Comprehensive nuclear materials 3 16 ceramic fuel–cladding interaction

Comprehensive nuclear materials 3 16 ceramic fuel–cladding interaction Comprehensive nuclear materials 3 16 ceramic fuel–cladding interaction Comprehensive nuclear materials 3 16 ceramic fuel–cladding interaction Comprehensive nuclear materials 3 16 ceramic fuel–cladding interaction Comprehensive nuclear materials 3 16 ceramic fuel–cladding interaction Comprehensive nuclear materials 3 16 ceramic fuel–cladding interaction

3.16 Ceramic Fuel–Cladding Interaction K Maeda Japan Atomic Energy Agency, O-arai, Ibaraki, Japan ß 2012 Elsevier Ltd All rights reserved 3.16.1 Introduction and Overview of Ceramic Fuel–Cladding Interaction 444 3.16.2 3.16.2.1 3.16.2.2 3.16.3 3.16.3.1 3.16.3.1.1 3.16.3.1.2 3.16.3.2 3.16.4 3.16.4.1 3.16.4.1.1 3.16.4.1.2 3.16.4.1.3 3.16.4.1.4 3.16.4.1.5 3.16.4.2 3.16.5 3.16.5.1 3.16.5.2 3.16.5.2.1 3.16.5.2.2 3.16.5.2.3 3.16.5.3 3.16.5.3.1 3.16.5.3.2 3.16.5.3.3 3.16.6 3.16.7 3.16.7.1 3.16.7.1.1 3.16.7.1.2 3.16.7.1.3 3.16.7.1.4 3.16.7.2 3.16.7.2.1 3.16.7.2.2 3.16.7.2.3 3.16.8 References Cladding Compatibility with Oxide Fuels and FPs Formation of Protective Oxides on Cladding Materials Chemical Interaction Among Oxide Fuels, FPs, and Cladding Morphology of Cladding Attack in Oxide Fuel Pins Observations of Cladding Attack Deep localized cladding attack FCCI at the top of the fuel column Types and Characteristics of Cladding Attack Occurrence of Interaction Between Oxide Fuels and Cladding Key Parameters in FCCI Development Fuel parameters Effect of temperature Effect of burnup Effect of temperature difference between fuel and cladding Effects of cladding materials FCCI Model and Wastage Equation Mechanism of Oxide Fuel and Cladding Interaction Oxygen Potential of Irradiated Fuel Characteristics of Major Corrosive FPs Iodine Cesium Tellurium Various Corrosion Reaction Mechanisms Corrosion early in life Iodine transport of cladding constituents Cladding corrosion by Cs–Te mixture Inhibition Methods for Oxide Fuel and Cladding Interaction Nonoxide Ceramic Fuels and Cladding Interaction FCCI of Carbide Fuel Chemical reactions with FPs Formation of intermetallic compounds Clad carburization Key parameters of clad carburization FCCI of Nitride Fuels Chemical reactions with FPs Formation of intermetallic compounds Clad nitriding Outlook 445 445 446 447 447 447 447 448 449 449 449 450 451 452 455 455 457 457 458 459 459 460 461 461 462 463 466 467 467 467 467 468 470 475 475 476 476 477 478 Abbreviations AISI ANL American Iron and Steel Institute Argonne National Laboratory CCCT C/M DFR Cladding component chemical transport Carbon-to-metal Dounreay fast reactor 443 444 Ceramic Fuel–Cladding Interaction EBR-II EPMA FAE FCCI FCMI FFTF FP FPLME FR GE HEDL JOG LWR MOX MX NMA N/M O/M PFR PIE PNC RIFF SIMS SNR Experimental breeder reactor-II Electron probe microanalysis Fuel adjacency effect Fuel and cladding chemical interaction Fuel–cladding mechanical interaction Fast flux test facility Fission product Fission product-induced liquid-metal embrittlement Fast reactor General Electric Company Hanford Engineering Development Laboratory Joint oxide-gaine (French) Light water reactor Mixed oxide Nonoxide, where M stands for U ỵ Pu and X stands for C or N Nuclear microprobe analysis Nitrogen-to-metal Oxygen-to-metal Prototype fast reactor Postirradiation examination Power Reactor and Nuclear Fuel Development Corporation, currently Japan Atomic Energy Agency Re´action a` l’Interface Fissile Fertile (French) Secondary ion mass spectrometer Schneller Natriumgekuăhlter Reactor (German) 3.16.1 Introduction and Overview of Ceramic Fuel–Cladding Interaction Ceramic fuels used for fast reactors (FRs) are oxide fuels and nonoxide ceramic (MX-type, where M stands for U ỵ Pu and X stands for C and N) fuels such as carbide and nitride fuels Ceramic fuel–cladding interactions (FCCI, fuel and cladding chemical interaction) are mainly divided according to oxide and MX-type fuels FCCI is more complicated in oxide fuels than in MX-type fuels because oxide fuels lead to oxidation of cladding materials and formation of various oxides of fission products (FPs) by irradiation Chemical interactions between uranium and plutonium mixed oxide (MOX) fuels and/or FPs and cladding materials are considered as one of the major factors limiting the lifetime of fuel pins in FRs This limitation is especially important for long-term irradiation of (U,Pu)O2 fuel pins clad in stainless steel Fuel pins for FRs are generally designed with Type 316 stainless steel cladding to operate with a peak cladding hot-spot temperature of 700  C Results and analyses of irradiation experiments related to FCCI with oxide fuels have been reported in various technical society conferences and topical and periodic reports since FCCI was first reported in the late 1960s FCCI is nowadays recognized as one of the major factors determining integrity and lifetime of oxide fuel pins as demonstrated in numerous in-pile and out-of-pile tests Some mechanisms of cladding attack have been proposed from the results of the many postirradiation investigations and thermodynamic analyses of the postulated chemical reactions Cladding and oxide fuels not violently react, even under a high oxygen potential condition; they only form a protective layer on the inner wall of the cladding But when fuel pins are irradiated in a reactor, the additional effect of the generated FPs induces cladding attack A number of experiments have shown that both stoichiometric and hypostoichiometric oxide fuels react with stainless steel cladding when irradiated in typical FRs On the other hand, out-of-pile tests between (U,Pu)O2 or (U,Pu) O2Àx and several stainless steels have shown that no detectable reaction took place within the times and temperatures of interest for FRs When hyperstoichiometric fuel, (U,Pu)O2 ỵ x , was tested, cladding attack was detected and the difference in reaction behavior was ascribed to the excess oxygen provided in the hyperstoichiometric fuel Inability to reasonably extrapolate the out-of-pile results to the in-pile results is of concern for the design of oxide fuel pins, and it indicates that the prediction of lifetime is complicated The thermochemistry of the fuel–cladding gap is complex as well and difficult to predict because it depends not only on concentrations of corrosive FPs, but also on major parameters such as fuel–cladding gap width, fuel oxygen-to-metal (O/M) ratio, cladding temperature, fuel temperature, and radial temperature gradient FCCI is further regarded as sensitive to linear heat rating and likely to change with fuel burnup When the swelling of the cladding is high and the fuel–cladding temperature gap is large, the probability of attack is enhanced Thus, cladding attack tends to be unpredictable, and it may be locally worse compared to the overall condition The possible consequence is complete penetration of the cladding by a chemical mechanism alone In addition to this, it may be considered that there is some form of stress corrosion cracking in the cladding Actual creep strain on the cladding is from fuel and cladding mechanical Ceramic Fuel–Cladding Interaction interaction (FCMI) and/or internal FP gas pressure FCCI has been identified as a contributing factor in the breaches of oxide fuel pins Observations at an axial location of a breach that was located at the approximate original top of the fuel column have shown extensive FCCI The breach was a consequence of FCMI and internal FP gas pressure As explanations of the observed effects of FCCI have been speculative, fuel pin design could rest only on empirical equations rather than on fundamental models Cladding wastage equations by FCCI have been developed for fuel pin designs Most observations of FCCI showed it to be the result of simple oxidation of the inner surface of the cladding Three principal types of cladding attack in stainless steel can be distinguished The first is a general oxidation of the inner surface of the cladding The second is intergranular attack and is the most important The third is advanced attack which appears to be a transport of the cladding constituents into the fuel It is typically seen as wastage of the cladding thickness in some local areas by mechanical or liquid-phase transport of cladding constituents into the outermost oxide layer on the fuel pellets FCCI with oxide fuels has been recognized as an important factor in the ability to achieve peak burnups in the range of 10 at.