R.; James, M R.; Ferguson, P D.; Maloy, S A.; Sommer, W F J Nucl Mater 2001, 296, 66–82 that of single-step threshold (n, a) reactions Since both steps of the sequence involve cross-sections that increase with decreasing energy and the second step exhibits a resonance at 203 eV, the generation rate per dpa in fast reactors increases near the core boundaries and out-of-core areas It is in thermalized neutron spectra characteristic of light and heavy water-cooled reactors, however, where the 59Ni(n, a) reaction can produce He/dpa generation rates that are significantly larger than those characteristic of fusion-derived spectra Nickel has five naturally occurring stable isotopes with 58Ni comprising 67.8% natural abundance, 60Ni comprising 26.2%, and $6.1% total of 61Ni, 62Ni, and 64 Ni There is no natural 59Ni or 63Ni at the beginning of radiation During irradiation in a highly thermalized neutron spectrum, all nickel isotopes are strongly transmuted, primarily to the next higher isotopic number of nickel 59Ni has a half-life of 76 000 years and is progressively transmuted to 60Ni while 58Ni is continuously reduced in concentration Therefore, the 59Ni concentration rises to a peak at a thermal neutron fluence of  1022 n cmÀ2 where the 59/58 ratio peaks at $0.04 and then declines, as shown in Figure 10 This transmutation sequence in nickel is an example of the isotopic shift category of transmutation defined earlier For other elements used to make stainless steels, there are no consequences to such a shift since the total amount of the element is unchanged 0.4 0.0 1021 59 Ni 1022 1023 Thermal fluence (n cm-2) 1024 Figure 10 Transmutation-induced evolution of three nickel isotopes during irradiation in thermalized neutron spectra Reproduced from Garner, F A.; Greenwood, L R In 11th International Conference on Environmental Degradation of Materials in Nuclear Power Systems – Water Reactors; 2003; pp 887–909 Reproduced from Garner, F A.; Griffiths, M.; Greenwood, L R.; Gilbert, E R In Proceedings of the 14th International Conference on Environmental Degradation of Materials in Nuclear Power Systems – Water Reactors; American Nuclear Society, 2010; pp 1344–1354 and isotope shifts induce no significant consequences However, in the case of nickel there is an intimate linkage between the displacement and transmutation processes that arises from the isotope shift The recoil of the 59Ni upon emission of the gamma ray produces only about five displacements per event, and usually is not a significant addition to the displacement dose However, the isotope 59Ni undergoes three strong reactions with thermal and resonance ($0.2 keV) neutrons, two of which are exceptionally exothermic and can significantly add to the dpa level These reactions, in order of highest-to-lowest thermal cross-section, are (n, g) to produce 60Ni, followed by (n, a) and (n, p) to produce helium and hydrogen, respectively Even at relatively low thermal-to-fast neutron ratios, the reaction sequence can produce significant amounts of helium For example, He/dpa ratios in the order of $3–8 appm dpaÀ1 can be experienced along the length of a 316 stainless baffle bolt in the baffle-former assembly of a pressurized water Radiation Damage in Austenitic Steels 81 316Ti DV (%) V0 Creep coefficient B Swelling ϫ 10−6 MPa−1 dpa−1 C = 0.046 wt% 400–420 ЊC 450–460 ЊC C = 0.006 wt% 316Ti+P 12 ϫ 1026 −2 Fluence (n m ) (E > 0.1 MeV) Figure 71 Acceleration of irradiation creep in two carbon variants of a stainless steel by a low rate of swelling at 350 and 420  C Reproduced from Neustroev, V S.; Shamardin, V K In Effects of Radiation on Materials: 16th International Symposium; 1993; pp 816–823 The lower carbon steel has a longer transient regime of swelling The height of the plateau is determined by the swelling rate B0 was determined to be $1  10À6 (MPa dpa)À1 and D to be 0.6  10À2 MPaÀ1 400–420 ЊC Irradiation creep Creep strain (%) 170 MPa 450–470 ЊC 400–420 ЊC 450–470 ЊC 90 MPa 0 330 ЊC 400 ЊC 2.0 e (%) 500 ЊC 600 ЊC 13.1 dpa 20% CW 316 20 40 60 dpa 80 100 120 Figure 73 Swelling and creep strains observed in two French steels irradiated as pressurized tubes in PHENIX, showing strong correlation between the two types of strain as the swelling rate increases Reproduced from Dubuisson, P.; Maillard, A.; Delalande, C.; Gilbon, D D.; Seran, J L In Effects of Radiation on Materials: 15th International Symposium; STP 1125; 1992; pp 995–1014 