Comprehensive nuclear materials 3 03 carbide fuel

Comprehensive nuclear materials 3 03 carbide fuel Comprehensive nuclear materials 3 03 carbide fuel Comprehensive nuclear materials 3 03 carbide fuel Comprehensive nuclear materials 3 03 carbide fuel Comprehensive nuclear materials 3 03 carbide fuel Comprehensive nuclear materials 3 03 carbide fuel

3.03 Carbide Fuel A K Sengupta, R Agarwal, and H S Kamath Bhabha Atomic Research Centre, Mumbai, India ß 2012 Elsevier Ltd All rights reserved 3.03.1 Introduction 56 3.03.1.1 3.03.1.2 3.03.2 3.03.2.1 3.03.2.2 3.03.3 3.03.3.1 3.03.3.2 3.03.3.3 3.03.3.3.1 3.03.3.3.2 3.03.3.4 3.03.3.4.1 3.03.3.4.2 3.03.4 3.03.4.1 3.03.4.1.1 3.03.4.1.2 3.03.4.1.3 3.03.4.1.4 3.03.4.1.5 3.03.4.1.6 3.03.4.2 3.03.4.3 3.03.5 3.03.6 References History of Carbide Fuel Glimpses of Carbide Fuel Physical Properties Thermophysical Properties of Carbide Fuel Thermochemistry of Carbide Fuels Fabrication of Carbide Fuel Melting Casting Hydride/Hydrocarbon Route Carbothermic Reduction Route Direct pressing method Sol–gel (wet) route Quality Control Chemical quality control method Physical quality control In-Pile Performance Introduction Burnup Swelling Performance of Na-bonded and He-bonded fuel pins Irradiation creep Experience on irradiation performance Fuel–clad chemical interaction Effects of Burnup on C/M Ratio and Chemical State of Fission Product MA-Containing Fuel Fuel Reprocessing and Waste Management Summary 56 57 59 59 60 64 64 64 64 65 66 67 68 69 70 70 70 71 72 73 75 78 79 80 82 83 84 Abbreviations ABR ADS BARC Advanced burner reactor Accelerated-driven system Bhabha Atomic Research Centre, Mumbai C/M, N/M Carbon to (U ỵ Pu) ratio, nitrogen to (U ỵ Pu) ratios CARLO Carbide low in oxygen CARRO Carbide rich in oxygen CDT Compounded doubling time CEA, France Commissariat a` l’e´nergie atomique, France EBR I Experimental breeder reactor I EBR II Experimental breeder reactor II EFR EPMA FBR FBTR fcc FCCI FCMI FFTF GEN-IV GFR IAEA ICP/MS European fast reactor Electron probe microanalysis Fast breeder reactor Fast Breeder Test Reactor, Kalpakkam, India Face-centered cubic Fuel–clad chemical interaction Fuel–clad mechanical interaction Fast flux test facility Generation IV Gas-cooled fast reactor International Atomic Agency Inductively couple plasma mass spectrometer 55 56 Carbide Fuel IDMS IGCAR ITU LDP LHR LMFBR LWR MA MKI & MKII NIMPHE PFR PGM PHWR PIE PUREX SEM SFR TD TIG TIMS TREAT TRU U/M, Pu/M UREX VPPM XGAR XRD XRF Isotopic dilution mass spectrometry Indira Gandhi Centre for Atomic Research, Kalpakkam Institute of Transuranium Elements Large development plant Linear heat rating Liquid metal fast breeder reactor Light water reactor Minor actinides Mark I and Mark II fuels of FBTR, Kalpakkam, India NItrure Mixte dans a Phenix Prototype fast reactor Platinum group metals Pressurized heavy water reactor Postirradiation examination Plutonium and uranium recovery by extraction Scanning electron microscopy Sodium-cooled fast reactor Theoretical density Tungsten inert gas Thermal ionization mass spectrometry Transient-overpower test Transuranium element Uranium to (U ỵ Pu) ratio, plutonium to (U ỵ Pu) ratio Uranium recovery by extraction Parts per million by volume X-ray gamma autoradiography X-ray diffractrometry X-ray fluorescence spectroscopy 3.03.1 Introduction 3.03.1.1 History of Carbide Fuel The world energy requirement from nuclear sources was mostly met by the use of uranium in pressurized heavy water reactors (PHWR) or light water reactors (LWR) However, uranium reserve is limited and it was felt necessary that the fuel cycle based on uranium must be a closed fuel cycle In the vast majority of LWRs or HWRs, nuclear power is generated by the fission of 235U by thermal neutrons The 235 U content in natural uranium is only 0.7% and the remaining is 238U which is nonfissionable 238 U can, however, be converted to fissionable 239Pu, which can be subsequently used as fuel for fast reactors The energy generated from natural uranium using FBRs is about 60 times greater than that obtained from LWR, even after allowing for losses in the fuel cycle.1 Hence, for effective utilization of the limited resources of U, the fast reactor concept came into existence Consequently, efforts were in progress to develop fast breeder reactors (FBRs) in several countries such as the United States, the United Kingdom, France, and the erstwhile USSR In the 1950s, the early fast reactor fuel developed was metallic (experimental breeder reactor I, EBR I) and it became the first fast reactor in the United States to generate electricity This led to the development of the second generation of larger fast reactors, which started operating in the 1970s With the inception of fast reactor technology, the power density of the nuclear reactor core went up manyfold, thus enabling compact reactors producing larger power These first few prototype reactors were BN350 (USSR), PHENIX (France), and PFR (United Kingdom) Large prototype fast reactors of capacity 500 MWe were developed in the USSR (BN600) in 1980 and another one of 1200 MWe capacity was developed in France (Super Phenix) The first-generation large breeder reactor used uranium–plutonium-mixed oxide fuels in austenitic stainless steel clad with sodium as coolant The technology for mixed oxide fuel fabrication and its irradiation performance are quite well established and a burnup of 10 at.% or more has been achieved in several reactors: for example, PFR (United Kingdom), Phenix, fast flux test facility (FFTF) (United States), and JOYO (Japan) The European fast reactor (EFR) program has also considered mixed oxide fuel for the first demonstration-type reactor Although FBRs based on mixed oxide fuel proved quite successful in enhancing energy production, this option was not a plausible futuristic solution particularly for developing countries which have limited uranium reserve to meet their long-term sustainable energy demand The anticipated shortage of fissile material in future led to the search for a fuel having higher breeding ratio and lower compounded doubling time (CDT) (less than 15 years) than oxide (about 25 years) The CDT is defined1 as the time needed to double the inventory of fissile material allowing for outof-reactor time, reprocessing, and losses during out-of-reactor time in a system of breeder reactors where the excess fissile material is used to start up new breeders as soon as the fuel from each cycle is reprocessed The high doubling time of oxide fuel is attributed to its lower heavy-atom density It also operates at a lower linear heat rating because of its low thermal conductivity These limitations of the oxide fuel led to Carbide Fuel the opening up of worldwide research for a suitable fuel with higher fissile atom density and thermal conductivity Metallic fuel based on U–Pu–Zr or U–Zr was presumed to be a possible solution, but it too had limitations such as low melting point and high swelling rates Other possible fuel compositions with higher fissile atom density were nonoxide ceramics, for example, uranium–plutonium-mixed carbide or nitride These fuels have higher thermal conductivity (about times higher than oxide; Figure 1), high fissile heavy-atom density, and a reasonably high melting point compared to metallic fuel.2 The higher thermal conductivity of carbide fuel results in more efficient heat transfer from the fuel to the coolant compared to oxide and this, in combination with a high melting point, makes carbide fuel suitable for operation at high specific power without causing any fuel central melting This also enables the use of large diameter fuel pins, more fissile material per pin, and more power generation Increase in the oxygen/metal ratio with burnup in case of oxide fuel increases the probability of fuel–clad chemical interaction (FCCI), whereas for carbide and nitride fuel, the nonmetal to metal ratio (C/M or N/M) remains constant or may even decrease depending upon the composition of the fuel High-specific-power operation permits fewer pins, compared to oxide, and a more compact core This in totality reduces the plant cost significantly These findings were very encouraging and led to initiation of a large amount of development work on carbide and nitride fuels from 1960 to 1970, and more than 5000 advanced fuel pins have been fabricated and irradiated However, nitride fuel has an inherent 30 Thermal conductivity λ (Wm-1 K-1) U–19Pu–10Zr 25 MC MN 20 15 10 500 Oxide MO1.98 problem of 14N because of the (n.p.) reaction leading to formation of 14C Hence to get the full benefit of a nitride fuel, 14N should be replaced by 15N, which is an expensive proposition The (n.p.) reaction will have a negative effect on the breeding ratio However, a fuel cycle based on carbide fuel could be the most feasible solution, and extensive research on different aspects of carbide fuel based on out-of-pile and in-pile experiments has been carried out in countries such as the United States, Germany, France, and India with the objective of developing their fast reactor program based on ‘carbide fuel.’ The development of the carbide-fueled FBR was a little sluggish during the 1970s The practical difficulty of fabrication of carbide fuel economically was probably the cause, since the high-purity inert cover gas required for fuel fabrication was expensive and maintenance of C/M ratio was difficult Apart from this, the thermo-chemical in-pile behavior of carbide fuels, e.g., extent of loss of ductility of the cladding due to carburization was not fully understood Behavior of the fuel under off-normal conditions, like loss of sodium bonding during transient-overpower conditions, was also not systematically investigated In the absence of this information, utilization of carbide fuel was limited to test pin or capsule irradiation only The Indian FBR program, however, started with the introduction of plutonium-rich mixed uranium– plutonium carbide as the driver fuel for 40 MWth (13 MWe) loop-type reactor (fast breeder test reactor, FBTR) The design of the Indian FBR was similar to that of Rapsodie (Rapid Sodium) Fortissimo version based on oxide fuel The reactor became critical in the year 1985 and it is the only reactor operative on a full core of carbide fuel Carbide fuel cannot be used in LWR because of its incompatibility with the coolant However, it can be safely used with liquid metal (Na or lead) or gas cooled (CO2 or He) in Generation IV (Gen-IV) type of high-temperature reactor Hence, carbide fuel is considered as an ‘advanced fuel’ with the basic characteristics of a breeding ratio of at least 1.30 or more and a doubling time of 15 years or less The burnup limit could be about 15 at.