% in FRs while maintaining high coolant bulk outlet temperatures However, in addition to cladding thickness losses due to FCCI, oxide fuels and FPs have the potential for reducing cladding load-bearing capabilities by mechanisms such as liquid-metal embrittlement (FPLME, FP-induced liquid-metal embrittlement) The other type of FCCI occurs for nonoxide ceramic (MX-type) fuels such as carbide and nitride fuels and the cladding MX-type fuels are chemically unreactive to sodium coolant, so sodium may also possibly be used as a medium for bonding between fuel and cladding instead of helium gas MX-type fuels are generally irradiated at lower temperatures and lower radial temperature gradients than oxide fuels, although at high linear heat rating, which results in low FPs release rate The volatile FPs (Br, I, Cs, and Rb) not form carbides or nitrides In particular, MX-type fuel pins are kept with low oxygen potential at the inner cladding surface; therefore, severe oxidative FCCI of the FPs is not expected A number of irradiation experiments have been performed with MX-type fuels to study FCCI The compatibility with cladding materials has been investigated in out-of-pile examinations and thermodynamic analyses As a consequence, unlike the case of oxide fuels, FPs from MX-type fuels not play a major role in FCCI Instead, the carburizing and 445 nitriding of cladding, and also the formation of intermetallic compounds of fuel and cladding, have been investigated as a major FCCI of MX-type fuels The carbide fuels (U,Pu)C, which are designed to be slightly hyperstoichiometric, will therefore be in the two-phase region (U,Pu)C ỵ (U,Pu)2C3 The presence of higher carbide phases carburizes the claddings Hyperstoichiometric carbides can embrittle the cladding by forming grain boundary carbides which can lead to intragranular failure of the steel after a moderate burnup The creep and swelling properties of stainless steels are sensitive to carburization and precipitation of M23C6 As sodium can act as a transfer agent, carbon transport rates through the gap in sodium-bonded fuel greatly increase Hypostoichiometric mixed carbides contain (U,Pu) metal as a second phase which may form low melting-point eutectics with iron or nickel base cladding alloys Hyperstoichiometric MN1 ỵ x-containing sesquinitride phase can cause nitrogen penetration and form a reaction layer at the cladding inner surface, which results in the clad nitriding The nitriding of cladding generally decreases the ductility and increases the mechanical strength Hypostoichiometric MN1 À x contains free metal leading to a eutectic melting reaction between the free (U,Pu) metal and the cladding, which results in formation of (U,Pu)Fe2and (U,Pu)Ni5-type intermetallic compounds At present, it is clear that the knowledge base for MX-type fuels is much smaller and less detailed than that for oxide fuels, and addition to the base is a work in progress However, MX-type fuels merit much less concern regarding cladding–fuel compatibility than oxide fuels In Sections 3.16.2–3.16.6, the causes of FCCI with oxide fuels are reviewed, considering the dependence on irradiation conditions and fuel parameters as well as types of cladding material Furthermore, the role of corrosive FPs in the FCCI, the mechanism of FCCI, and the FCCI enhancement by oxygen potential are summarized The MX-type fuel–cladding interaction is briefly described in Section 3.16.7 3.16.2 Cladding Compatibility with Oxide Fuels and FPs 3.16.2.1 Formation of Protective Oxides on Cladding Materials Out-of-pile tests between (U,Pu)O2 or (U,Pu)O2 À x and several stainless steels have shown that no detectable reaction took place for exposure times 446 Ceramic Fuel–Cladding Interaction and temperatures of a typical FR However, in hyperstoichiometric fuel, (U,Pu)O2 ỵ x , cladding attack was detected Excess oxygen was provided by the hyperstoichiometric fuel, which was considered to cause the difference in the reaction behavior Therefore, if the fuel surface O/M ratio can be maintained just below exact stoichiometry, oxidation of the cladding cannot take place Of the three major constituents of austenitic stainless steel cladding, Fe, Cr, and Ni, chromium has the greatest affinity for oxygen and forms the most stable oxide Initially, chromium begins to get oxidized when the oxygen partial pressure satisfies the equilibrium condition of the reaction 4=3Crcladdingị ỵ O2 gị ẳ 2=3Cr2 O3 sị; where the oxygen potential DG O2 ẳ RT lnpO2 Þ of the fuel surface reaches À554 kJ molÀ1 at 727  C.1 However, fuel and cladding not severely react, even when the oxygen potential is high; they only form a protective layer on the inner wall of the cladding The stable protective Cr2O3 thin layer prevents the fuel and cladding reaction from becoming thermochemically equilibrated In the initial stage of irradiation, the oxygen potential of the fuel surface rises because of oxygen redistribution Excess oxygen, after uranium and plutonium fission in the fuel, leads to an increase in fuel O/M ratio with burnup Radial redistribution of oxygen along the fuel radial temperature gradient enhances the increase of O/M ratio at the fuel surface It appears unlikely that the oxygen potential at the entire fuel–cladding interface can be kept low enough to prevent cladding oxidation throughout the entire lifetime of the fuel element Thus, a thin protective layer of oxide, mainly Cr2O3, soon forms on the inside surface of the cladding, thereby physically separating the substrate metal from the oxidizing medium Further growth of this layer requires that chromium ions diffuse from the substrate metal to the outer surface of the coating or that oxygen ions migrate in the opposite direction The rates of both these processes are very slow at 727  C because of the low values of the diffusion coefficients of the ions in the oxide layer If the thermochemically stable uniform layer is breached by mechanical forces or is dissolved by a component of the oxidizing environment, the substrate metal is exposed to rapid attack The integrity of the cladding relies on the kinetics of the chemical attack in an environment where oxidation is thermodynamically possible In addition, the inner wall temperature of the cladding in an FR-MOX fuel pin reaches the range at which the sensitization of stainless steel cladding occurs, !500  C.2 This suggests that the corrosion resistance of the stainless steel cladding might become degraded because of chromium being held in carbide particles in the cladding 3.16.2.2 Chemical Interaction Among Oxide Fuels, FPs, and Cladding As long as the protective layer stays intact, the stainless steel cladding is protected from further corrosion However, a number of irradiation experiments have shown that both stoichiometric and hypostoichiometric fuels reacted with stainless steel cladding Unlike irradiated fuel, fresh fuel does not corrode stainless steel cladding to the same extent It was suggested that irradiation damage might reduce the effectiveness of this protective layer But it was found that the extent of oxidation did not sufficiently increase while irradiation damage by fission fragments increased.3 The thermodynamic tendency of oxide fuels is to oxidize the cladding, and not to violently attack it in the absence of FPs because of the protection provided by the oxide film formed on the surface of the steel.4 However, the protective layer is impaired by a chemical reaction of reactive FPs and oxygen with chromic oxide Such evidence suggests that one or more of the FPs are responsible for accelerating the chemical reactions between fuel and cladding in irradiated fuel pins.3 FCCI is the FP-accelerated oxidative attack of the cladding that is frequently observed in FR fuel pins involving reactive FPs such as Cs, Te, and I.4 Specifically, cesium and tellurium are thought to contribute to the most aggressive intergranular attack modes.4,5 The FCCI phenomenon is generally recognized to be the result of the oxidation of chromium in the stainless steel cladding under the influence of FPs; cladding attack by Cs2Te has not been considered as an oxidation mechanism of the cladding materials In irradiation experiments, however, the protective oxide layer is breached in some places and cladding attack takes place, usually in a few isolated patches rather than uniformly Whether a chemical reaction between components of the irradiated fuel and constituents of the cladding can occur at all is determined by the thermodynamics of the reactions involved Local breakdown of the protective layer and subsequent corrosion appear to depend on the local accumulation of observed major FPs, such as Cs and Te or I, which are considered