12.0 dpa 1.0 BO = 2.8 ϫ 10−6 MPa−1 dpa−1 0.0 300 ЊC 400 ЊC 2.0 500 ЊC e (%) 600 ЊC 13.3 dpa 25% CW PCA 12.1 dpa 1.0 BO = 3.2 ϫ 10−6 MPa−1 dpa−1 0.0 100 200 300 400 500 Effective stress (MPa) Figure 72 Temperature-independent creep strains observed in 20% cold-worked 316 and 25% cold-worked PCA during irradiation in the ORR test reactor at a high He/ dpa ratio Reproduced from Grossbeck, M L.; Horak, J A J Nucl Mater 1988, 155–157, 1001–1005 Note that the two steels have very similar values of creep modulus B and are independent of irradiation temperature over a wide range The creep modulus B is about three times that of B0 ¼  10À6 (MPa dpa)À1, however, probably arising from observed high densities of helium bubbles to produce bubble-enhanced creep range of swelling behavior some unusual behaviors are often observed An example is shown in Figure 76 where the two-peaked swelling behavior frequently observed in 300 series steels is mirrored in the creep strains, but the relative proportions of the two strains are distorted.172 This is one manifestation of the creep disappearance phenomenon in which the attainment of significant swelling causes irradiation creep to strongly drop in rate or even to disappear under some conditions as seen in Figures 77 and 78 In early fuel pin studies it was often observed that irradiation creep strains would increase and then abruptly decrease and sometimes stop entirely, even though fission gas pressures continued to increase.173,174 These results were interpreted as evidence of fuel swelling very quickly to meet and thereby put stress on the cladding but later the onset of swelling in the clad caused it to out-swell the fuel and break contact Actually, the driving force 82 Radiation Damage in Austenitic Steels 2.5 MPa A094, T-415 ЊC 60 MPa Midwall creep strain/hoop stress (% per MPa) Swelling-induced diametral strain (%) C42, T-415 ЊC C38, T-390 ЊC 2.0 C39, T-390 ЊC C40, T-390 ЊC C44, T-390 ЊC 1.5 83508, T-420 ЊC 1.0 K280, T-395 ЊC A095, T-415 ЊC 0.5 83508 100 MPa 0.05 140 MPa 200 MPa Failed in next cycle 300 MPa 0.04 K280 0.03 0.02 A095 0.01 Failed in next cycle 0 100 50 150 dpa 50 100 150 dpa Figure 74 (left) Diametral strains resulting from void swelling at 400  C in neutron-irradiated stress-free tubes constructed from nine titanium-modified 316 stainless steels, (right) Stress-normalized midwall creep strains observed in three of these steels, showing a strong correlation of swelling and irradiation creep rates in each steel Reproduced from Toloczko, M B.; Garner, F A.; Eiholzer, C R J Nucl Mater 1992, 191–194, 803–807 16 BEQϫ10-6 (MPa dpa)-1 14 56 dpa 12 10 10 dpa 6.3 dpa 0 10 20 30 Ni-equivalent (%) 40 50 Figure 75 Creep modulus measured for six austenitic steels irradiated in BOR-60 fast reactor at 420  C, showing an enhancement of creep versus Ni-equivalent Reproduced from Neustroev, V S.; Shamardin, V K J Nucl Mater 2002, 307–311, 343–346 This behavior corresponds to the known effect of nickel on void swelling, indicating swelling-enhanced creep was primarily increasing levels of fission gas but irradiation creep had disappeared by ~7% burn-up Several features of creep disappearance are noteworthy The combined creep and swelling strain rate in a fuel pin or pressurized tube cannot exceed 0.33% per dpa or one-third of the eventual steady-state swelling rate As swelling approaches 1% per dpa the creep rate backs down proportionately to maintain this maximum rate as shown in Figures 78–80 The limit of 0.33% per dpa is reached before swelling gets to a significant fraction of 1% per dpa, as shown in Figure 80 Some tubes had already reached the maximum strain rate limit, but then lost their gas pressure and continued to swell at less than 1% per dpa As the creep cessation process gets underway the creep strain loses its responsiveness to the magnitude of the stress Note in Figures 79 and 80 that doubling the hoop stress did not double the strain rate in the tube The coupling coefficient D tends to fall to zero rather quickly when swelling-before-creep occurs but falls more slowly in creep-before-swelling scenarios (fuel pins vs pressurized tubes).175 A consensus explanation of the creep disappearance phenomena has not yet been reached Various models have been proposed involving voids acting to erase the anisotropy of dislocation Burgers vector176,177 and the involvement of precipitate sinks to serve as strong sinks that compete with