% or more Table shows the test irradiation program of carbide fuels in different reactors around the world 3.03.1.2 1000 1500 Temperature (ЊC) 2000 57 Glimpses of Carbide Fuel 2500 Figure Thermal conductivity of uranium–plutoniummixed carbide, nitride, oxide, and metallic fuels Many laboratories in the world were engaged in studies on the development of carbide fuels based on uranium and plutonium and thorium The results of these 58 Carbide Fuel Table History of carbide fuels used in different countries and in different reactors Fuel type Reactor Country/ organization Bond Density (% TD) Burnup (at.%) Clad References MC MC MC MC MC MC MC UC/MC MC MC MC MC RAPSODIE BOR60 EBRII RAPSODIE KNK II EBR II EBR II BOR 60 FFTF PX FFTF FBTR France/CEA USSR United States TUI FZK, Germany United States United States USSR United States CEA/TUI DOE/PSI India Na Na Na He He He He He He He He He 91.5 – – 86 85 80/87 81/87 – 80 80/82 – 91/86 12 – 12 12 16–20 10 10 – 10 16 – OX16H15M3G PE16 – 1.4970 316.20 cw 316.20 cw OX16H15M3GD9 15/15 Ti D9 SS316 cw 10 Table Properties of mixed uranium–plutonium oxide, carbide, nitride, and metallic fuels for SFR Properties (U0.8Pu0.2)O2 (U0.8Pu0.2)C (U0.8Pu0.2)N U–19Pu–10Zr Theoretical density (g cmÀ3) Melting point ( K) Thermal conductivity (W mÀ1  K) 1000 K 2000 K Crystal structure Breeding ratio Swelling Handling Compatibility Clad Coolant Dissolution and reprocessing amenability 11.04 13.58 14.32 15.73 3083 2750 3070 1400 2.6 2.4 Fluorite 1.1 Moderate In air 18.8 21.2 NaCl 1.2–1.3 High Inert atmosphere 15.8 20.1 NaCl 1.2–1.3 High (?) Inert atmosphere 40 Multiphase 1.4–1.5 High Inert atmosphere Average Average Demonstrated on industrial scale for aqueous and pilot scale for pyro-processes Large Carburization Good Dissolution not simple Process not yet demonstrated on industrial scale Good Good Dissolution easy but risk of C14 in waste management Eutectics Good Pyro-reprocessing demonstrated on pilot plant scale Limited Very little Limited Fabrication/irradiation experience Good studies were brought together in the International symposium on ‘Carbide in Nuclear Energy,’ Vols I and II held at Harwell in 196311 at a time when carbide was being seriously thought of as a reactor fuel Subsequently, following the oil crisis in 1974, a national advanced liquid-metal-cooled fast breeder reactor (LMFBR) fuels development program was initiated in the United States on the basis of the data available in the exploratory years of basic development work on carbide fuel, and a unified national approach was pursued to look into all the aspects of carbide fuels Both carbide and nitride fuels offer the best potential for LMFBR performance in the long run because of their higher thermal conductivity, fissile-atom density, and chemical compatibility with liquid sodium Mixed oxide fuel has a higher melting point than carbide or nitride fuels but the higher thermal conductivity of the carbide fuels compensates for it Some important parameters of these different types of fuel are given in Table There are two concepts available for the carbide fuel pin depending upon the type of bond between Carbide Fuel the fuel pellet and the cladding material: He-bonded and Na-bonded carbide fuels The average operating fuel temperature of the He-bonded pin is high because of low thermal conductivity of the He bond compared to the Na bond This design requires a larger fuel–clad gap and low fuel density (85%) compared to oxide fuel so that the higher swelling of the fuel due to high operating temperature can be accommodated The fission gas release will also be higher compared to that from a sodium-bonded pin In case of the sodium-bonded pin, the fuel–clad gap is larger and the fuel density higher so that there is enough radial space for accommodating fuel swelling and the end of life of fuel is determined by the fuel–clad gap closure Sodium-bonded fuel does not undergo much of fuel restructuring at comparable linear powers used for oxide-fueled fast reactors because of the lower central temperature The purity of the sodium bond is very important and is limited to a maximum of 10 ppm High O in sodium reduces thermal conductivity, affecting thermal performance of the fuel A hyperstoichiometric (C/M > 1, carbon to metal ratio) fuel composition is chosen so that it contains some amount of sesquicarbide M2C3 phase (M ẳ U ỵ Pu), which takes care of the decrease in (C/M) ratio with burnup This decrease in (C/M) ratio with burnup may result in the formation of actinide metal phase, which may form low-melting eutectic with the cladding, thereby limiting the life of the fuel pin This is more so for fuels with the components of high plutonium content because, for Pu fission by fast neutron, the fission yield curve shifts to the right, generating more noble fission products Depending upon the temperature of operation (type of bond), they may cause bonding of the fuel with cladding material resulting in fuel–clad mechanical interaction (FCMI) O and N impurities also play important roles, as they act as ‘carbon equivalent’, which decides the carbon potential of the fuel O and N are impurities that are picked up during fabrication and their contents are very important to make the fuel of the desired specification Solubility of O is more in PuC than in UC-rich carbide fuel Hence, plutonium-rich fuel has more O solubility than UC-rich fuel The details of all these aspects are discussed in the following section, which deals with the fuel properties, characterization, fabrication, and postirradiation examination (PIE) The reprocessing and waste management of carbide fuel has also been covered, to give a glimpse of the overall carbide fuel cycle Table shows the typical 59 Table Specifications of high- and low-plutoniumcontaining fuels Plutonium (wt%) Pu/(U + Pu) Oxygen (ppm) Oxygen + nitrogen (ppm) M2C3 (wt%) Density (% TD) Grain size (Pu0.7U0.3)C (Pu0.2U0.8)C 66 Ỉ 0.70 6000 7500 5–20 90 Ỉ $10–12 mm 21.3 Ỉ 0.225 360 $400 12.5 Æ 1.4 80 Æ 12 mm fuel specification for carbide fuels and Figure shows the design of the Pu-rich carbide fuel pin used in FBTR 3.03.2 Physical Properties 3.03.2.1 Thermophysical Properties of Carbide Fuel The thermophysical properties that are of importance and affect the fuel performance are solidus/ liquidus temperature, thermal conductivity, coefficient of thermal expansion, elastic/fracture properties, creep, and hardness at ambient and at high temperatures The solidus/liquidus temperatures along with thermal conductivity limit the fuel operating temperature in terms of linear heat rating (W cmÀ1), and the fuel center which ‘sees’ the highest temperature does not exceed the solidus temperature The liquidus temperature gives some indication of the physical state of the fuel in case of core meltdown under accidental conditions Thermal conductivity is an important factor that determines the rate of heat transfer from the fuel to the clad As mentioned above, these properties also put a limit on the fuel surface temperature Thermal conductivity, though an intrinsic property, varies with a number of parameters which are characteristic to the sample Some of these parameters are density or porosity (shape, size, and distribution), composition, presence of a second phase, grain size, etc Coefficient of thermal expansion is an important design parameter for the fuel pin (both for fuel and cladding material) It depends on the composition as well as the extent of second phase present The stresses generated by the fuel over the cladding material are partly due to the difference in the coefficient of thermal expansion between the fuel and the cladding material The elastic property and the fracture property of the fuel are primarily responsible for the extent of FCMI Hardness 60 Carbide Fuel Angular position of top plug with respect to bottom plug 263Њ View in direction S Ỉ5.1 Ỉ5.1 Top plug 13 13 Leak tight weld (TIG) Spring S S1 Wire weld on the top plug 106 Spring support Insulation pellet SS316 Clad Ỉ 0.76 mm wire wound throughout the length pitch = 90 mm ±1.5 320 515.3 531.5 505.5 Fuel pellet 531.5 Insulation pellet Disk 61.5 1.8 Pellet support tube Leak tight weld (TIG) Bottom plug 13 S2 13 FBTR fuel pin sectional view FBTR fuel pin Figure Schematic diagram of the fuel pin design of a fast breeder test reactor of the fuel determines the extent of FCMI: a softer fuel will exert lower stresses on the cladding, thus have lower FCMI Both thermal-induced and irradiation-induced creep of the fuel also determine the extent of pellet–clad mechanical interaction Creep properties also depend on a number of variables such as composition, presence of a second phase as precipitates, grain size, etc Thermophysical properties of actinide carbide have been extensively described and discussed in Chapter 2.04, Thermodynamic and Thermophysical Properties of the Actinide Carbides of this Comprehensive A summary of the data for high (70%) and low (20%) plutonium-containing fuel12 is given in Table 3.03.2.2 Thermochemistry of Carbide Fuels Uranium forms three compounds with carbon, that is, UC, U2C3, and UC2, out of which only U2C3 is a stoichiometric compound UC is stable over a wide Carbide Fuel temperature and composition range and melts congruently at slightly substoichiometric composition, at 2780 K UC2 is stable in two phases, a-UC2 and b-UC2 U2C3 decomposes into UC ỵ a-UC2 on heating from 2096 to 2110 K and into UC ỵ C below $1400 K The Pu–C system has four compounds: Pu3C2, PuC1–x, Pu2C3, and PuC2 The compound Pu3C2 decomposes into Pu ỵ PuC at 848 K Though the crystal structures of carbides of U and Pu are very similar, phase diagrams of U–C and Pu–C are very different These differences are mainly due to (i) the presence of Pu3C2 compound, (ii) the low stability of PuC compared to UC, and (iii) the high stability of Pu2C3 compared to U2C3 (Chapter 2.04, Thermodynamic and Thermophysical Properties of the Actinide Carbides) UC and PuC are highly dense, face-centered cubic (fcc) packed metal atoms with octahedral holes occupied by carbon atoms The brittleness of carbides is due to alternate close-packed planes of metal and nonmetal atoms, the latter restricting the slip and thereby hardening the crystal In addition, the two p-states of carbon not favor the formation of octahedral ligands; hence many of the fission products are not soluble in monocarbide lattice During irradiation, carbide fuels are known to swell more than the oxide fuel Swelling is a complicated phenomenon, controlled by many factors However, the close-packed structure of monocarbides is known to contribute to higher swelling in carbide fuels, especially at low temperatures (T/Tm < 0.3), as bulky fission product atoms of Xe/Kr cannot be accommodated in the carbide lattice The transport properties of vacancies and interstitials in these structures also add to this problem The presence of the more open type structure of the M2C3 phase in hyperstoichiometric mixed carbide fuel reduces the swelling to some extent The structures of UC and PuC are isomorphous with monocarbides