important corrosive elements The generated volatile FPs are released and accumulate at the fuel–cladding gap with increasing Ceramic Fuel–Cladding Interaction burnup When fuel surface oxygen potential exceeds the threshold necessary for oxygen transport to the cladding inner surface, excess oxygen and corrosive FPs can interact with the cladding inner surface leading to FCCI Internal wastage of the stainless steel cladding is related to the complex phenomenon of corrosion established by the presence of FPs (Cs, I, and Te) and oxygen at the fuel–cladding interface The threshold temperature for cladding attack is around 500  C 3.16.3 Morphology of Cladding Attack in Oxide Fuel Pins 3.16.3.1 Observations of Cladding Attack 447 considered as intergranular corrosion accelerated by sensitization which is seen as the loss of Cr by Cr23C6 precipitation in the grain boundary.9 As this was the most aggressive form of FCCI observed along grain boundaries deep in the cladding in the case of initial O/M ratios above 1.98, this type of attack has been largely eliminated by using fuel with O/M ratios of 1.98 and below By utilizing an appropriate lower O/M fuel associated with longer irradiation for excess oxygen in the fuel pin, a more uniform matrix interaction tends to take place A combined interaction form, consisting of matrix FCCI proceeded by intergranular FCCI, occurs in fuel with moderate O/M ratio and high burnup.2 3.16.3.1.1 Deep localized cladding attack 3.16.3.1.2 FCCI at the top of the fuel column Regions of chemical reactions between the fuel and cladding have been generally observed, especially the hotter cladding temperature regions Examination of metallography samples showed occurrence of nonuniform and deep localized cladding attack in irradiated fuel pins.6,7 Cladding attack usually occurred in an irregular manner over the inner surface of the cladding and in the case of intergranular attack, its depth of penetration varied from site to site In addition, the observed deep localized interaction was usually of a different type than in the rest of the sample When access to the substrate metal was established, cladding attack by FPs occurred, either uniformly or only locally, but in some cases it penetrated more than 100 mm into the cladding This would be a significant reduction of the effective thickness of the cladding Despite the reduction in cladding thickness, actual fuel pin failure has rarely been observed The occurrence of a deep localized interaction of more than 100 mm in depth was observed in a sample which had an initial O/M ratio larger than 1.98, and was irradiated to less than at.% burnup at cladding temperature higher than 650  C.6 That suggests that this type of interaction occurs primarily in hightemperature regions with relatively low burnup This interaction is called deep localized FCCI and is an intergranular type of cladding attack, characterized by a highly localized reaction product Because cladding attack tends to be random, it becomes locally worse compared to the overall condition For a fuel with an initial O/M ratio of 1.99–2.00, there is evidence that intergranular attack of sensitized stainless steel cladding occurs in the matrix around the carbide particles in the grain boundaries.8 Microprobe examinations have shown this area to be depleted in chromium and manganese, with significant quantities of the FP cesium present in the reaction product This was The top of the fuel column at a cladding temperature near the maximum corresponds to the boundary of fissile–fertile fuel pellets At this location, axially migrated and accumulated volatile FPs react with the cladding material Axial isotopic gamma scans for high burnup pins have shown that there are larger amounts of cesium in the area of the upper insulator pellets than in the area of the fuel.10 Because of the migration of cesium to the cold region in the irradiated fuel pins, cesium peaks are generally found at both ends of the fuel column These accumulations were generally related to the formation of a phase consisting of U–Cs–O (Cs2UO4) at the UO2 blanket or insulator pellets, which caused localized inelastic deformations of the cladding (up to 30 mm) at the fuel–blanket interfaces by a volumetric change.10,11 But Kleykamp12 confirmed that instead of Cs2UO4, a cesium uranoplutonate Cs2(U,Pu)4O12 layer was formed on the grain boundaries of the UO2 blanket pellets in the irradiation experiment Furthermore, formation of compounds at the UO2 blanket or insulator pellets led to a severe intergranular attack of the cladding (up to 100 mm) in this region.10,11 Figure shows ceramographs for a longitudinal section removed from a fuel pin (maximum burnup 14.5 at.%, 695  C cladding inner surface temperature, and initial O/M ratio 1.984).6 Both the depth and character of the FCCI had changed significantly at the fissile–fertile transition zone The maximum depth of cladding attack at the fertile and fissile fuel pellets was approximately 90 and 135 mm, respectively A similar localized form of cladding attack occurred at higher temperatures of >600  C at the fissile–fertile interface.13 This fissile– fertile interface reaction, termed RIFF (Re´action a` l’Interface Fissile Fertile, in French), is associated with migration of volatile FPs to the end of the fuel 448 Ceramic Fuel–Cladding Interaction column There is no evidence of RIFF in PE-16 (high Ni alloy) and EM-12 (ferittic–martensitic steel alloy) The occurrence of RIFF appears to depend on the choice of cladding materials.13 The change in character of the cladding attack at the top of the fuel column suggests a change in the mechanism of chemical interactions at locations of fissile and fertile fuel pellets The reaction of fuel and cesium suggests the presence of a high oxygen potential in the fuel.14 Large cesium pressures, which are generally expected in hypostoichiometric fuel, lead to the formation of cesium uranate in the UO2 blanket or insulator pellets The FP inventory and the radial temperature gradient in the region of the UO2 blanket or insulator pellets are significantly different from those at the region of the fissile fuel column The predominantly radial heat transfer in the upper region of the fuel column and the absence of heat-generating material in the UO2 blanket or insulator pellet suggest little or no thermal gradient across the UO2-cladding gap This effect has been taken into account in the design of other irradiation experiments by reducing the volume of the blanket pellets.15 The depth of chemical interaction between cladding and fuel outside the blanket–fuel interface has always been lower than 60 mm 3.16.3.2 Types and Characteristics of Cladding Attack UO2 (U,Pu)O2 Figure Cladding attack in the vicinity of the fissile–fertile fuel interface Reproduced from Lawrence, L A Nucl Technol 1984, 64, 139–153, with permission from ANS (a) From the metallographic examination of stainless steel–clad fuel pins irradiated to various burnup levels, it is generally possible to observe the character of the evolving cladding attack along the cladding temperature distribution in the fuel pin The cladding attack is classified into three types: (1) matrix, (2) intergranular, and (3) combined matrix and intergranular (also called ‘advanced’ or ‘evolved’).16 Figure 2(a) and 2(b) show typical photomicrographs of matrix and intergranular types of attack in a fuel pin with Type 316 stainless steel cladding (burnup: 50 MWd kgÀ1 M), respectively In addition, Figure 2(c) shows a severe combined intergranular and matrix attack observed in the fuel pin with Type 347 stainless steel cladding (burnup: 140 MWd kgÀ1 M) The first type of cladding attack is a general oxidation and is confined mainly to the shallow inner surface of the cladding The entire body of the inner wall of the cladding is converted to a reaction zone containing the oxides of Fe, Cr, and Ni In the regions of matrix attack, EPMA (electron probe microanalysis) results show a depletion of iron and nickel and (b) Matrix (c) Intergranular Combined or evolved Figure Types of representative cladding attack: (a) matrix, (b) intergranular, and (c) combined or evolved Reproduced from Perry, K J.; Melde, G F.; McCarthy, W H.; Duncan, R N., In Fast Reactor Fuel Element Technology, Proceedings of Conference, New Orleans, Luisiana, Apr 13–15, 1971; Farmakes, R., Ed.; American Nuclear Society: Hinsdale, IL, 1971; pp 411–429, with permission from ANS Ceramic Fuel–Cladding Interaction an enhancement of chromium and cesium Trace amounts of iodine and tellurium are also observed in the region of matrix attack.16 When this