dislocations.175 Radiation Damage in Austenitic Steels 83 Pin 32 Swelling (%) Swelling (%) Pin 31 Creep modulus (MPa dpa F)-1 Creep strains (%) Pin 47 Pin 32 1.2 0.8 Pin 31 0.4 Bottom Pin Top 6´10-6 Pin Pin 47 400 Fuel column length 500 600 Temperature (ºC) 700 Figure 76 Swelling and creep behavior observed along the length of AISI 316 fuel pins irradiated in the RAPSODIE fast reactor; (left) solution annealed and (right) 20% cold-worked Reproduced from Boutard, J L.; Carteret, Y.; Cauvin, R.; Guerin, Y.; Maillard, A In Proceedings Conference on Dimensional Stability and Mechanical Behavior of Irradiated Metals and Alloys; British Nuclear Energy Society: London, 1983; pp 109–112 30 Onset of creep disappearance Instantaneous 20 creep coefficient 1030 psi n cm-2 10 Swelling-enhanced creep 10 20 Swelling in the absence of creep 30 dpa 40 50 60 Figure 77 Instantaneous creep coefficient B derived from strain measurements on pressurized tubes constructed from a double-aged higher-swelling condition of 316 stainless steel irradiated in EBRII at 550  C Reproduced from Porter, D L.; Garner, F A J Nucl Mater 1988, 159, 114–121 4.02.9.5 Recent Revisions in Understanding of Irradiation Creep 4.02.9.5.1 Dependence of irradiation creep on dpa rate As mentioned earlier, once swelling begins, irradiation creep quickly assumes all the parametric dependencies of void swelling However, for many years it was assumed that the B0 component of creep was also strongly dependent on dpa rate, increasing as the dpa rate fell, as shown in Figure 81 The original research that established this perception was performed by Lewthwaite and Mosedale on various cold-worked steels in the Dounreay Fast Reactor at temperatures in the 270–350  C range.178 84 Radiation Damage in Austenitic Steels 10 30 ksi 487-543 ЊC 15 ksi 0.33% per dpa Total Diameter change (%) DD (%) D Swelling deformation Plastic deformation Hoop stress = ksi -2 (a) 30 ksi Irradiation creep 0.33% per dpa Stressfree swelling 0 10 12 Local atomic burnup (%) Figure 78 Creep and swelling strains observed in annealed 347 stainless clad fuel pins irradiated in EBR-II, showing the disappearance of further creep strain as irradiation continues Reproduced from Appleby, W K.; Hilbert, R F.; Bailey, R W In Proceedings Conference on Irradiation Embrittlement and Creep in Fuel Cladding and Core Components; British Nuclear Energy Society: London, 1972, pp 209–216 These data were originally explained in terms of fuel-clad interaction acting as the major source of stress in the cladding, with fuel contact and stress-driven creep eventually terminated by the onset of clad swelling to move the clad away from the fuel Continually increasing gas loading was actually the primary loading on the cladding, not the fuel The explanation advanced for such a dependence was the decreasing amount of annihilation of point defects by recombination at lower dpa rates, where such an effect is expected to be more pronounced at the lower irradiation temperatures characteristic of this experiment An earlier review article was published where this and other data sets were assessed to determine the appropriate rate dependence.1 Some data sets available at that time supported a flux dependence and other data sets supported an independence of dpa rate On balance it appeared that a strong dependence of irradiation creep rate on dpa rate was the more defendable conclusion With hindsight and additional published data supporting the opposite conclusion, it was later realized that apparent dependence of creep rate on dpa rate was an artifact of the analysis procedure used by Stress-affected swelling at 30 ksi -2 (b) 20 40 60 80 100 dpa Figure 79 (a) Deformation observed in pressurized tubes of 20% cold-worked AISI 316 irradiated in EBR-II at 550  C Reproduced from Porter, D L.; Garner, F A J Nucl Mater 1988, 159, 114–121; Porter, D L.; Garner, F A In Effects of Radiation on Materials: 13 International Symposium (Part II) Influence of Radiation on Material Properties; ASTM STP 956; 1987; pp 11–21 Note that doubling the hoop stress from (from 15 to 30 ksi: 103 to 206 MPa) does not double the deformation rate, which never exceeds 0.33% per dpa (b) Density measurements on the 30 ksi (206 MPa) tube show that stress accelerates the rate of swelling, but also causes the creep rate to approach zero at high swelling levels Mosedale and Lewthwaite The authors had not properly separated the transient and