of transition metals and some of the actinide metals, and thus these fission products dissolve in the monocarbide phase to varying degrees An increase in O and N impurities results in an increase of carbon activity and CO pressure of the fuel (Figure 3) O impurities are also known to contribute significantly to the actinide redistribution in the carbide fuels and fuel restructuring during burnup Much experimental data as well as assessed and critically reviewed works are available in the literature investigating the impact of O ỵ N impurities on the behavior of carbide fuels.13–17 In order to understand Table Thermophysical properties of high- and lowplutonium-containing carbide fuels Properties (U0.3Pu0.7)C (U0.8Pu0.2)C Solidus temperature (K) Thermal conductivity (W mÀ1 K) at 1273 K Coefficient of thermal expansion (300–1800 K) Hardness (MPa) at 1250 K 2148 12.0 3023 19.0 13.8  10À6 10.9  10À6 1200 1400 61 ´ 10–16 0.011 T = 1000 K pCO 0.010 aC aC Pu/M = 0.55 pCO ´ 10–16 0.009 0.008 pCO ´ 10–16 0.007 ´ 10–17 aC of SS-clad at 1000 K ´ 10–17 Carbon activity ´ 10–16 0.006 pCO ´ 10 –17 aC 0.005 Pu/M = 0.7 ´ 10–17 600 800 1000 1200 1400 1600 1800 2000 0.004 2200 [O] in ppm Figure Effect of O content and Pu/(Pu + U) on carbon potential and carbon monoxide partial pressure of carbide fuel 62 Carbide Fuel w −45 −50 −55 :8 C pm 00 p 60 ]: N + −60 985 K −40 935 K −35 [O Carbon potential (kJ mol-1) −15 −30 )C ,M U −10 −25 % U2 C -U wt +x C :4 )C C3 0.8 U M % t , u w u0 (P pm (P :8 C3 0p M2 00 , m ]: pp +N 00 [O ]: N + [O x : wt% )C 1+ m, M 2C 6000 pp U0 ]: N + [O (Pu U )C 1+x : (Pu 0.55 0.45 MARK II 15 wt% m, M 2C 3: N]: 6000 pp )C 1+x [O + U u 0.3 MARK I: (P 0.7 SS 31 Pu2C -PuC +x −5 −20 example, MC, MC2 (with high U), and M2C3 (with high Pu) Solubility of nitrogen stabilizes UC2 and PuC2 below their decomposition temperatures Carbon activity of mixed carbide decreases with increase in Pu content because of negative Gibbs energy change for the reaction PuC ỵ U2C3 ! UC ỵ Pu2C3 This also results in Pu enrichment of M2C3 phase Opposite effects of Pu and N ỵ O contents also mean that Pu-rich fuel can accommodate higher nonmetallic impurities than U-rich fuel (Figure 4) The effect of Pu and O ỵ N content on carbon activity can also be seen in the phase diagram of carbide system (Figure 5) Carbon precipitation results in shrinking of the MC ỵ M2C3 phase field with increase in O ỵ N content and its expansion with increase in the Pu fraction Mixed carbide shows actinide segregation, with a higher plutonium content in the M2C3 phase than in the MC phase,20,21 as shown in Figure The segregation of actinides in the two phases MC and MC1.5 reduces with increase in the nonmetallic impurity contents O and N.12 This can be explained from the negative Gibbs energy changes of the reactions Pu2C3 ỵ UN ! PuN ỵ U2C3 and Pu2C3 ỵ UO ! PuO ỵ U2C3 Hence, the presence of O ỵ N impurities stabilizes more U in M2C3 and more Pu in MC Segregation of actinides in two phases is more pronounced in plutonium-rich fuel and the effect reduces with increase in temperature and also with increase in sesquicarbide fraction t% the effect of O and N on the behavior of carbide fuels, it is important to understand their interactions with individual carbides The structures of UC and PuC are isomorphous with UO/PuO and UN/PuN, and therefore these compounds show reasonable solubility in the monocarbide phase Though UO and PuO are not stable compounds, O replaces C and gets stabilized in the monocarbide lattice PuC can accommodate more oxygen ( 65 mol% PuO) than UC ( 35 mol% UO) PuC–PuO and PuC–PuN follow a near-ideal-solution behavior, whereas the UC–UO system shows a negative deviation from ideality and has limited solubility UC–UN and PuC–PuN form solutions over the complete composition range UC–UN solution shows a slight positive deviation from ideality on the UC side Because of the increased ionic character of nitrogen and oxygen compared to carbon, addition of oxygen or nitrogen impurity in UC, PuC, or MC results in slight contraction in the lattice with a small positive deviation from Vegards law.18,19 As discussed in Chapter 2.04, Thermodynamic and Thermophysical Properties of the Actinide Carbides, nonstoichiometry of carbides decreases by substitution of C by O or N MC and M2C3 phases of Pu-rich mixed carbide fuel become stoichiometric at high N ỵ O content (!6000 ppm) Pu-rich hyperstoichiometric carbide fuel is biphasic MC ỵ M2C3 in the temperature range of operation of reactor However, U-rich hyperstoichiometric fuel may have three phases at high temperatures, for PuC-Pu −108 −110 UC-U 600 800 1000 1200 1400 1600 1800 2000 2200 Temperature (K) Figure Comparison of carbon potentials of carbide system for different Pu/(Pu + U) and (O + N) values Carbide Fuel Pu/(Pu+U) = 0.2 Pu/(Pu+U) = 0.55 Pu/(Pu+U) = 0.7 C 0.0 63 1.0 X M 1000 K 0.8 0.2 (U, +C )C 1.5 Pu 0.0 [O pp ]= m ] = 200 p 20 p m 00 [O] pp =2 m pp m 0p pm 0.4 0.2 0.2 [O [O ]= 20 [O ] 00 =2 00 0p pm 00 [O ] =2 (U,Pu) XC (U,Pu) + (U,Pu)CNO 1.0 0.4 C 0.8 O+ )CN 0.6 0.6 Pu (U, (U,Pu)C1 + 0.4 (U,Pu)CNO XMN 0.6 0.8 0.0 (U,Pu)N 1.0 Figure Effect of Pu/(Pu + U) and oxygen impurity on the phase diagram of (U,Pu)–C–N system PuC1.5 PuCNO PuC1.5 PuCNO PuC1.5 1.0 0.9 0.9 0.8 0.8 0.7 0.7 0.6 0.6 0.5 0.5 0.4 0.4 0.3 0.3 0.2 0.1 0.0 UCNO C/M = 1.005 [N] = 1000 ppm [O] = 1000 ppm T = 1000 K C/M = 1.005 [N] = 1000 ppm [O] = 5000 ppm T = 1000 K UC1.5 UCNO C/M = 1.005 [N] = 1000 ppm [O] = 1000 ppm T = 1500 K UC1.5 UCNO xPuC1.5 in MC1.5 xPuCNO in MCNO PuCNO 1.0 0.2 0.1 0.0 UC1.5 Figure Effect of O and N contents and temperature on segregation of U/Pu in (U,Pu)C and (U,Pu)C1.5 phases In case of M ỵ MC system, when metal phase precipitates, the metal phase is richer in Pu, but this effect decreases with increase in temperature In this phase field, increase in nonmetallic impurities results in a slight increase in segregation Segregation behavior is also important for back-end processing of the mixed carbide by oxidation and dissolution in HNO3 Excess oxidation of the carbide leads to the formation of M3O8 phase, which has limited Pu solubility This results in excessive segregation of plutonium in the MO2 ỵ x phase, giving problems during dissolution in HNO3 64 Carbide Fuel 3.03.3 Fabrication of Carbide Fuel Fabrication of carbide fuel on commercial scale is a difficult task and needs additional care because of its pyrophorocity apart from high radio toxicity, and the concern for criticality restricts the batch size Moreover, carbide powders formed during carbothermic reduction of oxides are prone to oxidation and hydrolysis This requires high-purity inert-gas cover (nitrogen or argon) in the fabrication line consisting of glove boxes The O and moisture content should be less than 25 vppm (each) to minimize O pickup during the fuel fabrication process and reduce the possibility of any fire hazards due to pyrophorocity A mixed carbide fuel fabrication facility consisting of a series of interconnected glove boxes, maintained under once-through inert (nitrogen) cover gas is shown in Figure In 1960, when research on carbide fuel was initiated, three different methods were followed, namely melting casting, metal hydriding– dehydriding, and carbothermic reduction of oxide The carbide produced by the latter two techniques is processed further by powder metallurgy techniques for the manufacture of fuel pellets 3.03.3.1 Melting Casting Melting casting process, also known as ‘deep casting,’ was followed during the early days of carbide fuel development This process was very reliable because of some advantages over conventional powder metallurgy route; for example, the products were highly dense and very pure In this method, UO2 or U metal chips with graphite are arc-melted and made into the form of a button This button is partially melted many times for homogenization before finally melting and dropping into a mold Pellets as large as 1.8 cm in diameter and 15 cm long could be made by this technique For larger pellets, the skull-melting technique was followed, in which molten carbide forms a shell-type cast in a copper mold and acts as its own containment Melting casting route results in large-grained materials compared to that obtained by powder metallurgical methods.22 Melting casting method is, however, uneconomical due to the high cost of metal fabrication 3.03.3.2 Hydride/Hydrocarbon Route This method is followed for small-scale production of high-purity carbides, where the metal hydride reacts with graphite Actinide carbides MC and M2C3 can be prepared from a mixture of hydride and graphite: PuH2 ỵ 0:85C ! PuC0:85 ỵ H2 gị ẵI The reaction occurs under vacuum at high temperature (1800–2600  C) and this is followed by sintering at 1500  C for h.23,24 For making UC, the reaction between uranium metal and a hydrocarbon gas is carried out at 600–800  C, which yields a fine and easily sinterable product However, the reaction needs to be controlled carefully, as continued flow of methane produces UC2 Hence, the methane flow must be judged fairly accurately to produce only UC In this process, a fine powder of UH3 is prepared by reacting bulk uranium metal with hydrogen at 200–275  C, and the powder formed is decomposed above 430  C to produce fine uranium powder This metal powder is then reacted with methane (or propane) at 600–800  C25–27 to produce UC 3.03.3.3 Carbothermic Reduction Route Uranium monocarbide is produced by carbothermic reduction of UO2 and carbon following the reaction Figure Carbide fuel fabrication facility showing chains of interconnected glove boxes maintained under high-purity inert gas UO2 ỵ 2C ! UC2 þ CO ½IIŠ UO2 þ UC2 ! 2UC þ CO ½IIIŠ In this process, UC2 may be formed as an intermediate product In this method, a homogeneous mixture of UO2 and carbon is blended together and the mixture is compacted at 300–600 MPa pressure along 72 Carbide Fuel Cladding material Fuel Initial fuel–clad gap Fragmented fuel Reduced fuel–clad gap (b) Beginning of life stage ‘A’ (a) As-fabricated fuel Porosity IV III I Zone of transition from low to high swelling Wedge-type crack Swelled fuel No fuel–clad gap (c) End of stage B, Zone I: cracks transformed to porosity, wedge cracks in Zone III and Zone IV Figure 12 Cross-section of helium-bonded carbide fuel pin irradiated to 11.2 at.% burnup indicating three structural zones: (a) as-fabricated fuel, (b) high-density central rich in ‘Pu’ and metallic fission product, and (c) porous finegrained with coarse and fine grain Reproduced from Blank, H Material Science and Technology, Nuclear Materials, Pt I; Cahn, R W., Haasen, P., Kramaer, E J., Eds.; VCH Publisher: New York, 1994; Vol 10A T/Tm > 0.5, and after the gap closure the fuel temperature drops and major part of the fuel operates below 0.5 Only the inner part operates at a temperature >0.5 For helium-bonded pin, the thermal stability limit of the as-fabricated fuel can be matched to the initial fuel operating temperature and some advantages of the hot fuel can be achieved For the purpose, the porosity in the central zone can be increased at the cost of fuel–clad gap, thus enhancing the initial short period ‘A.’ The porosity for gas release takes place at the later period of ‘C.’ The outer zone in contact with the clad remains cool during the period ‘C’, and the swelling is partly accommodated into the porosity of the as-fabricated fuel 3.03.4.1.3 Performance of Na-bonded and He-bonded fuel pins 3.03.4.1.3.1 Sodium-bonded pin For the sodium-bonded fuel pin, the fuel density is higher than that of He-bonded fuel and the fuel–clad gap is more So, under normal operating conditions for a linear heat rating 2000  C The high swelling cannot be fully accommodated by the fuel–clad gap for a high-burnup fuel Increasing the cladding thickness can make the clad stronger to accommodate restraint swelling 3.03.4.1.4 Irradiation creep 3.03.4.1.3.2 He-bonded pin The performance of He-bonded fuel to a great extent depends on the design parameters: namely, fuel–clad gap, smear density, type of clad, pin diameter, linear power, and burnup He bonding is the most preferred bonding concept of carbide fuel, partly because Na bonding, apart from the cost, may deteriorate with burnup and hence the bond quality The in-pile performance of the He-bonded pin depends to a large extent on the porosity of the as-fabricated fuel, which undergoes structural changes when it passes through two stages of burnup, ‘A’ and ‘B.’ The structural changes at the end of stage B decide the safe burning of the fuel in stage C For He-bonded fuel, the initial temperature rise is much higher at the beginning of life because of the lower thermal conductivity of the He bond compared to Na bond The lower thermal conductivity of the He bond also warrants the reduction of the fuel–clad gap compared to the Na-bonded fuel As a consequence, the fuel density for Hebonded fuel is lower and a fabrication porosity of about 15% is recommended (equivalent to 85% pellet density) The fuel–clad gap closure in the early burnup period results in lowering of the fuel temperature, and free swelling changes into restrained swelling under the contact pressure developed at the clad– fuel interface Hence, the mechanical properties of the fuel and clad (creep, fracture toughness) to a great extent predict the fuel behavior during the remaining period of burnup It has been observed that the Hebonded pin can be operated safely up to 15 at.% burnup (peak burnup 20 at.%) The irradiation experience with various design parameters has been summarized by Matzke.29 Cladding breaches due to FCMI or FCCI can be due to loss of ductility of the clad, carburization of the clad, or fuel swelling It is desirable that the hoop Creep is a time- and temperature-dependent deformation mechanism under stress The source of stresses in the fuel under irradiation is the pressure generated by fission gases produced within the fuel The growth of fission gas bubbles within the fuel results in swelling of the fuel by creep deformation The fuel–clad gap provided in the fuel pin design is utilized to accommodate this swelling known as ‘free swelling.’ However, after the fuel–clad gap closure, free swelling is restrained and a back stress is generated by the clad on the fuel This results in restrained swelling of the fuel, which is accommodated within the available pores of the fuel by creep deformation Carbide has a close-packed fcc structure of the NaCl type Unlike close-packed metals, deformation of carbide requires much higher stresses because of the strong covalent bonding existing between the metal atom and the carbon atom It also needs sufficient thermal excitation to leave the lattice site Accommodation of fuel swelling (restrained) within the porosity of the ceramic carbide fuel results in the extension of the primary creep region and this causes an increase in the strain rate before it reduces and attains the steady-state creep The tertiary creep region lies outside the life span of the carbide fuel and hence is not important Two types of creep deformation are operative in the fuel: temperature-dependent thermal creep and radiation-induced or radiation-assisted creep According to Seitzer et al.56 and Dienst,57 thermal creep starts at 1000  C and dominates over radiation-induced creep However, it was also inferred that the relatively flat and lower temperature profile in carbide due to higher thermal conductivity compared to mixed oxide and the high neutron fluence provides sufficient means for the reduction of stresses in the fuel by radiationinduced creep The details of plasticity of carbide can be obtained in the reviews by Routbort and Singh58 and 74 Carbide Fuel Matzke.29 The steady-state creep curve is a function of stress (s), temperature (T ), composition, grain size, and impurity content in the fuel The steady-state creep rate can be expressed by the relation e¼ Ad Àm sn expDH =RT ị ẵ2 where A, m, and n are constants for a particular composition and structure For hyperstoichiometric fuel, the constants A, n, and DH were estimated by Hall59 and found to be 1.57  1011(hÀ1), 2.4, and 506 (kJ molÀ1) Blank60 recommended that the above equation and the values of the constants be different for different materials and should not be used as an input data for a model These data are specific to the fuel composition and type Apart from the porosity correction and presence of a second phase (higher carbides), inhomogeneous deformation across the cross section of the fuel may also change the mechanisms of the creep deformation Figure 13 shows the steady-state creep rate of UC and MC1 ỵ x at a stress of s ¼ 20 MPa These data have been taken from the assessment of Hall59 and Tokar.61 As indicated by Figure 13, for a small portion of fuel, that is, fuel center to periphery, there will be creep Otherwise, the fuel will mostly behave like a brittle material The creep data presented in Figure 13 represent those of a high-density material and need to be corrected for porosity if they are to be used for any other material The horizontal line shows the temperature-independent irradiation-induced creep for hyperstoichiometric mixed carbide fuel at a stress of 20 MPa The irradiation-induced creep, however, is a function of fluence The temperature of operation of mixed carbide fuel under steady-state condition is also shown in this figure by a horizontal line indicated by ‘Ts’ (fuel surface) and ‘Tc’ (fuel center) For mixed carbide fuel with high plutonium content (>55% Pu), no thermal or irradiation creep data is available in the literature However, Tokar et al.62–64 estimated, qualitatively, the creep behavior of high plutonium carbide fuel from hot hardness data by drawing an analogy of creep and those data Sengupta et al.65 also measured the hot hardness data of carbide as a function of Pu content and sesquicarbide content Hot hardness data are also useful in predicting the FCMI behavior of the fuel when the fuel–clad gap is closed Hot hardness of mixed carbide fuels (PuC: 70 and 55%) were measured in a high-temperature microhardness tester using Vickers pyramid indenters.65 The result (Figure 14) showed decreases in hardness with increase in temperature, with MKII (55% PuC) having higher hardness at all temperatures MKI (70% PuC) shows a sharp decrease in hardness at 1123 K ($0.52 Tm; where Tm is the solidus temperature), indicating the onset of creep deformation For MKII fuel, no such sharp transition was observed The data generated for MKI were in close agreement with those of Tokar et al.62–64 up to 1100 K The M2C3 phase is harder than MC and is uniformly MC1+x (U0.45Pu0.55)C - MKII10 (U0.31Pu0.69)C0.93 63 (U0.3Pu0.7)C - MKI10 (U0.79Pu0.21)C63 1´10-2 104 1´10-4 UC1–x 1´10-5 Irradiation creep s 1´10-6 TC 1´10 Log hardness (MPa) de/dt (h-1) 1´10-3 103 TS -7 10 104/T (K-1) Figure 13 Steady-state creep data of UC1Àx and MC1 + x at a stress of 20 MPa Ts and Tc are the fuel surface and center temperatures, respectively Reproduced from Blank, H Material Science and Technology, Nuclear Materials, Pt I; Cahn, R W., Haasen, P., Kramer, E J., Eds.; VCH Publisher: New York, 1994; Vol 10A 200 400 600 800 1000 1200 1400 1600 Temperature (K) Figure 14 Hot hardness data of uranium–plutoniummixed carbide fuels Carbide Fuel distributed in the MC phase, hindering dislocation motion and hence reducing creep deformation It is recommended that the amount of M2C3 phase in the fuel should be optimized to allow creep to occur Though Pu-rich carbide fuel is harder than U-rich fuel, beyond a temperature of 1553 K (average volumetric temperature of fuel), the plutonium-rich fuel behaves in the same way as the U-rich carbide fuel, for which in-reactor performance indicated no failure due to FCMI Hence, from extrapolation it could be presumed that the Pu-rich fuel will also behave in a similar manner 3.03.4.1.5 Experience on irradiation performance 3.03.4.1.5.1 US experience The US experience of carbide fuel irradiation has been summarized recently by Crawford et al.66 A large number of He-bonded and sodium-bonded pins were irradiated in EBR II as an initial test for subsequent loading in FFTF to simulate conditions for a large development plant (LDP) The irradiation condition simulates the peak-cladding temperature condition and the peak power condition of FFTF The objective of this study was to see the effects of pin diameter (7.89–9.40 mm), pellet density (81 and 87% TD), pellet–clad gap (0.13–0.78 mm), type of cladding alloy, and bond type (He or Na) on in-reactor lifetime, fuel and cladding swelling, fission product behavior, fuel–cladding mechanical interaction, and fuel restructuring About 470 MC fuel rods with sodium or helium bonding were irradiated in EBR II with different types of cladding: for example, SS316, PE-16, stainless steel D9, and D21 Over 200 MC fuel rods were irradiated in FFTF in two assemblies: the ACN-1 experiments with rods fabricated using 316SS and D9 cladding; and the FC-1 test, a full-size, 91-rod FFTF assembly with 316SS and D9 cladding and ducts The AC-3 test consisted of 91 full-size, D9-clad rods of which 25 rods contained the sphere-pac fuel made in PSI, Switzerland, and 66 rods contained pellet fuel irradiated to at.% burnup without breach The higher thermal conductivity of carbide makes the average fuel operating temperature very low and it behaves like a brittle material showing cracks However, the cracks resulted in fuel relocation and did not cause premature pin failure In the EBR II test 21, fuel breaches were observed before reaching the goal burnup Out of this, 15 were PE-16-cladded rods, and clad failure was attributed to irradiation embrittlement of the cladding alloy, rendering it less capable of enduring the stress 75 induced by FCMI and fission gas pressure The FC-1 FFTF experiment (a full-size, 91-rod FFTF assembly) attained goal burnup with breach A peak fuel burnup of 20 at.