type of cladding attack evolves, there is a definite segregation of the cladding constituents in the reaction product layer The reaction product of the matrix attack is a mixture of metal particles and nonmetallic compounds in the fuel–cladding gap In addition to the three major constituents of the cladding, the reaction zone contains the FPs (Cs, Mo, and lesser amounts of I, Te, and Pd) The reaction zone does not appear to contain the heavy metals (U and Pu), and neither does the cladding The attack on the grains is uniform with no strong preference for attack along the grain boundaries The second type of cladding attack is penetrating the stainless steel cladding along grain boundaries and it is the most relevant for fuel pin failure Intergranular attack occurs where the steel is sensitized Such attack on the area of chromium depletion from a steel layer adjacent to the grain boundaries by precipitation of carbides at the grain boundaries is in accord with metallographic observations Opening of the grain boundaries from the cladding inner surface indicates that attack has occurred along them In addition, metallic and nonmetallic reaction products are also detected in the fuel–cladding gap This indicates that the grains have been chemically attacked, as evidenced by the roughened surface The third type of cladding attack is the combined matrix and intergranular attack that is characteristically observed in local areas, and is often accompanied by wastage of the cladding thickness caused by mechanical interaction or liquid-phase transport of cladding constituents in the outermost oxide layer adjacent to the fuel The dissolution of iron, chromium, and nickel in a medium of liquid cesium and tellurium present in the fuel–cladding gap is known as cladding component chemical transport (CCCT) It is interesting to note that the constituents of the cladding are not uniformly distributed in the reaction zone.17–19 3.16.4 Occurrence of Interaction Between Oxide Fuels and Cladding 3.16.4.1 Key Parameters in FCCI Development The thermochemistry in the fuel–cladding gap is complex and is difficult to predict because it depends not only on concentrations of corrosive FPs, but also on major parameters such as the fuel–cladding gap width, fuel O/M ratio, cladding temperature, fuel 449 temperature, and temperature gradient It is essential to develop correlations between the loss of cladding strength and the various parameter groups such as irradiation conditions and fuel specifications The qualitative characteristics of FCCI and the observed effects of various fuel and irradiation parameters on FCCI are described next 3.16.4.1.1 Fuel parameters Severity and frequency of internal cladding attack appear to be independent of both fuel form and fuel density In the case of vibrocompacted fuel, reduction in fuel density might contribute to increased FP release and to eased radial migration via pores in the fuel, which has a minor influence on the severity of the cladding attack (see Chapter 2.02, Thermodynamic and Thermophysical Properties of the Actinide Oxides) Annular pellet fuel having a theoretical intrinsic density of approximately 96%, which corresponds to a smear density of 80%, showed no clear difference in cladding attack in comparison with vibrocompacted fuel of the same smear density.20 From the results, a general increasing trend of the depth of cladding attack at cladding temperatures above 500  C at approximately and at.% burnup was indicated in both pellet and vibrocompacted fuels On the other hand, Batey and Bagley21 showed that the cladding attack in vibrocompacted fuel was significantly more severe in comparison with pellet fuels Also, vibrocompacted fuels pins that were irradiated at higher power levels (57–79 kW mÀ1) showed two to eight times the depth of cladding attack expected from fuel irradiations having the same inner surface cladding temperatures.22,23 In contrast, sphere-packed fuel pins which are loaded with spheres of mixed UO2–PuO2 and UO2 have exhibited decreased depths of cladding attack in comparison with pellet fuels with a similar initial O/M ratio and irradiation history.24–30 Thus, there is no clear explanation of the evidence for different cladding attack behavior caused by different fuel forms From the results of metallographic observations, the influence of the initial O/M ratio on the severity of cladding attack was emphasized in addition to the influence of the cladding temperature,14 and it appeared that the type of the attack was controlled by the initial O/M ratio The initial O/M ratio has a significant influence on the depth of cladding attack.31–34 Figure shows the influence of initial O/M ratio on the depth of cladding attack.31,32 The effects of initial fuel stoichiometry on the 450 Ceramic Fuel–Cladding Interaction Maximum cladding temperature (ЊC) Depth of cladding penetration (mils) 5.0 4.0 538 593 649 704 127 102 O/M 1.94 1.96 2.00 3.0 76 O/M 2.00 Threshold temperature 2.0 51 1.0 25 O/M 1.96 Depth of cladding penetration (mm) 482 O/M 1.94 1300 900 1100 1200 Maximum cladding temperature (ЊF) Depth of cladding penetration (mm) (a) 1000 Rapsodie I £ O/M 1.96 ³ O/M 1.98 150 (210) 100 ³1.98 50 £1.96 450 (b) 500 550 600 650 Cladding inner surface temperature (ЊC) 700 Figure (a) Maximum depth of cladding attack as a function of cladding inner surface temperature The data represent burnup ranging from to 13 at.% Reproduced from Weber, J W.; Jensen, E D Trans Am Nucl Soc 1971, 14, 175–176, with permission from ANS (b) O/M influence on cladding attack as a function of cladding inner surface temperature Reproduced from Goătzmann, O.; Duănner, Ph In Technical Committee Meeting on Fuel and Cladding Interaction, Proceedings of the International Working Group on Fast Reactors, IWGFR/16, Tokyo, Japan, Feb 21–25, 1977; IAEA: Austria, 1977; pp 43–48, with permission from IAEA characteristics of the FCCI have been examined and the results suggested that hypostoichiometric fuel has advantages with respect to cladding attack Furthermore, effects of initial O/M ratio on the maximum depth of cladding attack with increasing burnup have been confirmed.35 The effects of other parameters could not be clearly evidenced This suggests that variations of the other parameters are not so critical However, fuel impurities such as C, Si, Ni, or halogens might aggravate internal cladding corrosion either by independent interactions or through reinforcement of FP attack as catalysts Therefore, the changes in characteristics and the correlation of depth of FCCI have been determined as a function of the initial O/M ratio 3.16.4.1.2 Effect of temperature The changes in characteristics and the correlation of depths of FCCI have also been determined as a function of the cladding inner surface temperature The experimental data for this and the results show a remarkable increase in the depth of cladding attack Ceramic Fuel–Cladding Interaction 200 Depth of cladding penetration (mm) Experiment (210) O/M ³1.98 £1.96 Rapsodie I RAPS.-MON MFBS DFR 304 DFR 350 DFR 435 MOL 7A MOL 7B 150 451 2s (³1.98) 100 s X 50 X O/M ³1.98 X O/M £1.96 500 600 550 650 700 Cladding inner surface temperature (ЊC) 750 Figure Measured depth of cladding attack and the results of statistical analysis as a function of cladding inner surface temperature Reproduced from Goătzmann, O.; Duănner, Ph In Technical Committee Meeting on Fuel and Cladding Interaction, Proceedings of the International Working Group on Fast Reactors, IWGFR/16, Tokyo, Japan, Feb 21–25, 1977; IAEA: Austria, 1977; pp 43–48, with permission from IAEA with temperature above a threshold of about 500  C Measured depth of cladding attack in specimens of various irradiation experiments and the results of statistical analysis as a function of cladding inner surface temperature are shown in Figure 4.32 The depth of cladding attack increases with higher initial O/M ratio The threshold temperature of the cladding attack is higher with lower initial O/M ratio There is general agreement that the temperature threshold for cladding attack is an inner surface temperature of 500  C.4,17,21,32,36–52 From out-of-pile tests, the temperature threshold of cladding attack was identified as between 450 and 500  C, which is consistent with in-pile tests Random nonuniform cladding attack has been observed at all cladding temperatures down to $500  C in FR fuel pins Although high temperature appeared to promote more widespread attack, the depth of penetration showed no consistent variation with temperature in the range normally employed in FRs Certainly, the cladding attack generally