post-transient strains, and all of the lower flux data were in the higher-rate transient regime When the DFR creep data were reanalyzed by Garner and Toloczko, the creep compliance B0 was found to be independent of dpa rate.179 4.02.9.5.2 Dependence of creep and creep relaxation on neutron spectra It is sometimes assumed that thermalized neutron spectra can produce more effectively surviving point defects since gamma-recoil events not produce cascades and therefore there is less in-cascade annihilation Thus, a larger fraction of thermally produced defects are postulated to survive to contribute to creep or embrittlement.180,181 Radiation Damage in Austenitic Steels 550 ЊC Core 12 550 ЊC Core 0.33% per dpa MPa 35 MPa 70 MPa 117 MPa 163 MPa 233 MPa DD (%) D 85 575 ЊC Core 12 575 ЊC Core MPa 16 MPa 31 MPa 63 MPa 104 MPa 146 MPa 0 20 40 60 dpa 20 40 60 80 Figure 80 Diametral strains observed in two related heats of 20% cold-worked AISI 36 irradiated in FFTF as pressurized tubes Reproduced from Garner, F A.; Toloczko, M B.; Puigh, R J In Effects of Radiation on Materials: 19th International Symposium; ASTM STP 1366; 2000; pp 667–678 Note that many of the tubes have reached the limiting deformation rate of 0.33% per dpa Those tubes which subsequently fail show that swelling had not yet reached its limiting rate of 1% per dpa Foster and coworkers have published three papers over the past several decades where it appeared that irradiation creep indeed occurred at a higher rate in thermal reactors than in fast reactors.182–184 In the last of these papers it was noted that, as proposed by Garner 34 the previously unsuspected 59Ni contributions to dpa might account for the apparent but possibly misleading increase in creep rate The T/F ratio in the experimental test reactors cited by Foster was rather high compared to that in PWRs An additional reason for such enhancement of creep probably lies in the large amounts of transmuted helium and stored hydrogen in thermalized spectra that results from the 59Ni sequence and the stored hydrogen concept, producing bubbles and voids that accelerate the creep rate Therefore, it does not appear necessary to invoke an enhanced survivability or displacement effectiveness of gamma recoil events to explain the apparently higher creep rates in thermal reactors 4.02.9.5.3 Dependence of creep modulus on hydrostatic stress Although it is well known that it is the deviatoric component of any stress state that drives creep, there were previously very little data to show whether the creep coefficient is identical in both dilational and compressive stress states Recent papers by Hall,185,186 Neustroev,187 and Garzarolli188 show that creep coefficients are unchanged by the sign of the hydrostatic stress As shown in the next section, additional confirmation of the independence of creep compliance on the sign of the hydrostatic stress component can be found in some stress relaxation experiments 86 Radiation Damage in Austenitic Steels 4.8 Normalized creep rate 4.0 1.0 Stress reduction ratio J B EN58 E FV548 Mk helices 347 S.S 240–360 ЊC H6 helices (1) Annealed M316 T < 304 ЊC (2) (3) 0.5 s0 = 216 MPa 3.0 0 0.5 1.0 2.0 1.5 dpa 1.0 0.9 Preload 23.6 N 36.5 N 1.0 0.6 Displacement rate (dpa s–1) ´ 107 Figure 81 Dependence of irradiation creep rate of springs made from various austenitic steels on dpa rate in and below the DFR core, normalized to the highest displacement rate studied Reproduced from Lewthwaite, G W.; Mosedale, D J Nucl Mater 1980, 90, 205–215 4.02.9.6 Creep Stress Relaxation by Irradiation There are situations where the applied load is initially fixed and then declines during irradiation There is usually a transient followed by an instantaneous creep rate defined by B0, but the load is constantly falling, leading to an exponentially declining load Two examples of in-reactor creep relaxation experiments are shown in Figure 82, both conducted on a high-nickel alloy Inconel X-750 Foster and coworkers have very convincingly demonstrated that creep coefficients derived from creep experiments could be used to successfully predict stress relaxation for the same steel in similar neutron spectra.163 Note that the creep coefficient derived for X-750 from the EBR-II experiment is 1.6  10À6 (MPa dpa)À1, just slightly larger than B0 and probably Stress reduction ratio 2.0 0.7 0.5 0.3 Neutron dose (dpa) Figure 82 (top) Stress relaxation experiment conducted on