% in 10 MC fuel rods clad in type 316 stainless steel was achieved in EBR II Of those rods, five had experienced a 15% transient-overpower test in EBR II after attaining 12 at.% burnup Thirteen other He-bonded rods and three Na-bonded rods attained 16 at.% burnup in EBR II without breach The FFTF AC-3 experiment results showed that, for the relatively low-temperature conditions of the test, the pellet fuel and sphere-pac exhibited only minor observed differences in behavior, and both types of fuel performed in a manner consistent with the rest of the MC fuel database Carbide fuel failures typically result from FCMI, due to high fuel swelling, which leads to early fuel–cladding gap closure Also, because it generally operates at relatively low temperature, fuel creep is not effective in relieving cladding stress For this reason, MC fuel pin design must incorporate a large fuel–cladding gap and make use of a low-density fuel to delay the onset of FCMI No fuel failures have been attributed to the cladding carburization phenomenon Ten transientoverpower tests of MC fuels were conducted in TREAT using fuel irradiated in EBR II (burnups up to 12 at.%) for the purposes of establishing that cladding breach would occur above 115 and 125% overpower The results suggested FCMI-induced breaches, but most importantly indicated comfortable margins to failure The rods indicated only small cladding strains and small amounts of liquidphase penetration of the cladding The conclusion of these tests was that nothing in fuel transient-overpower response would prevent or limit application of MC fuels to fast reactors The EBR II tests also included rods irradiated beyond goal burnup to breach, and one intentionally defected rod irradiated for 100 days beyond cladding breach The defected rod exhibited a reaction between the fuel and the coolant (O in the coolant), which resulted in a higher specific-volume reaction product and caused expansion as well as widening of the defect with no release of activity into the coolant Other rods irradiated to natural breach in EBR II did not exhibit that phenomenon MC fuels appear to operate benignly after cladding breach In another experiment, in EBR II, the effect of sodium void was tested by irradiating an MC fuel rod with a purposely induced void, simulating void formation due to Na expulsion during irradiation The fuel rod exhibited microstructural changes, reflecting a local high fuel temperature 76 Carbide Fuel and no loss of cladding integrity This experiment indicated that MC fuel would withstand an Na bond expulsion of some magnitude The US experience with MC fuels was not very large; however, it was sufficient to instill confidence that such fuels have irradiation performance adequate for use in SFRs In particular, He-bonded rods clad in lower swelling cladding alloys with around 80% fuel smeared density appeared to show the best performance potential Kummerer54 summarized the results of irradiation experiments under different conditions based on a reference concept Altogether, 101 pins were irradiated, out of which 83 pins were irradiated under steady-state condition, which includes 15 failed pins Eighteen pins underwent steady-state and cycling, which reported one pin failure The parameters, for example, pin diameter, type of bond, fuel–clad gap, smear density, cladding material and wall thickness, linear power, and burnup, were varied Based on the results of these irradiation, a reference pin concept with cold-worked austenitic steel (1.4970) cladding, pellet diameter 7.0 mm, pellet density 84% TD, fuel–cladding gap of 400 mm, helium bond, smear density 75% TD, pin diameter 8.5 mm, and clad wall thickness of 0.55 mm evolved for further irradiation for final performance test The ‘NItrure Mixte dans a Phenix (NIMPHE)’ fuel irradiation experiments were designed to study the influence of the fabrication route on the fuel stability The direct pressing route, as proposed by Richter,67 produces pellets with improved thermal stability; however, the starting materials, the reaction temperature, and the pressing conditions, all influence the thermal stability Also, direct pressing of granules decreases the densification under irradiation by direct pressing and irradiated at a linear power of 730 W cmÀ1 with a burnup of 5.8 at.% is being carried out and results are awaited A short-term irradiation experiment at the beginning of life in the High Flux Reactor Petten was carried out for He-bonded mixed-carbide low in oxygen (CARLO) and rich in oxygen (CARRO).53 The U:Pu ratio was 80:20, the O content was 290 and 1590 ppm, and the carbon content was 4.85 and 5.05 wt% for CARLO and CARRO, respectively The pellet diameter was 6.4–6.2 mm The clad ovalization was Ỉ4 and Ỉ6 mm before irradiation and Ỉ16 and Æ20 mm after a burnup of 0.5 at.% for CARLO and CARRO, respectively The main objective of this test was to study the thermal stability of the fuel Axial g-scanning of the pins indicated that for CARLO the 137Cs activity was more at the extremities of the pin having a lower temperature It was also observed for CARRO, but the concentration profile was less pronounced For CARLO fuel, wedge type crack was observed at the periphery, indicating strong difference in the swelling between the fuel center and periphery CARRO did not show any such crack indicating swelling at the periphery There was no fuel–clad gap for CARLO but for CARRO a gap existed, indicating in-pile densification at high temperature From the analysis of the pellet geometry after irradiation, it was observed that in-pile densification in He-bonded carbide fuel is not a general phenomenon but is restricted to unstable CARRO fuel which is exposed to high temperature at the beginning of life and that this condition was expected to exist for fuel pins with large initial gap and/or higher linear heat rating O plays a very vital role in stronger radial transportation of Pu and less axial Cs transportation and increase in FCCI by carburization of the clad The excess O combines with Cs and forms a complex, thus reducing the mobility of either O or Cs 3.03.4.1.5.3 The NIMPHE irradiation experiment 3.03.4.1.5.4 Fernandez et al.28 reported the influence of three different fabrication routes (conventional carbothermic reduction and direct pressing of pellets and granules) on the irradiation performance of carbide fuels having densities between 80% and 85% TD in the Phe´nix reactor NIMPHE consisted of a capsule containing two (U0.8, Pu0.2)C pins along with some mixed nitride pins irradiated at a higher linear power of 730 W cmÀ1 The pin geometry was similar to Superphenix standard, with pellet and pin diameters of 7.11 and 8.5 mm, respectively Postirradiation experiments of NIMPHE 2, a carbide fuel fabricated MC fuel containing 70% and 55% PuC was used as the driver fuel for FBTR, India This unique fuel composition created many differences in the performance of the fuel This prompted stringent fuel specification at the beginning of the campaign and was followed by a very conservative approach of raising the burnup and LHR with intermittent PIE.68 PIE started with experimental fuel pin irradiation at the beginning of life followed by fuel subassembly examination at burnups of 2.5, 5, 10, and 15.5 at.% Evaluation of the radiograph of the fuel pins after 2.5 and at.% revealed presences of gaps between the pellets 3.03.4.1.5.2 German fast breeder project Indian experience Carbide Fuel (a) 77 (b) (U0.8 Pu0.2)C1.0 (U0.8 Pu0.2)C1.1 Figure 16 Cross-section of mixed carbide fuel after 1.5 at.% burnup Reproduced from Suzuki, Y.; Arm, Y.; Handa, M.; Shiba, K Research and development of uranium plutonium mixed carbide and nitride fuels at JAERI; IAEA TECDOC-466; Vienna, 1988; pp 73–82 (c) (d) Figure 15 Photomicrographs of (U0.3Pu0.7)C1 + x fuel pin at different burnups: (a) 25 GWd tÀ1, (b) 50 GWd tÀ1, (c) 100 GWd tÀ1, and (d) 155 GWd tÀ1 Reproduced from Sengupta, A K.; Basak, U.; Kumar, A.; Kamath, H S.; Banerjee, S Experience on mixed carbide fuel with high plutonium content for Indian fast breeder reactor- an overview; Journal of nuclear materials 385; 2009; pp 161–164 and clad After 10 at.% burnup, the pellet–clad gaps were not observed at the center of the fuel column The maximum increase in stack length was 2.61% Radiography of fuel pins after 15.5 at.% burnup showed pellet–pellet gap at the end of the column, and the stack length increase varied from 2.7% to 3.7% Maximum fission gas release was 16% and the corresponding internal pressure in the fuel pin was 2.1 MPa after 15.5 at.% burnup The Xe/Kr ratio was around 13 The cross-sections of the fuel at various burnups up to 15.5 at.% are shown in Figure 15 It indicates circumferential cracks due to thermal stresses, and discrete zone free from any porosity is observed near the outer circumferential area This was attributed to creep of the fuel due to FCMI The end of the fuel column reveals circumferential cracking and fuel–clad gap closure 3.03.4.1.5.5 Japanese experience The capsule irradiation tests of carbide fuels (C/M ratio of 1.0 and 1.1) started in 1983 in JFR and JMTR.69 Three capsules of carbide fuels containing two pins in each were irradiated at about 1.5 and at.% burnup The pin diameter was 6.5 and 9.4 mm, and the linear heat rate was 450 and 650 W cmÀ1 After 1.5 at.% burnup, characteristic cracking and restructuring of the pellet (Figure 16) took place at the central part of the pellets with the formation of large pores Migration of semivolatile Cs to plenum and other cooler parts of the pin was detected Clad carburization at the grain boundary near the surface was also observed The results of the other reported irradiation up to at.% are awaited 3.03.4.1.5.6 French experience Between the years 1960 and 1970, about 80 MC fuel pins (80% sodium-bonded), including two NIMPHE pins in collaboration with TUI as mentioned above, were irradiated in MTRs (OSIRIS, SOLOE) and then in Rapsodie and PHENIX reactors.70 The sodiumbonded fuel had a high density (>90% TD) but low smear density of about 70–80% TD with a large fuel– clad gap The fuel could reach a burnup of 14 at.% without a limited number of clad failures The swelling rate was $1.6%/at.