increases with temperature only in the interval of 500–600  C; however, saturation and a decrease of the cladding attack were observed above 600  C.37 Figure shows the temperature dependence of depth of cladding attack and of neutron-induced swelling in stainless steel cladding used for the fuel pins of Phenix.53 The maximum swelling occurred near the relative axial position of 0.7 in the Phenix fuel pins Generally, swelling of FR-MOX fuel increases with burnup and results in fuel–cladding gap closure.18,54 At further burnup, swelling of the cladding begins to occur depending on the swelling properties of each material, such as length of the incubation period A large fuel– cladding gap forms again, and the FPs are released into the gap in accordance with the formation of the socalled JOG (joint oxide-gaine in French).18,19,54,55 As FPs in the gap gradually migrate to a colder region of the fissile column, the gap conductance should be degraded.56,57 Therefore, the maximum temperature increase occurred across the fuel–cladding gap at near a relative axial position of 0.7 in the Phenix fuel pin as shown in Figure This large temperature difference across the gap would lead to a thermodynamic driving force for cladding attack 3.16.4.1.3 Effect of burnup With an increase in fuel burnup, there is significant generation and migration of FPs such as Cs, Te, I, and Mo In addition, the oxygen potential increases with irradiation in the fuel pin and fuel periphery This determines the thermochemically stable chemical reactions which occur between FPs and cladding constituents A quasilinear relation exists between the maximum cladding depth of FCCI and burnup However, the influence of burnup on the depth of FP penetration into the cladding is not clear Table summarizes the characterization of FCCI as a Swelling DV/V (arbitrary scale) Ceramic Fuel–Cladding Interaction Average depth of attack (mm) 452 40 Phenix clad attack 20 Swelling 700 800 900 Cladding inner surface temperature (K) 1000 Figure Comparison of distribution of maximum depth of cladding attack with the profile of neutron-induced cladding swelling of fuel pin Reproduced from Fee, D C.; Johnson, C E J Nucl Mater 1981, 96, 80–104, with permission from Elsevier Table Characterization of cladding attack as a function of O/M ratio, burnup, and cladding inner surface temperature O/M Burnup Low (0–3 at.%) High (1.98–1.99) Moderate (1.96–1.97) Low (1.94–1.95)  Moderate (3–6 at.%)  >675 C intergranular 525  C matrix 600  C shallow intergranular and matrix 600  C combined with linear heating rates >80 kW mÀ1 Depth of significant carburization was up to 0.120 mm at 630  C and 0.200 mm at 670  C Clad cracking was only associated with 3.16.7.1.4 Key parameters of clad carburization The carburization depth increases approximately linearly with irradiation time or with burnup The amount of carburization is dependent on carbon activity (fuel stoichiometry and oxygen content), temperature (cladding temperature and fuel temperature gradient), and on the bonding materials (helium gas and liquid sodium) between fuels and cladding 3.16.7.1.4.1 Carbon activity Carbon activity in austenitic stainless steels depends on the temperature and heat treatment history.181 The solubility of carbon in sodium increases significantly on increasing the temperature.182 Hence, the carbon activity decreases on increasing the temperature at constant carbon concentration The carbon activity in MC pins depends on the C/M ratio, but not on the presence of FPs.183 Clad carburization with sodium- or helium-bonded fuels increases with increasing x in MC1 ỵ x Moreover, the effective carbon activity is increased by radial carbon transport in the thermal gradient and reduced strongly by the effect of oxygen impurities.184 The oxygen content of the sodium can affect the rate of carbon transfer.185,186 Carbon (wt%) Cracked 40% M2C3, cladding temp 610 ЊC, 91 kW m-1 >5% M2C3, cladding MAX temp 650 ЊC, 80 kW m-1 Intact 600  C and heavy carburization >20% penetration Grains and grain boundaries were heavily decorated with M23C6 and M7C3 but there was no large carbide-rich layer Clad cracking occurred In contrast, sphere-pac mixed carbide fuel containing 0.3%) Out-of-pile chemical compatibility experiments between plutonium-rich MC fuels and Type 316 stainless steel (20% cold-worked) cladding were carried out at 700  C for 1000 h by using a capsule containing four fuel pellets sandwiched between disks of the cladding The pellets containing up to 0.7% oxygen and 20% M2C3 caused insignificant carburization of the cladding to less than 12 mm in depth However, the pellets containing relatively high oxygen ($1%) and high M2C3 (60%) caused carburization of the cladding to depths of up to 90 mm.191 The depth of clad carburization in an identical test carried out with uranium-rich MC containing 1% M2C3 and around 0.2% oxygen was 150 mm without breaching in the Type 316 stainless steel cladding.192,193 In addition, dissolved oxygen in liquid sodium has caused some surface oxidation of the fuel and mass transfer of carbon, which lead to the decomposition of UC2 platelets and to carburization of the cladding.194 Hyperstoichiometric (U,Pu)C1 ỵ x generally includes (U,Pu)2C3 instead of (U,Pu)C2 as a second phase The U2C3 phase is more compatible with sodium than the UC2 phase, because the carbon activity of U2C3 is lower than that of UC2.195 In addition, Pu2C3 phase has carbon potential lower than that of U2C3 at all temperatures and for Type 316 stainless steel up to 705  C.196 471 Migration of carbon from the fuel to the cladding would occur if the carbon activity in the cladding is less than in the fuel.164 The carbon activity in the cladding depends strongly on its structure and chromium content; it equals approximately 10À3 for AISI 316 stainless steel and 10À2 for AISI 410 stainless steel.197,198 The carbon activity in UC is approximately 10À2.199 In an exposure experiment of UC containing U2C3 phase to sodium for 2000 h at 700  C, carbon transfer did not occur.196 This was explained by comparison of carbon activity between fuel and cladding The carbon activity in the carbide in equilibrium with U was calculated to be  10À6 at 727  C, and the carbon activity in (Cr,Fe)23C6 in equilibrium with austenitic stainless steel is  10À2 at 727  C The carbon activity in the fuel is influenced by oxygen and nitrogen impurities in the carbide An ‘equivalent’ carbon content is evaluated by the following equation by taking into account nitrogen and oxygen content: C ¼ wt%C ỵ 0:75wt%O ỵ 0:86wt%N where only the molecular weights are considered Stoichiometric fuel has an equivalent carbon content of 4.8 wt% The ‘equivalent’ carbon content suitable for FRs was evaluated as between 4.9 and 5.2 wt% of hyperstoichiometric carbide, and the oxygen content of carbide for helium-bonded fuel was limited to 0.4 wt%.156 The amount of oxygen is specified for sodium-bonded fuel in accordance with the ‘equivalent’ carbon content upper limit of wt% for phase stability reasons, which reduces the risk of unacceptable clad carburization.200 A preliminary model for the clad carburization by sodium-containing carbon has been proposed.197 It was assumed that the carbon activity would establish the equilibrium and that carbon diffusion into the cladding along the gradient of the activity would control the growth of carburized phase Chromium-rich carbide would be precipitated so as to achieve equilibrium of dissolved chromium in the matrix This model also assumed the carbide formation of three types Cr23C6, Cr7C3, and Cr3C according to chromium and carbon activities 3.16.7.1.4.2 Cladding temperature The cladding temperature plays the most important role in clad carburization The temperature gradient in the fuel causes the excess carbon to migrate to the colder zones of the fuel.201 Carburization depth in various stainless steels at medium burnup of sodium-bonded carbide fuel pins as a function of temperature is shown in Figure 16 472 Ceramic Fuel–Cladding Interaction Carburization depth generally increases depending on temperature increase Cold-worked Type 316 stainless steel shows less carburization in comparison to solution-annealed Type 316 stainless steel at 400 Type 316 stainless steel (solution annealed) 8–12 at.% Type 316 stainless steel (cold-worked) + 3–5 at.%204,205 Carburization depth (mm) 300 200 + + at.%204,205 + 10–12 at.%203 DIN 1.4970, cold worked at.% DIN 1.4988, cold worked at.% 7.5–9 at.