X-750 in the NRU heavy-water reactor at 300  C using constant curvature bent beams Reproduced from Causey, A R., Carpenter, C K C.; MacEwen, S R J Nucl Mater 1980, 90, 216–223; (bottom) stress relaxation of compressed springs in EBR-II at 375–415  C Reproduced from Walters, L C.; Reuther, W E J Nucl Mater 1977, 68, 324–333 enhanced by low levels of voids or bubbles in this high-nickel alloy In NRU, however, the creep relaxation proceeded much faster, partially due to a larger transient, but also because the steady-state creep rate is larger In this experiment the thermal-to-fast ratio was $10, so there was significant 59Ni enhancement of dpa rate and probably also bubble formation to enhance the creep rate The greater scatter at very low residual stresses in the EBR-II experiment is mostly due to frictional variations on the compressed Radiation Damage in Austenitic Steels springs and grain-to-grain interactions that come into play at low stress levels Stress relaxation experiments can be conducted using a wide variety of specimen types and usually yield similar results, although the transient regimes often vary with specimen geometry, preparation, and texture versus stress field relationship, as shown in Figure 83 304 0.8 Stress ratio (a/d) C ring (A2 = 1.2 ´ 10-6 MPa dpa–1) 0.6 0.4 Bend (A2 = 1.8 ´ 10-8 MPa dpa–1) 0.2 Irradiated temperature: - 561 K 0 Dose (dpa) Figure 83 Stress relaxation experiments conducted on 304 stainless steel at 288  C in water-cooled JMTR at 0.82–1.7  10À7 dpa sÀ1, showing creep coefficients close to B0, and also demonstrating different transient behavior in different test geometries Reproduced from Ishiyama, Y.; Nakata, K.; Obata, M.; et al In Proceedings of 11th International Conference on Environmental Degradation of Materials in Nuclear Systems; 2003; pp 920–929 87 Creep relaxation by irradiation is important in that it can reduce the opportunity for irradiationassisted stress corrosion cracking It does so by decreasing internal or surface stresses produced by deliberate or inadvertent damage, as well as by reducing internal stresses arising from welding, abrupt cooling, etc Figure 84 demonstrates the radiation-induced relaxation that occurs in a weld that proceeds with a creep compliance of B0 that is independent of the sign of the hydrostatic stress.189 Therefore, it appears that the creep compliance B0 can be confidently applied to any stress state As a rule of thumb one can anticipate that by 10 dpa, >90% of any preload will be relaxed even in the absence of a transient The fractional unloading is not dependent on the magnitude of the preload as long as the bolt or component was not loaded beyond the yield point Stress relaxation in structural components of operating reactors is not always operating in isolation Frequently, a component experiences time-dependent stresses that develop with time as a result of the growth or movement of adjacent components In pressurized water reactors there are bolts that join baffle plates to former plates These bolts are usually cold-worked 316 but the plates they join are annealed 304 stainless, a higher swelling steel Initially, the bolt will start to relax its preload but if the plates are swelling faster than the bolts, then differential swelling will begin to reload the bolt Additionally, if a bolt is replaced with a fresh bolt, the reloading can be even stronger due to larger amount of 250 Before irradiation 200 After 3–6 dpa irradiation sy 3W-H 0.8 Stress relaxation s/s0 150 sy (MPa) 100 50 -50 Irradiation temperature: 561 K 0.6 0.4 Tensile (6 mm from surface) 0.2 -100 -150 Compressive (4 mm from surface) Distance from surface, mm 10 20 30 40 Distance from left edge (mm) 50 60 0 Dose (dpa) Figure 84 Residual stresses in SA 304 associated with a one-pass weld with mechanical constraint Stress reversals occur with depth from the surface Reproduced from Obata, M.; Ishiyama, Y.; Nakata, K.; Sakamoto, H.; Anzai, H.; Asano, K J ASTM Int 2006, 3, 15–31 Residual stresses before and after irradiation were measured by neutron diffraction Note that B0 was determined to be $1  10À6 (MPa dpa)À1 and is independent of the sign of the hydrostatic stress 88 Radiation Damage in Austenitic Steels 350 200 10 years 40 years 300 180 140 200 Torque (nm) Axial stress (MPa) A B C 160 250 150 100 50 120 100 80 60 0 10 20 30 40 50 60 70 Time (year) Figure 85 Calculated bolt relaxation and reloading is shown for two conditions