% burnup below 1000  C They perceived some major design risks, such as lack of sodium bond, locally resulting in overheating of the fuel, excessive swelling, and hence a possible clad failure Clad carburization in case of sodium-bonded pin was also presumed to be an issue which might cause clad embrittlement Also, trapping of small carbide chips in the large gap between the fuel and the clad may also cause localized clad straining For the He-bonded pins, it was observed that smear density should be limited to 70–75% TD for high LHRs (800–1000 W cmÀ1) and high burnup operation In summarizing the international experience on carbide fuel performance, the following conclusions can be drawn:  Sodium-bonded fuel design has some intrinsically good performance characteristics, such as higher 78 Carbide Fuel LHR compared to the He-bonded design; however, it has several drawbacks such as a more complex process of sodium-bonding technique, their more stringent quality assurance, and a greater extent of clad carburization  A He-bonded pin design can achieve better performance in terms of LHR and burnup provided it fulfills the design criteria concerning carbide fuel fabrication, pin design, and the balance between fuel–clad gap width and fuel porosity  Swelling of carbide fuel is an issue when compared with oxide fuel Both solid fission products as well as fission gases dissolved in matrix contribute to swelling Swelling rate increases by growth and merging of bubbles and it is a function of temperature Beyond a critical temperature, swelling increases drastically 3.03.4.1.6 Fuel–clad chemical interaction Carbon activity and partial pressure of CO are important parameters responsible for clad carburization, which make the clad brittle, resulting in clad breach In sodium-bonded fuel, carbon transfer from the fuel to the clad takes place by dissolution of carbon in sodium liquid However, in He-bonded fuel pins, carbon transfer takes place through CO After 2–3 at.% burnup, the fuel may swell and come in direct contact with the clad But, transport of carbon down the temperature gradient within the fuel also takes place through CO Thermodynamic computations and experimental studies on the chemical compatibility between mixed carbide, SS316, and sodium coolant have been restricted to uranium-rich (U/(U ỵ Pu) > 0.7) compositions The results of the out-of-pile and inpile experiments as summarized by Ganguly and Sengupta71 are given below:  The chemical compatibility between mixed carbide fuel and sodium was excellent Out-of-pile test with helium-bonded MC1 ỵ x containing up to 12 wt% M2C3 at 873 K for 1000 h showed no clad carburization, but at 1000 K carbide precipitation along the grain boundaries and slip lines in a zone up to 50 mm in depth from the fuel–cladding interface, was observed The hardening of the SS316 cladding up to 100 mm was observed Similarly, out-of-pile experiments with sodium-bonded MC1 ỵ x showed no carburization at 873 K, but at 1000 K an increase in carbide precipitation in grain boundaries and along slip lines was noticed up to 90 mm of the surface The depth of carburization for sodium-bonded fuel was more compared to the He-bonded pins  Out-of-pile burnup simulation experiments showed that neither individual fission products nor the global chemical effect of burnup was detrimental to clad carburization  In-pile experiments in EBR II up to 12 at.% burnup with stoichiometric MC and hyperstoichiometric MC containing up to 20% M2C3 revealed that the carburization in sodium-bonded carbide elements was 2–3 times greater than that in helium-bonded elements The carburized layer in sodium-bonded pins was as high as 150 mm However, no fuel element failures in both He- and Na-bonded pins had been attributed to clad carburization Saibaba et al.17 have estimated the carbon potential and Pco of hyperstoichiometric (Pu0.7U0.3)C1 þ x up to 2200 K as a function of O, N, and M2C3 contents, and Pillai et al.72 have reported the carbon potential data of SS316 up to 1000 K These calculations show that  the carbon potential of hyperstoichiometric (Pu0.7U0.3)C containing relatively high M2C3 (À20%) is lower than that of (U0.8Pu0.2)C with very low M2C3($2%);  the Pco of (Pu0.7U0.3)C1 ỵ x containing relatively high oxygen (1 wt%) is lower than that of its uranium-rich counterpart with very low oxygen ($0.1 wt%);  carbon activity of the fuel increases with increase in temperature but that of SS316 decreases with temperature In the MC ỵ M2C3 phase mixture, plutonium concentration is greater in the M2C3 phase On extending the same basis of calculations to hyperstoichiometric (Pu0.7U0.3)C1 ỵ x, the extent of plutonium segregation in the M2C3 phase has been found to be more than 90% The mixed sesquicarbide in this case could, therefore, be assumed to be pure Pu2C3 On the basis of thermodynamic data, the calculated carbon potentials of PuC, UC, U2C3, and Pu2C3 as a function of temperature are shown in Figure The carbon potentials of PuC and Pu2C3 are lower than that of SS316 up to 1300 and 1200 K, respectively The carbon potential of UC is lower than that of PuC, Pu2C3, and SS316 at all temperatures but the values for U2C3 are relatively high The carbon potential of hyperstoichiometric plutonium-rich mixed carbide has been found to be lower than its uranium-rich counterpart at all temperatures Theoretical calculations, therefore, indicate that Carbide Fuel 79 The cooling time is an out-of-pile phenomenon, and it has been observed by Blank et al.53 that the amount of Mo, Pd, and Nd increases while that of Pu and other fission products decreases The types of fission products and their chemical state are given below:     Indentations Figure 17 Microstructure of SS316 cladding material after compatibility test Reproduced from Sengupta, A K.; Basak, U.; Kumar, A.; Kamath, H S.; Banerjee, S J Nucl Mater 2009, 385, 161–164 hyperstoichiometric (Pu0.7U0.3)C containing high M2C3 and oxygen are not likely to cause any major carburization of SS316 cladding This was also observed in the microstructure of SS316 cladding (Figure 17) after the out-of-pile experiments carried out by Ganguly and Sengupta.71 3.03.4.2 Effects of Burnup on C/M Ratio and Chemical State of Fission Product The fission of uranium–plutonium MC results in the formation of different types of fission products, but mainly lanthanides, rare earths, and noble metals The chemical state of the fission products will affect the in-pile behavior of the fuel to a large extent The extent of the fission products formed will depend on the burnup and the cooling time of the fuel The fission products generated may form dicarbides or monocarbides depending upon the carbon potential (C/M ratio) of the fuel prevailing in the fuel crosssection and the temperature gradient This may cause decrease in the C/M ratio and result in the formation of metal phase U/Pu, which forms low-melting eutectic with the components of the cladding material Fission products and their chemical states will also depend upon the Pu content of the fuel Fast fission of Pu results in the formation of more noble metals The chemical state of each element and their quantity (depending upon their half lives) depend on     fuel composition, neutron spectrum, residence time of the fuel in the reactor, and cooling time Xe, Kr: inert gas Cs, Rb, Te, I: volatile fission products Sr, Ba: alkaline metals form dicarbide Zr, Mo, Ru, Pd, Y, Nb, Tc, Rh, Ag: 4d metals; mono-, di-, and sesquicarbide  Ce, Nd, La, Pr, Pm, Sm, En, Gd, Tb: monoand sesquicarbide: Lanthanides form mono- and sesquicarbide Hence, the resultant C/M ratio after irradiation will decrease due to formation of the higher carbides of alkaline-earth metals, the 4d metals, and the lanthanides This decrease in C/M ratio may result in shifting of the hyperstoichiometry fuel to hypo and the resultant metal phase may lead to formation of low-melting eutectic Hence, a hyperstoichiometric fuel is always preferred in fuel stoichiometry The extent of hyperstoichiometry (M2C3 content) in the initial fuel composition is determined by the target burnup In their investigation based on an experimental measurement for (U0.85Pu0.15)C1 ỵ x, Blank et al concluded that at 10 at.% burn up, the mean C/Me, where, Me stands for sum of the actinide and the dissolved fission product atom, decreases to less than in hyperstoichiometric carbide Beyond 10 at.% burnup, the decrease in C/M ratio and formation of the metal phase of actinides will be a serious issue from the point of view of low-melting eutectic formation To have an in-depth understanding of the variation of C/M ratio with burnup, theoretical calculations were carried out by Agarwal and Venugopal.73 They observed that the chemical state of rare earth metals plays a crucial role in controlling the carbon activity of the fuel with burnup Rare earth elements show more solubility in M2C3 than in the MC phase of the fuel Lanthanides make stable sesquicarbides and dicarbides The carbon potential of Pu2C3/PuC is higher than that of LnC2/Ln2C3, and therefore LnC2 can coexist with MC ỵ M2C3 fuel But the small difference in Gibbs energies of formation of the compounds is offset by dissolution of Ln2C3 in M2C3 phase of fuel However, LnC2 is also known to dissolve completely in an isomorphous MC2 phase, but absence of MC2 phase in most of the carbide fuels destabilizes LnC2 Lanthanides with reasonable fission yield not form stable monocarbides, but they get stabilized to a limited extent by 80 Carbide Fuel dissolution in MC lattice ($6 mol.%) Ln2C3 shows limited solubility in isostructural U2C3 but considerable solubility in Pu2C3.74,75 It was observed that, in high-burnup, irradiated, and simulated carbide fuels, the solubility of lanthanides in carbide fuel is much lower than the sum total of solubilities of the individual lanthanide carbide In Pu-rich carbide fuel, higher solubility of lanthanides in the sesquicarbide phase compared to the monocarbide phase results in a decrease in carbon potential of the fuel with burnup and this effect is more pronounced at high temperatures With increase in burnup, the difference in carbon activity of the center and the surface of the fuel decreases, which should result in reduced mass transfer of carbon from the center to the surface In case of U-rich fuel, limited solubility of Ln2C3 in U2C3 will result in a smaller decrease in carbon potential with burnup Depending upon the temperature of fuel, U-rich fuel with MC2 phase will stabilize LnC2 and reduce the carbon potential of the fuel Kleykamp76 indicated that, depending upon the amount of O present in the fuel, there could be formation of oxides of lanthanides instead of carbides and this would prevent decrease in the C/M ratio to some extent The sesquioxides of lanthanides are more stable than the oxides of uranium or plutonium In fact, both lanthanides and alkaline earths act as oxygen getters in the carbide fuels Thus, with increase in burnup, the oxygen dissolved in the MC phase reduces The PIEs of carbide fuel reported in the literature have also confirmed the presence of oxides of lanthanides and alkaline earths Thus, the formation of stable sesquicarbide of lanthanides decreases the carbon potential of the carbide fuel with burnup, but oxygen in the fuel binds some of the lanthanides, thereby diluting this effect on the carbon potential Another important chemical species detrimental to carbide fuel performance is iodine, which is very corrosive for the SS cladding Gotzman and Ohse77 carried out out-of-pile experiments on the chemical compatibility of UC with Cs, Te, I, and Se simulated fuel corresponding to 10 at.