% + + + ++ + + + + 100 + + + + + + 700 800 900 1000 Temperature (K) Figure 16 Carburization depth in various stainless steels up to high burnup of sodium-bonded mixed carbide fuel pins as a function of cladding temperature The shaded area shows cold-worked stainless steels Reproduced from Ronchi, C.; Coquerelle, M.; Blank, H.; Rouault, J Nucl Technol 1984, 67, 73–91, with permission from ANS least below $630  C.202–205 However, carburization in cold-worked stainless steels shows a steep increase above about 630  C Carbon penetration into the cladding as measured by EPMA is shown in Figure 17 Carbon penetration seems to follow a simple diffusion process with a constant surface concentration.202 Clad carburization corresponds to conditions of chemical diffusion Because a part of the carbon precipitates as (Fe,Cr)23C6 during carbon penetration, the effective diffusion coefficients are lower than the value of carbon tracer diffusion.206 The absolute carbon concentration profile of cladding for the irradiated (U, Pu)C fuel pin (42 MWd kgÀ1) has been measured using a secondary ion mass spectrometer (SIMS)207 and compared to earlier results by nuclear microprobe analysis (NMA) of protons from the reaction 12 (d,p)13C.208 The result is shown in Figure 18(a) Bart et al.207 quantified the metallographic results and confirmed the formation of metal carbide phase starting at the inner cladding surface (Figure 18(b)) Metallographic and nuclear microprobe studies of carbon concentration profiles in helium- and sodium-bonded claddings have led to the following conclusions.156 (1) Clad carburization increased with temperature in the range 400–700  C although the temperature dependence was not strong above 600  C (2) Carbon penetration of irradiated cladding increased with time but, unlike the out-of-pile case, the concentration gradient followed a nonideal diffusion law The irradiation experience at high linear heating rates !900 kW mÀ1 with oxygen-rich fuels Carbon (%) P1 = 870 W cm-1 Burnup = 7.4 at.% Tclad = 810 K %M2C3 = 3.8% t = 337 equivalent power days 0 100 200 300 400 Depth (µm) Figure 17 Typical carbon concentration profile in the Type 316 stainless steel cladding Cladding was carburized by mixed carbide containing 3.8% sesquicarbide at 550  C of cladding temperature and burnup of 7.4 at.% Reproduced from Ronchi, C.; Coquerelle, M.; Blank, H.; Rouault, J Nucl Technol 1984, 67, 73–91, with permission from ANS Ceramic Fuel–Cladding Interaction 473 0.8 0.7 0.5 AERE Harwell 0.4 Eir 0.3 0.2 0.1 Carbon concentration band as specified in M-316 steel Carbon (wt%) 0.6 0 (a) 10 Sodium side 20 30 40 50 60 70 Cladding thickness (%) 80 90 100 Fuel side Sodium side (b) Standard austenitic microstructure (0.04 w/o carbon) M23C3 MC3 Fuel side Figure 18 (a) Carbon concentration profile in the M 316 stainless steel cladding after irradiation with mixed carbide fuel (EIR: measured by SIMS, AERE Harwell measured by nuclear microprobe) Reproduced from Bart, G.; Aerne, E.T.; Burri, M.; Zwicky, H.-U Nucl Instrrum and Meth B, 1986, 17, 360–362, with permission from Elsevier (b) Appearance of carburized stainless steel cladding specimen The specimen was carburized by mixed carbide fuel (burnup 42 GWd tÀ1) Reproduced from Bart, G.; Aerne, E.T.; Burri, M.; Zwicky, H.-U Nucl Instrrum and Meth B, 1986, 17, 360–362, with permission from Elsevier (oxygen ¼ 1000 ppm) showed that the excess carbon of the M2C3 phases migrated toward the cladding, so the M2C3 content increased in the cold outer zone.176 Clad carburization was severe compared to the case with carbide fuel with small oxygen content Therefore, radial redistribution of M2C3 by migration of excess carbon toward the pellet periphery has to be considered Carburization of solution-annealed Type 316 stainless steel was measured across the cladding thickness of an irradiated sodium-bonded fuel pin with a low initial M2C3 content 2%.209 The thickness of the carburized zone was 0.120 mm at 850  C for a burnup of 27 MWd kgÀ1, which depends on cladding temperature The mean carbon concentration in the cladding was about 2000 ppm A carbon diffusion coefficient in the fuel calculated using the above results was D % 3.10À11 cm2 sÀ1, which suggested that diffusion mechanisms may be enhanced with irradiation 3.16.7.1.4.3 Bonding medium MX-type fuels are chemically unreactive to sodium coolant, so sodium may also possibly be used as a medium for bonding between fuel and cladding instead of helium gas The clad carburization can be developed in direct contact with fuel, or through the bonding mediums between the fuel and cladding gap such as helium gas or liquid sodium The extensive data available for helium- and sodium-bonded fuels210–220 consistently show clad carburization The influence of the carbon content of the fuel has been identified both in helium- and sodium-bonded pins Clad carburization occurs less in helium-bonded fuels, but more in sodium-bonded fuels where excessive carburization can occur with hyperstoichiometric (U,Pu)C1 ỵ x 221,222 The depth of carburization of Type 316 stainless steel and Incoloy 800 in sodium-bonded carbide fuel is two to three times greater than in comparable helium-bonded fuels.13,216 Behavior of carbon redistribution in helium- and sodium-bonded pins is different The rate of carbon transport toward the cladding is determined by the carbon diffusion with sodium-bonding or with gas-bonding.181 As sodium can act as a transfer agent, carbon transport rates through the gap in sodiumbonded fuel greatly increase In contrast, carbon transfer is slow in helium-bonded fuel with a gas-filled gap In helium-bonded pins, carburization seemed to become detrimental only after the establishment of fuel and cladding contact The hyperstoichiometric carbide fuel containing high oxygen content can cause carbon redistribution by carbon monoxide and increase M2C3 content in the pellet periphery which results in severe clad carburization Carburization rates increase with increasing M2C3 content of the Ceramic Fuel–Cladding Interaction 3.16.7.1.4.4 Buffer addition An Ellingham-type diagram of some cladding constituent elements and MX fuel systems is shown in Figure 19.226 The free energies of formation of (U,Pu)C, (U,Pu)2C3, and carbides of cladding constituent elements show that chromium forms the most stable carbides Therefore, carbon transport from the fuel to the cladding should occur at a rate determined by the kinetics of the mechanism of carbon transfer The driving force for a chemical reaction is the tendency to decrease the chemical potential Carburization of cladding steels is a diffusion controlled process in which the difference between the thermodynamic stability of carbides in the carbon source and carbides produced in the cladding is a driving force In an immersion test for about 336 h at 650  C, carbon was transferred from carburized Type 316L stainless steel to V–15 wt% Cr–5 wt% Ti alloy.176 The free energy of formation of titanium carbide (TiC) is about –172 kJ molÀ1 as compared with about À71 kJ molÀ1 for chromium carbide (Cr23C6) formation This indicated that carbon was transported in static sodium from a carbon-bearing material to another material in order to form a more stable carbide Thus, the thermodynamic stability of Cr23C6-type carbide relative to the other carbides U2 C + C = 2U C 2UC + C = U2C3 3Fe + C = Fe3C 2PuC + C = Pu2C3 23Cr + 6C = Cr23C6 –10 7Cr 23C6 + 27C = 23C r7C3 –20 –42 2Mo + C = Mo2C Pu+ C = PuC 23Cr + 6C = Cr 2W + C = W C –84 23 C6 U + C = UC –30 2V + C = V 2C –126 ΔGC = RT In ac (kJ g−1 atom C) carbide At low linear heat ratings the detrimental effect of impurity oxygen is not apparent The compatibility between higher carbide fuel and austenitic stainless steel cladding bonded with sodium has been studied.223 In sodium-bonded fuels, carbon transfer is effective from the beginning of the irradiation As the fuel pellets form many cracks during irradiation, liquid sodium can penetrate into the cracks of the pellets This results in shortening the diffusion path of carbon from the fuel and presenting new surfaces to the sodium Because of the limited solubility of carbon in sodium,224 carbon probably enters the liquid sodium as atoms Carbon is dissolved in the liquid sodium with an apparent heat of solution of 109 kJ molÀ1 and a saturation concentration at about 700  C of approximately 100 ppm in weight,225 and it is transferred to the cladding The maximum cladding inner surface carbon content observed in pins of low linear heating rate ($40 kW mÀ1) is the same in sodium- and helium-bonded carbide and gas-bonded oxide pins and appears to saturate at approximately 0.4 wt%.156 This level of carburization should not affect mechanical properties significantly ΔGC = RT In aC (kcal g–1 atom C) 474 2Nb + C = Nb2C –40 –168 Ti + C = TiC C 2Ta + C = Ta 400 Zr + C = ZrC 800 1200 Temperature ( ЊC) 1400 Figure 19 The carbon potential of various cladding components and UC and PuC as a function of temperature Reproduced from Matzke, Hj Science of Advanced LMFBR Fuels; Elsevier, UK, North Holland, Amsterdam, Netherlands, 1986, with permission from Elsevier in the system would determine whether the movement of carbon occurred in the sodium To have good compatibility and good irradiation behavior, it is thus desirable to stabilize the monocarbide One way is to retain the NaCl structure by alloying with a material such as TiC or ZrC, or forming carbonitrides or carbosulfides However, sodium- and vacuum-bonded compatibility tests at 600 and 700  C between stainless steels and carbosulfide fuels of differing sulfur contents and stoichiometries revealed that the amount of clad carburization is a linear function of only the total nonmetal content (C ỵ O ỵ N ỵ S) of the fuel, and does not depend on the actual sulfur content of the fuel.227 No apparent differences in carburization of Type 316 stainless steel for a range of carbonitrides with C/(C ỵ N) ratios varying between 0.85 and 0.20 were identified.228 Another way to stabilize the monocarbide is to make a ternary metallic addition to obtain a thermodynamically stable three-phase field in which UC is a stable phase From a study on carbonitride, it was concluded that it was not easier to prepare single-phase carbonitride than pure carbide, at least up to 15 mol% UN.229 The principle of the buffering action of several metals Ceramic Fuel–Cladding Interaction or carbides has been described by Chubb.166 Among other materials studied, Fe, Cr, V, Re, and W were found to be effective as buffering agents Attempts to stabilize the MC phase by buffer addition were made using a buffer acting as a carbon acceptor from M2C3 and as a carbon donor to free metal The compounds used as buffers for UC fuel were W–UWC2,166 Mo–UMoC2,230 V–UVC2,231 Cr–Cr23C6,231 Fe–UFe2,233 and Ti–TiC.230,234 Attempts for buffering (U,Pu)C by using W, Fe, Cr, Mo, and Ti were also made.230,234 Molybdenum was reported to stabilize C23C6-type carbides in austenitic stainless steels.235 Although sodium-bonded cladding materials such as AISI 316 and AISI 410 stainless steels were carburized by uranium carbide after 2–16 weeks at 800  C, both Cr and V additions to uranium carbide are effective to improve the compatibility of the carbide fuel with the stainless steel claddings The carbon activity in uranium monocarbide has been buffered with Cr or V to avoid carbon transfer from carbide fuels to the cladding.164 AISI 316 stainless steel, in contact with threephase alloys UC–Cr–Cr23C6, did not show any sign of carburization after 20 weeks at 800  C But, when insufficient Cr was present to keep the three-phase composition, formation of higher chromium carbides (Cr7C3) and carbon transfer to the steel occurred.164 Compatibility tests between vanadium stabilized UC and AISI 410 or 316 stainless steels were performed and the compatibility was better depending on increased vanadium content However, a good compatibility between uranium carbide and vanadium alloys could be observed only if the reaction rate was very slow.236 It was also reported that chromium addition to modified monocarbide fuels was effective to improve the compatibility of the carbide fuel with stainless steel claddings.236 The burnup of (U,Pu) atoms in-pile tends to shift the fuel toward the hyperstoichiometric structure Therefore, a modified fuel was proposed to buffer the carbon activity against irradiation-induced changes and minimize the formation of potentially deleterious second phases Two systems employing Cr- and Fe-based additions were selected for examination The Cr-based modification, which yields fuel containing stoichiometric (U,Pu)C with Cr and chromium carbides as grain boundary second phases, is considered the better system of the two studied This preference is strongly influenced by the minimum melting points of second phases 475 in Cr-modified fuels (1300–1350  C) compared with those in Fe-modified fuels (1000–1050  C) Fe-modified mixed carbide fuel may contain the ternary (U,Pu)Fe2–Fe-(U,Pu)C eutectic which melts at 1045  C Investigations into the constitution of these systems and systems using W and V additions have been reported.166,237 Out-of-pile compatibility tests of hypostoichiometric UC, Cr23C6-modified UC, and Fe-modified UC with stainless steels in sodium showed that fuel–cladding reactions were eliminated by the additions.238 Similarly, modified (U,Pu)C has also been shown to be stable in sodium.239 In these latter tests, modified (U,Pu)C fuels in sodium at 800  C showed no metallographic evidence of carburization of Type 304 stainless steel after approximately 1400 h of testing A number of researchers have suggested the use of buffering agents or stabilizers to minimize the formation of the potentially deleterious second phases.164,189,238,240,241 Chromium, vanadium, and iron seem to be the three elements referred to most consistently They promote the formation of a polyphase fuel structure without the compatibility problems of either free metal or higher carbide phases Van Lierde et al.242 suggested that very fine precipitates of vanadium carbide would work well as a buffering agent 3.16.7.2 FCCI of Nitride Fuels FCCI of mixed nitride fuels is similar to carbide fuels except regarding clad carburization FCCI in nitride fuel pins consists of clad nitriding, slight reactions with FPs, and formation of intermetallic compounds Clad nitriding by nitride fuels is less concerned with FCCI than clad carburization by carbide fuels, and there have been few problems reported on compatibility with nitride fuels and cladding (see Chapter 2.03, Thermodynamic and Thermophysical Properties of the Actinide Nitrides and Chapter 3.02, Nitride Fuel) 3.16.7.2.1 Chemical reactions with FPs In nitride fuel, oxidation of the fuel is excluded The information on the behavior of major volatile FPs such as Ba, Cs, and Te is limited.243–246 Chemical forms in the irradiated fuel are indicated as Cs(1), Ba3N2, BaTe, and Cs2Te by the thermodynamic calculation.247 The halogens (Br and I) and alkali metals (Rb and Cs) not form nitrides but tend to form compounds with the other volatile elements, such as CsI Only the chalcogenides (Te, Se) form nitrides 476 Ceramic Fuel–Cladding Interaction (SeNx and Te3N4), but these probably decompose even at low temperatures Therefore, in irradiated nitride fuels, Se and Te are expected to be present either in elemental form or in compounds with the other volatile elements such as Cs2Te or CsI Chemical reactions with FPs were not observed in a capsule of sodium-bonded fuel and Type 304 stainless steel cladding at high cladding temperature (795  C) and high burnup ($15 at.%).248 The cladding was sensitized after irradiation; however, electronmicroprobe examinations have revealed no trace of either fuel or FPs in the cladding 3.16.7.2.2 Formation of intermetallic compounds The binary systems U–N249–253 and Pu–N226 are less complicated than the systems of carbides The main compatibility problems of MN with cladding are also because of its narrow homogeneity range The two main uranium nitrides are the mononitride UN and the sesquinitride U2N3 Precipitations of U within grains or at grain boundaries are seen in the U-rich UN phase Hypostoichiometric MN1 À x-containing free U or Pu should be avoided as a eutectic melting reaction with the cladding can occur Hypostoichiometric MN1 À x contains free metal leading to a eutectic melting reaction between the free (U,Pu) metal and the cladding, which results in formation of (U,Pu)Fe2 and (U,Pu)Ni5-type intermetallic compounds.254 An intermetallic reaction phase (probably (U,Pu)(Ni,Fe)3) has been observed in the fuel–cladding gap.216 A low melting eutectic composed of free metal of the fuel and the constituents of the cladding can cause melting of cladding at 1000  C in the hypostoichiometric sodium-bonded fuel.255 At 1000  C, the mixed nitride fuel appears to get oxygen from a variety of cladding materials; as a result, the activity of nitrogen is increased by oxidizing the fuel surface The prediction of a stable reaction with cladding constituents considering the displacement of nitrogen from the fuel to form a nitride was made by thermodynamic calculations.255 Pure UN and PuN are stable with respect to most major cladding constituents such as Nb, Ta, V, Mn, Cr, Fe, and Mo Zirconium and titanium possibly form nitrides