of bolt replacement in a 304 stainless baffle-former assembly Reproduced from Simonen, E P.; Garner, F A.; Klymyshyn, N A.; Toloczko, M B In Proceedings of 12th International Conference on Environmental Degradation of Materials in Nuclear Power Systems – Water Reactors; 2005; pp 449–456 The cold-worked 316 bolt is replaced and reloaded at either 10 or 40 years Note that differential swelling does not reverse the loading until almost 10 dpa as the bolt approaches full relaxation differential swelling Figure 85 shows several calculated histories of bolt loading for PWR-relevant temperatures and dpa rates.190 While bolts are generally preloaded to a specified level, there is always some range of attained loads It is difficult to measure the stress level in a bolt while it is still in place, but a rough measure of the remaining load can be made from the torque required to remove the bolt While this is not an exact measurement with friction, corrosion, irradiation-induced self-welding, and other complications possibly participating to define the torque, Figure 86 shows that the measured torques are in reasonable agreement with predictions of creep equations based on experiments conducted in BOR-60 fast reactor The fact that most of the data lie above the predictions may indicate that many of the bolts are indeed being reloaded by differential swelling to some degree 4.02.9.7 Stress Rupture While irradiation creep is relatively well understood the effect of radiation on thermal creep and thereby 10 20 30 40 Dose (dpa) Figure 86 Torques measured during removal of bolts from French PWRs of the CPO series Only bolts showing no indication of cracking are included The results are in agreement with predicted creep relaxation when applied to upper or lower preload values, but the predictions not include any reloading A, B, and C denote measurements from three different CPO plants Reproduced from Massoud, J P.; Dubuisson, P.; Scott, P.; Ligneau, N.; Lemaire, E In Proceedings of Fontevraud; 2002; Vol 5; paper 62, 417 creep rupture is not as well defined In general it appears that creep rupture properties are not improved by irradiation and are adversely affected as shown in the example of Figure 87.191,192 As shown in Figure 88 Ukai and coworkers have compared the reduction in rupture life in air, sodium, and after irradiation in FFTF, demonstrating that the largest influence is due to irradiation.193 There is some evidence that irradiation in neutron spectra that produce high He/dpa ratios will decrease rupture life, especially at higher temperatures, compared to irradiation in fast reactors due to the accumulation of helium bubbles on grain boundaries and triple points.191,192 It is possible to improve the in-reactor stress rupture properties of a given steel by additions of selected trace elements such as P and B, both of which are known to affect the distribution and stability of carbide phases An example is shown in Figure 89.194 Fortuitously, such additions also add to the swelling resistance of such steels Radiation Damage in Austenitic Steels 4.02.9.8 Fatigue Fatigue loading can be very detrimental for situations involving cyclic loading, especially when associated with thermal cycling such as might occur in the first wall of a fusion device As shown in preceding sections, radiation changes the microstructure and affects the phase stability of steels as well as generating deleterious gases such as helium and hydrogen 103 700 ЊC Stress (MPa) Annealed Thermal aged Cold-worked 102 Irradiated in BR-2 101 1.3 1.4 1.5 1.6 1.7 1.8 T(13.5 + log tR) (K) Figure 87 Effect of starting condition and irradiation in the BR-2 reactor on stress rupture behavior of DIN 1.4970 at 700  C Reproduced from Wassilew, C.; Ehrlich, K.; Bergmann, H J In Influence of Radiation on Material Properties: 13th International Symposium; ASTM STP 956; 1987; pp 30–53; Grossbeck, M L.; Ehrlich, K.; Wassilew, C J Nucl Mater 1990, 174, 264–281 Data are plotted versus the Larson Miller Parameter (LMP) The effect of radiation is stronger than the effect of cold-working Therefore it is not unexpected that fatigue life will be adversely affected by irradiation as shown in Figure 90.192 Fatigue tests are by necessity conducted out-ofreactor and therefore are not fully representative of in-reactor conditions, especially not being subject to the mitigating influence of radiation creep to reduce local stress concentrations In