% burnup They concluded from their experimental results that carburization was possible only when the virgin fuel (no simulated fission product) was in contact with stainless steel However, with Cs, Te, I, and Se addition to the fuel UC1 ỵ x, there was no incompatibility This is due to the high thermodynamic stability of gaseous Cs2I2 and CsI; most of the iodine available in the irradiated carbide fuel is present as Cs compounds, resulting in a drastic reduction in the corrosive effects of iodine However, the g-radiation field of the reactor dissociates cesium iodide molecules, which is followed by ionization of iodine,78 thereby increasing its pressure by a factor of $107 in a FBR.79 Therefore, iodine is still an element of concern in carbide fuels, though its effect is significantly reduced due to its bonding with Cs However, fuel–clad chemical compatibility will not be an issue, as the life-limiting factor in carbide fuels will be FCMI at a stage when the fuel– clad gap is closed.73 3.03.4.3 MA-Containing Fuel Advanced fast reactor fuels (carbide or nitride) such as metal or oxide fuels are likely to contain recycled MA, for example, Am, Np, and Cm They are also referred to as transmutation fuels The goal of development of such a fuel is to develop and qualify a nuclear fuel system that performs all the functions of a traditional fast-spectrum nuclear fuel while destroying recycled actinides The transmuted fuel has a high burnup potential, resulting in long, uninterrupted burning of the fuel thereby reducing the spent fuel volume and eventually the fuel cycle cost The potential of fully closed actinide recycle is well understood, as the separation of the MA from the spent fuel and their recycling substantially reduce the long-term radiotoxicity in the waste, for example, from 10 000 to 500 years The Superfact experiments performed by CEA and ITU in the Phenix reactor using oxide fuel are considered to be a milestone in this respect The irradiation performance database for MA-containing MC or MN fuels is limited and in the very initial stage of development compared to mixed oxide or metal fuels Carbide fuels can be used in MA incinerator reactors because of their hard neutron flux Very limited information is available on the carbides of MA Np forms stable monocarbide and sesquicarbide, whereas only sesquicarbide of Am is known so far The Gibbs energies of formation of these carbides are similar to those of uranium and plutonium carbides (DfG (1000 K) in kcal molÀ1 ¼ À25.4 (UC), À19.3 (NpC), À12.7 (PuC), À66.5 (U2C3), À51.2 (Np2C3), À38.5 (Pu2C3), À18 (Am2C3)).80 As MA carbides and U/Pu carbides are isostructural, carbides of MA are expected to dissolve in fuel matrix to some degree As the fission yield of MA is very low, they will not affect the fuel’s thermochemical behavior in a regular fuel But in case of incinerator carbide fuels with a high fraction of MA, the stability of MA in different Carbide Fuel phases will have an appreciable effect on the carbon activity of the fuel The procedure for the fabrication of MAcontaining carbide fuel is the same as that of conventional carbide fuel fabrication, that is, by carbothermic reduction of the oxide feed materials In this case, the feed material should be (U,Pu,MA)O2 The reduction temperature is around 1500  C under low vacuum to prevent volatilization loss of Pu and Am An additional consideration for pellet-type fuel fabrication is the effect of radiation and decay heat on the binders used during pellet pressing Traditional organic or metalorganic binders tend to decompose quickly upon incorporation into a system with high MA loading High a activity, high decay heat, high neutron emission rate, and high g activity, all require special consideration during the fabrication of MA-containing fuels During fabrication, one needs to take precautions such as use of biological shield from g and neutron emitters and containment to prevent the risk of incorporation of toxic radioactive material into the body The presence of MA in fuel necessitates the use of hot cells or hybrid laboratories, combining standard glove boxes with remote operation by telemanipulators Lead (Pb) glasses provide shielding from g emitters, and a combination of water, lead, and cadmium (or boron) provides shielding from neutrons emanating directly through radioactive decay (e.g., 244Cm) or by (a, n) reactions, which typically occur with light elements (O, N, or C) in the fuel Handling of MAs and fabrication of MA containing fuel is best undertaken in a dustfree atmosphere, and hence processes that reduce dust production are not just desirable but essential In this respect, sodium-bonded fuel with large pellet–clad gaps allows more tolerances in the fuel pellet diameter and elimination of the centerless grinding step which generates waste in the form of dust and contaminates the cell’s inside surface The wet route with automated handling facility in a hot cell under nitrogen cover gas is the best possible option In the ITU, Karlsruhe,28 a qualified advanced process of fabrication of MAcontaining fuel exists, which consists of glove boxes shielded by water (50 cm) and lead (5 cm) for handling a limiting mass of 150 g of 241Am and g 244Cm In this process, all operations are carried out by remote control telemanipulators The irradiation behavior of MA-containing fuels is different from that of conventional fuels Increased fuel pin helium gas inventory is expected due to capture and decay sequences associated with 241 Am and 242Cm, which decay by a emission (halflife of 169 days) to 238Pu This could lead to higher 81 fuel swelling and is also an additional source for fuel pin overpressurization The data on in-pile performance of MA-containing carbide fuel are limited, unlike oxide or metal fuel, and the scope of its use is still wide open Under the US fast reactor program, reference and alternative 1000 MWth advanced burner reactor (ABR) core concepts were developed using ternary metallic (U–TRU–Zr) and mixed oxide fuel with UO2 ỵ TRU However, better thermal and neutronic properties of carbide and nitride fuels prompted them81 to consider TRU carbide as an alternative fuel form ABR core concepts with these fuels were developed based on the reference 1000 MWth ABR core with the expectation that the carbide fuel can mitigate the disadvantages of both metallic and oxide fuels in the ABR: for example, favorable passive safety features in a severe accident compared to the oxide core, a higher discharge burnup compared to the metallic core, and a potential to increase thermal efficiency For consistency, the potential design goals used in the reference ABR core concepts were also employed in this study: a 1000 MWth power rating, medium TRU conversion ratio of $0.75, a compact core, 1-year operational cycle length at least with a capacity factor of 90%, and sufficient shutdown margin Generally, it has been observed that the core performance parameters of the carbide and nitride fuels in the ABR core lie between the values of the metallic and oxide cores The neutron spectrum is softer than that of the metallic core, but harder than that of the oxide core The fuel resident time of 60 months is longer than that of the metallic core, but shorter than that of the oxide core The required TRU enrichment (25% for the carbide fuel) and average discharge burnup of $100 GWd tÀ1 are also higher than those of the metallic core, but lower than those of the oxide core This indicates that the core can be more compact or the operating temperature increased further to reduce the plant cost The replacement of the bonding material by helium gas increases the fuel temperature by $200  C, but overall impacts on the kinetic parameters and reactivity feedback coefficients are negligibly small As a result, the carbide core satisfies all sufficient conditions for core passive safety that are not met by the oxide core The good thermal conductivity and high melting temperature lead to a significant decrease in the average fuel temperature (comparable to metallic fuel), and hence provide a large margin to fuel melt and favorable passive safety features without additional design fixes that were required in the oxide core concepts 82 Carbide Fuel 3.03.5 Fuel Reprocessing and Waste Management This aspect of the nuclear fuel cycle has been discussed in Chapter 5.14, Spent Fuel Dissolution and Reprocessing Processes; Chapter 5.16, Spent Fuel as Waste Material; Chapter 5.18, Waste Glass; Chapter 5.19, Ceramic Waste Forms; Chapter 5.20, Metallic Waste Forms; Chapter 5.21, Graphite and Chapter 5.22, Minerals and Natural Analogues of this comprehensive and may be referred to for further information This section highlights carbidespecific issues Different nations have different back-end strategies: for example, direct disposal (once-through fuel cycle), storage and postponed decision (wait and watch), and reprocessing and recycling (closed fuel cycle) The reprocessing of the carbide fuel can be carried out by the hydrometallurgy route or the pyrometallurgy route Due to the extensive work carried out on the oxide fuel, especially LWR fuel, the hydrometallurgy route via the PUREX process is the most developed technology for the fuel processing The same process can be used for burnt carbide fuel with some modifications since carbide fuel from an FR differs from the burnt fuel of LWR because of the following issues:  high specific activity due to high burnup;  high fissile content;  presence of sodium in the fuel (Na-bonded pins) or on subassemblies (Na-cooled reactors);  enhanced formation of Pt-group metals (PGM), which form insoluble residues along with fissile elements;  pyrophoric nature of carbide fuels; and  formation of some complex organic compounds during dissolution which interfere during separation process by forming a third phase.82 If the source of spent fuel is an MA incinerator, the presence of high concentrations of MA may complicate fuel reprocessing further due to the high g dose The dissolution of carbide fuels in nitric acid has been a subject of considerable interest, as it leads to the formation of a number of complex organic compounds (e.g., mellitic–benzene hexacarboxylic acid, oxalic acid) along with some soluble organic compounds and aromatic carboxylic acids It was found that up to 50% of the original carbon may remain in the solution due to formation of these organic compounds These organic compounds can interfere in the separation process during solvent extraction by forming a third phase Much effort has been devoted in different laboratories across the world to find a solution to this problem It was found that the