Aluminum is unstable with respect to UN, and its additions must be restricted Nickel is stable with respect to both UN and PuN on the basis of nitride formation except for the formation of UNi5 However, (U0.8Pu0.2)N is unstable with respect to pure nickel and Inconel-625 at low temperatures (527 and 727  C) With U2N3, only iron, nickel, and molybdenum are stable on the basis of the potential for nitride formation 3.16.7.2.3 Clad nitriding In the U–Pu–N system, the mononitrides show complete miscibility Plutonium sesquinitrides not exist in the Pu–N system However, after Lorenzelli256 reported a sesquinitride phase with a Pu/(U ỵ Pu) ratio of 0.15, Potter257 suggested the (U,Pu)N phase diagram with either very small Pu solubility or with up to 15 mol% ‘PuN1.5’ solubility in UN1.5 existed It should be noted that the most frequent composition of advanced fuels will be at about 20 mol% PuN Also, MN fuels are usually specified to be singlephased mononitrides (in contrast to MC fuels which are frequently specified to contain 10–15 wt% M2C3) As the binary systems U–N and Pu–N show very narrow single-phase fields for the mononitrides, it seems reasonable to assume that the mononitride solid solutions (U,Pu)N also exist over a relatively narrow composition range High-purity stoichiometric MN fuel is not likely to have any chemical interaction with Type 316 stainless steel cladding Single-phased MN fuel also shows excellent compatibility with stainless steel Plutonium sesquinitrides are not usually formed, but a sesquinitride phase can be formed with a Pu/(U ỵ Pu) ratio of 0.15 in the UPuN system.256 Uranium-rich MN1 ỵ x containing a second phase of mixed sesquinitride (M2N3) caused nitriding of Type 316 stainless steel cladding and consequent loss of its ductility.134 The sesquinitride in (U,Pu)N1 þ x is essentially U2N3 (with some Pu in solution) which decomposes at about 1350  C Hyperstoichiometric MN1 ỵ x-containing sesquinitride phase can cause nitrogen penetration and form a reaction layer at the cladding inner surface, which results in the clad nitriding Diffusion of nitrogen into the cladding and nitride precipitation can occur according to the following reaction.134,258 M2 N3 ỵ 2Cr ! Cr2 N ỵ 2MN As the N/M (nitrogen-to-metal) ratio of (U,Pu)N increases with burnup, the amount of sesquinitride (essentially U2N3) would increase with burnup However, excessive nitriding of cladding by nitride fuels irradiated to high burnups have not been reported Irradiation tests confirmed the excellent compatibility of nitrides with stainless steel.176,248,259–271 Oxygen impurity in MN fuel plays a vital role in controlling nitriding of Type 316 stainless steel Ceramic Fuel–Cladding Interaction cladding The nitrogen activity is decreased by oxygen impurities in the fuel.134 Some 2–3 wt% of oxide in MN stabilizes the U2N3 phase.270 Oxide stabilizes the M2N3 phase and decreases the nitrogen activity, thus reducing clad nitriding On the other hand, if oxygen is present in MN as solid solution or exists as impurity in the filling gas or in the sodium bond, or originates from the cladding, it causes surface oxidation of MN by the following reaction, leading to release of N2 and clad nitriding.134 MN ỵ 2O2 ! MO2 ỵ 1=2O2 The compatibility test of (Pu0.3U0.7)N pellets containing a large amount of oxygen ($0.5 wt%) with Type 316 stainless steel cladding showed a discontinuous clad nitriding around 35 mm.192 Similar effect of oxygen content was observed in uranium-rich mixed nitride.258 The helium-bonded mixed nitride fuel pin irradiated in EBR-II achieved a fuel burnup of 32.1 MWd kgÀ1 M at a peak linear power of 81.4 kW mÀ1.248 Excellent compatibility was exhibited between Type 304 stainless steel cladding and single-phase mixed nitride fuel Cladding in contact with fuel containing the U2N3 phase exhibited a reaction layer indicative of nitrogen penetration A typical section of an irradiated stainless steel cladding specimen is shown in Figure 20.248 The darkening of the cladding near the fuel indicated both reaction with the cladding and 477 formation of a higher nitride phase No reactions in the fuel were noted that could be associated with the cladding Mixed nitride fuel and Type 304 stainless steel cladding, sodium-bonded to one another, were completely compatible at temperatures of 700– 800  C for at least 10 000 h.255 As expected, helium bonding did not lead to significant nitriding either Hyperstoichiometric nitride fuels containing sesquinitride may possibly nitride the cladding, based on thermodynamic considerations The formation of Cr2N and the compatibility of fuel and iron were confirmed by a thermodynamic analysis of UN and cladding materials.138 Chromium nitride precipitates were observed in the cladding, but no important detrimental effect was identified The main impurities likely to be introduced during the nitride fabrication process are carbon and oxygen The influence of such impurities on the fuel properties of UN and (U0.8Pu0.2) N has been investigated by Arai et al.271 The nitrided cladding generally decreases the ductility and increases the mechanical strength Compatibility of mononitride fuel with stainless steel was confirmed by out-of-pile and in-pile tests with gasand sodium-bonded fuel pins.226 A thin layer of the cladding inner surface was nitrided in gas-bonded fuel pins but no deterioration of the mechanical properties of the cladding was shown.176,259 The microhardness measurements in the compatibility test of (Pu0.3U0.7)N pellets containing higher oxygen ($0.5 wt%) with Type 316 steel cladding showed Vickers Hardness Numbers of 265 and 210 at locations in the reaction zone and away from it, respectively.192 3.16.8 Outlook 100 µm Figure 20 Appearance of clad nitriding of Type 304 stainless steel cladding Specimen was obtained from a helium-bounded mixed nitride fuel pin irradiated in EBR-II (etched) Reproduced from Bauer, A A.; Brown, J B.; Fromm, E O.; Storhok, V W In Fast Reactor Fuel Element Technology, Proceedings of Conference, New Orleans, LA, Apr 13–15, 1971; Farmakes, R., Ed.; American Nuclear Society: Hinsdale, IL, 1971; pp 785–817, with permission from ANS While considerable progress has been made in the explanation of FCCI mechanisms, the models are not able to provide accurate predictions at this time FCCI of oxide fuels has come to be understood through irradiation experiences gathered over a long period FCCI data of MX-type fuels also have been accumulated but are still insufficient Although there are several cladding materials that can be used for fuel elements, existing chemical compatibility data indicate their capability for withstanding only moderate irradiation periods In oxide fuels, technological breakthroughs regarding inhibition of FCCI will be needed to extend the lifetime of fuel pins further However, MX-type fuels merit much less concern regarding cladding–fuel compatibility than 478 Ceramic Fuel–Cladding Interaction oxide fuels Further FCCI data of MX-type fuels irradiated at higher linear heating rate will be needed to examine the compatibility of fuel and cladding at high temperatures The restructuring of MX fuel might provide indications of the mechanisms associated with FCCI In the case of oxide fuels, the observed regimes of complex intergranular and matrix attack in oxide fuels cannot be explained by simple mechanisms Generally, the design approach with respect to FCCI includes consideration of cladding wastage at the 2s confidence level in order to reserve assurance of achieving the burnup goal without cladding failure Therefore, intergranular attack must be prevented or minimized The actual embrittlement of cladding by intergranular attack is difficult to examine in compatibility tests Experiments designed to clarify the phenomenon should be encouraged References 10 11 12 13 14 15 16 Olander, D R Fundamental Aspects of Nuclear Reactor Fuel Elements USERDA, TID-26711-P1, 1976 Lambert, J D B.; 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The MX-type fuel–cladding interaction is briefly described in Section 3. 16. 7 3. 16. 2 Cladding Compatibility with Oxide Fuels and FPs 3. 16. 2.1 Formation of Protective Oxides on Cladding Materials. .. Cr2 O3 O2 4Cs 2Mo ỵ 2Cs2 ẵU; PuO3:56 ỵ 52 ỵ x O2 2Cs2 MoO4 ỵ 2ẵU; PuO2x 525 35 1 2Cs2 ẵU; PuO3:56 ỵ 12O2 ⇄2Cs2 ½U; PuŠO4 431 424 37 6 37 2 Mo þ Cs2 Te þ 2O2 ⇄2Cs2 MoO4 þ Te Mo þ O2 ⇄MoO2 10 31 2... solution-annealed Type 31 6 stainless steel at 400 Type 31 6 stainless steel (solution annealed) 8–12 at.% Type 31 6 stainless steel (cold-worked) + 3 5 at.%204,205 Carburization depth (mm) 30 0 200 + + at.%204,205