this sense out-ofreactor results may be conservative The tests can be conducted in a variety of ways, however, generally using either strain-controlled or load-controlled methods, with the former being more relevant to low cycle fatigue arising from thermal cycling Guidance on the application of fatigue data is provided by Tavassoli.195 Figure 90 presents the usual engineering curves of total strain versus the number of cycles to failure In this representation the lifetimes of irradiated and unirradiated materials are not really so dissimilar The observed difference is the result of competing influences, degradation due to irradiation, and improvement due to hardening As pointed out by Boutard,196 it is better to isolate the irradiation effect on the lifetime in which the controlling parameter is the plastic strain range As shown in Figure 91, there is a significant effect of radiation on the lifetime at a given plastic strain.196,197 The lower the plastic strain, the greater the decrease in lifetime Under conditions where the crack initiation phase controls the lifetime of the unirradiated material, irradiation will result in much earlier crack formation In-air In-sodium In-reactor 500 Hoop stress (MPa) 89 300 Irradiation effect Sodium effect 100 80 60 14 14.5 15 15.5 16 16.5 17 17.5 18 LMP = T (14.04 + log tR) / 1000 Figure 88 Creep rupture behavior of 20% cold-worked modified 316 stainless steel, showing effect of sodium and irradiation to reduce failure lifetimes Reproduced from Ukai, S.; Mizuta, S.; Kaito, T.; Okada, H J Nucl Mater 2000, 278, 320–327 90 Radiation Damage in Austenitic Steels and much earlier failure Other researchers have reached the same conclusion.198 In general it appears that most researchers agree that helium is a contributing but not primary cause of the radiation-induced degradation in lifetime.195–199 4.02.10 Conclusions In general there are no beneficial aspects of radiation on austenitic steels when exposed to neutron irradiation Structural components used in various nuclear reactors may have been constructed from alloys with carefully tailored and optimized properties, but there is an inevitable degradation of almost all engineering properties of interest as irradiation proceeds Even more importantly, having labored to build a device with well-defined dimensions, separations, and tolerances, it must be recognized that these dimensional attributes can also change dramatically, requiring that the design anticipate such changes in order to maximize safe and efficient operation for the longest possible lifetime Hoop stress (MPa) 1000 100 D9I D9 575 ЊC 605 670 750 10 12 D9 D9I 575 ЊC 630 695 775 14 16 18 LMP, T (13.5 + log tR) ´ 10-3 (K, h) 20 Figure 89 Improvement of in-reactor [FFTF fast reactor] stress rupture properties of D9 stainless steel by controlled additions of B and P Reproduced from Hamilton, M L.; Johnson, G D.; Puigh, R J.; et al In Proceedings ASTM Symposium on Residual Elements in Steel; ASTM STP 1042; 1989, pp 124–149 10.0 Total strain range, D ' (%) - Unirradiated - f1 = 0.7-2 ´ 1026 n m–2 - Unirradiated, aged 115 days at 430 ЊC 1.0 0.1 102 103 104 105 106 Cycles to failure 107 108 Figure 90 Fatigue life of 20% cold-worked AISI 316 stainless steel irradiated in HFIR to a maximum dose of 15 dpa and 900 appm He Reproduced from Grossbeck, M L.; Ehrlich, K.; Wassilew, C J Nucl Mater 1990, 174, 264–281 Radiation Damage in Austenitic Steels Plastic strain range (%) 100 EC - 316L : 430 °C Nonirrad Irrad.: 10 dpa 5 10-1 10 11 10-2 103 106 12 Figure 91 Plastic strain versus number of cycles to failure of annealed EC-316L irradiated to 10 dpa at $430  C in BR2 Reproduced from Grossbeck, M L.; Ehrlich, K.; Wassilew, C J Nucl Mater 1990, 174, 264–281; Vandermueulen, W.; Hendrix, W.; Massault, V.; Van de Velde, J J Nucl Mater 1988, 155–157, 953–956 Using total strain rather than plastic strain, the reduction of life was only a factor of $2, relatively independent of strain range 13 104 105 Number of cycles to failure 14 15 16 17 18 This evolution of properties and dimensions frequently determines the lifetime of any given structural component, a lifetime that will be very specific to each nuclear environment It is important to recognize that all potential degradation processes may not yet have been identified and that others may lie over the current exposure 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Austenitic Steels 2.8 ´ 1022 n