advanced dissolution process with silver-catalyzed electrolysis could circumvent this problem A study with the use of ozone for the destruction of organic species has also been successful.83 Most of the experiences in the processing of carbide fuel are limited to laboratory-scale experiments A compact hot-cell facility has been commissioned in Kalpakkam, India,84 for reprocessing FBTR spent fuel and to establish the process and equipment for high-burnup, plutonium-rich fuels (Figure 18) For this purpose an advanced Purex process has been developed The fuel pin is chopped and dissolved in an electrolytic dissolver For recycling, Pu and U are separated in the pure form after centrifugation and three cycles of solvent extraction The solvent extraction process is based on the tri-n-butyl phosphate (TBP) solvent Detailed studies have been carried out on the various parameters that influence the third phase formation: for example, effects of carbon chain length of the diluents, acidity of the aqueous phase, etc TAP (tri-n-amyl phosphate)/n-dodecane has been found to have high capacity to load Pu(IV) without third phase formation under normal extraction conditions as compared to TBP/n-dodecane Long-chain monamides are also being investigated as alternate extractants The amides, for example, dihexyloctanamide (DHOA), have special advantages of simple synthesis, complete incinerability, innocuous degradation products, low aqueous solubility, high decontamination factor (DF), and reasonably high U/Pu loading capacity Figure 18 Compact hot-cell facility for mixed carbide (FBTR) spent fuel reprocessing at IGCAR, Kalpakkam, India Reproduced from Natarajan, R.; Raj, B J Nucl Sci Technol 2007, 44(3), 393–397 Carbide Fuel 3.03.6 Summary The potential of carbide as a fuel for fast reactors was well known since its inception Better thermal properties, higher fissile atom density, and compatibility with liquid metal coolants are some of the properties of the carbide fuel that made it a superior fuel compared to oxide fuel These properties result in a higher breeding ratio and lower doubling time However, large-scale utilization of carbide fuel has not been explored so far and its potential as a promising and commercially viable fuel option is yet to be established The challenges of carbide fuel developments Heat load (108 sieverts/t) There are many advanced hydrometallurgical processes in different stages of development Most of the solvent extraction methods are liquid–liquid type extraction and are adaptations of the PUREX process to some extent Pyroprocessing is a generic term used for several kinds of pyrometallurgical reprocessing These processes are the main alternative to ‘aqueous’ processes They can be adapted for carbide fuel directly As this process is already established for metallic fuel, reduction of carbide fuel to metal or conversion to a chloride can prepare carbide fuels for pyroprocessing These processes can be used for reprocessing of fuel with high burnup and are the subject of renewed developmental effort worldwide In these processes, the fuel is first dissolved in a bath of molten salts (chlorides, fluorides, etc.) at very high temperatures (>500  C) and then desired elements are separated by various techniques such as liquid metal extraction, electrolysis, or selective precipitation According to the IAEA definition, waste is any material that contains or is contaminated by radionuclides at concentrations or radioactivity levels greater than the exempted quantities established by the competent authorities and for which no use is foreseen Different countries have different interpretations of ‘waste,’ depending upon their fuel cycle policies The partitioning of MA and important fission products can reduce many concerns of waste disposal.85 The removal of fissile elements and MA addresses the long-term concerns of the immobilization matrices of the wastes, and medium-depth repositories will suffice for disposal of these wastes Due to the hard-neutron flux in carbide fuels, it is possible to use carbide fuel for the purpose of burning MA This can greatly reduce the heat-load and amount of the final high-level waste (Figure 19) 83 U Minor actinides (Np, Am, Cm etc.) Pu Fission products 10 102 103 104 105 106 Time (years) Figure 19 Contribution of different radioactive elements in the burnt fuel to radiotoxicity of waste as a function of time include primarily its pyrophorocity, which needs special care during fabrication while handling the carbide material This issue becomes more critical while handling fine carbide powders during milling, precompaction, and granulation after carbothermic reduction of oxide feed material Use of high-purity dry inert cover gases such as nitrogen or argon in glove boxes in once-through mode has been the general approach Sintered carbide fuel pellets have been made for the FBTR in India and in other countries as well for test pin irradiation using such an approach without much constraint Optimization of process parameters during the carbothermic reduction step to minimize the actinide losses (particularly for MA containing fuel) or plutonium losses by vaporization is another challenge This issue becomes more critical for mixed carbide fuel with high plutonium content: for example, MKI and MKII fuels for FBTR containing 70% and 55% Pu, respectively In such high Pu-bearing fuels, one may have to compromise on the O content (impurity) and Pu losses of the product A high reduction temperature may reduce the O content but will increase Pu losses It is recommended that the oxide feed material be a coprecipitated solid solution of oxides of uranium and plutonium of the type (U,Pu)O2.86 In such cases, apart from the advantages derived from microhomogeneity of uranium and plutonium in the final product, the length of time required for the high-temperature carbothermic reduction step will also be shorter The economics of carbide fuel fabrication is a matter of concern because of an additional step of carbothermic reduction of the oxide feed To make commercial production of carbide fuel an attractive 84 Carbide Fuel and economically viable option, the fuel cycle should be a closed cycle where the technology of fuel reprocessing and waste management also plays a vital role In such a case, the recycling of MA needs to be considered, as it will significantly reduce the waste load and also make the fuel proliferation-resistant The technology for reprocessing of carbide fuel is not well established, although a small-scale facility has been set up and demonstrated by IGCAR, Kalpakkam, India, for reprocessing of mixed carbide fuels based on the PUREX process The problem of enhanced formation of PGM, which form insoluble residues along with fissile elements, and the formation of some complex organic compounds during dissolution, which interfere during separation process by forming a third phase, need to be understood and the process steps need to be modified suitably The fabrication of MA-containing carbide fuels needs remote and automated operations in hot cells A sodium-bonded fuel option having a large fuel– clad gap (having greater tolerance for fuel pellet diameter) will have some advantages over the He-bonded fuel having low tolerance for fuel pellet diameter The centerless grinding step of the fuel pellet diameter, which generates dust, is not required in this type of fuel A remote and automated wet route (gelation method) of production of carbide fuel will be an attractive option, as it would reduce the dust hazards of radioactive powder material Of late, there has been renewed interest86 in the use of carbide fuel for ‘gas-cooled fast reactor (GFR).’ as part of ‘The Generation IV Project.’ The GFR has the advantage of a fast neutron spectrum as well as of high temperature The high outlet temperatures provide improved economy in the power conversion unit and permit process heat application The design and development of an innovative refractory fuel in an advanced ceramic cladding has been the aim of this concept One of the concepts uses mixed carbide pellets (cylindrical pellet or disk with low L/D ratio), and fiber-reinforced SiC–SiCf composite has been selected as cladding material Mixed nitride and oxide has been considered as reserve options The fuel configuration will be either a classical pellet in tube or an innovative plate type: that is, the fuel pellet is disk shaped with a low length-to-diameter ratio in a honeycomb structure and the cladding is in the form of a plate covering both sides A tungsten layer is provided between the fuel and the clad to prevent any interaction between them and also to contain the fission products Helium gas will be used as coolant instead of sodium The use of He as a coolant in GEN-IV reactor concept eliminates the possibility of any fire hazards due to sodium, of which there have been some reported instances (See Chapter 2.04, Thermodynamic and Thermophysical Properties of the Actinide Carbides and Chapter 5.14, Spent Fuel Dissolution and Reprocessing Processes) Acknowledgments The authors express their deep sense of gratitude to Dr Srikumar Banerjee, Chairman, Atomic Energy Commission, Department of Atomic Energy, Government of India, for his guidance, support, and valuable suggestions in structuring this review They also thank Dr V Venugopal, Director, Radiochemistry and Isotope Group, BARC, India, for his keen interest and encouragement during the course of this work Thanks are also due to Mr Kasi Viswanathan, Associate Director, GRIP Group, for providing the irradiation examination data of MKI carbide fuel for FBTR The authors also sincerely thank Ms Mareena Michael for her help in preparing the manuscript Contributions made by staff members of the Radiometallurgy Division and Fuel Chemistry Division, BARC, and Reprocessing Division, IGCAR, Kalpakkam, 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Apostolidis, C.; Glatz, J P J Nucl Mater 1989, 166, 181–188 Natarajan, R.; Raj, B J Nucl Sci Technol 2007, 44(3), 393–397 Schapira, J P Nucl Instrum Methods Phys Res 1989, A280, 568–582 Brunel, L.; Chauvin, N.; Mizuno, T.; Pouchon, M A.; Somers, J The Generation IV Project GFR fuel and other core materials, GIF Symposium, Paris (France), Sept 9–10, 2009 ... and M2C3 can be calculated (precision Ỉ5% of the absolute value): VMC KM C IMC ẳ VM2 C3 KMC IM2 C3 ẵ1 69 3. 03. 3.4.2 Physical quality control 3. 03. 3.4.2.1 Fuel pellet inspection Sintered fuel pellets... as-fabricated fuel 3. 03. 4.1 .3 Performance of Na-bonded and He-bonded fuel pins 3. 03. 4.1 .3. 1 Sodium-bonded pin For the sodium-bonded fuel pin, the fuel density is higher than that of He-bonded fuel and... for mixed carbide fuel fabrication method 3. 03. 3 .3. 2 Sol–gel (wet) route The internal gelation route (wet method )32 ,33 of particle fuel fabrication is dust-free and provides ceramic fuel with