4.1. Introduction
Regarding plutonium and minor actinides (MAs) from spent LWR fuel, several fuel cycle scenarios are envisioned. The LWR plutonium can be seen as a long term asset, promoting a rapid expansion of fast (self-)breeders (CR ≥ 1) and transition to a pure fast reactor scheme. On the other hand, if plutonium and minor actinides are rather perceived as a waste and the desire is to destroy them quickly, the reactors will work in a transuranium (TRU) ‘burner’ mode. These systems will then operate in two-component schemes together with LWRs and/or in a concert with LWRs and sub-critical MA burners in a double-strata scenario. Since the beginning of the nuclear program, about 1 600 tonnes of transuranics have been produced by 2005. Two Generation IV reactors, the sodium-cooled and lead-cooled fast reactors (SFRs and LFRs) are interesting future options that can be used both as (self-)breeders with long burnups and TRU burners. In several countries, the latter could be fast reactors’ first mission
4.2. Method 4.2.1. Design
This study aims at an indicative comparison of SFR and LFR cores with similar safety coefficients (Doppler, coolant temperature reactivity), which also accommodate large fractions of minor actinides and plutonium in the fuel. The latter then facilitates high MA and Pu consumption rates. Based on the proposal of European Lead-Cooled Fast Reactor (ELSY) project [1], the 600 MW(e) power level was chosen as a basis. However, an up-rated version of an LFR burner (900 MW(e)) was also investigated.
For LFRs, the use of the improved supercritical steam cycle was considered, providing a thermal efficiency of 42%. Similarly high thermal efficiency, up to 45%, could be achieved with an SFR employing a supercritical CO2 Brayton cycle. Design parameters concerning pin and pellet are based on established LFR and SFR core designs and summarized in Table 1.
TABLE 1. DESIGN PARAMETERS OF SFR AND LFR-BURNER CORE-CONCEPTS CONSIDERED IN THIS STUDY
The different core parameters were chosen to comply with restrictions imposed by the thermo-mechanical characteristics of the coolant, cladding, and fuel.
Our SFR core is based on a modified model core for the WAC benchmark reactor [2], where the axial and radial reflectors were removed. As sodium allows for significantly higher coolant velocities (8-10 m/s) than lead (2-3 m/s), SFR pin lattices can be much tighter than those for LFRs. This leads to more compact SFR cores in comparison to LFR cores. The original WAC benchmark reactor had a power of 800 MW(e), whereas the one used here has a reduced power of only 600 MW(e) lowering the linear power by 25%. This is a safety relevant change and facilitates the comparison with the LFR.
Averaged inlet and outlet sodium temperatures are 653 and 823 K, respectively.
In order to improve the coolant temperature reactivity coefficient, the concept of a high-leakage, pancake-like core was chosen for the LFR core. Active pin length is 100 cm and the core has the same pin and pellet design as the BREST reactor [3]. An 80 K axial coolant temperature increase in the core is based on the ELSY design. Average coolant outlet temperature is 753 K, which is comfortably below the limit of 870 K, guaranteeing the stability of protective oxide layers under nominal operation. Seismic stability requirements constrain the height of the reactor vessel to 11 m.
The burners operate in concert with LWRs in two-component scheme recycling both Pu and MAs, which are homogeneously admixed to the core fuel. To give an indicative inter-comparison of both systems with (U,TRU)O2-92Mo fuel, a fuel cycle length of 330 days with 35 days refuelling period was tentatively chosen.
Consideration has been given to the option of including uniformly distributed moderating pins (or thermalizing zones) in fast reactors. The reason is the significant deterioration of the coolant temperature and Doppler reactivity coefficients due to the presence of a sizeable amount of minor actinides. By tailoring of the neutron spectra by moderators, the spectral gradient during coolant heat-up/voiding is diminished. Additionally, more neutrons are scattered down to the resonance region, which profoundly improves the Doppler feedback.
For this purpose, hydrides were considered as moderators, but they have the disadvantage of having relatively low decomposition temperature (e.g. ~1100 K under a H2 atmosphere for CaH2), which excludes their incorporation in the fuel directly. In this paper, we investigated introduction of BeO moderators located in dedicated pins within a sub-assembly. It should be noted that, however, that the use of BeO could be problematic due to its high chemical toxicity. Another option would be to use metallic beryllium or 11B4C, which was also considered for the CAPRA reactor.
4.3. Computational model
The Monte Carlo code MCB [4] was used in our neutronic and burnup analyses. Doppler reactivity feedback was estimated by evaluating a reactivity change upon the increase of fuel temperature from 300 to 1500 K. Coolant temperature reactivity coefficients correspond to a change in keff due to a heat-up of coolant in the active core only. The 1-σ statistical deviations in keff were under 10 pcm.
Nuclear data libraries were adjusted for the temperature dependence by the NJOY code. The averaged temperatures of the core components were assumed as follows: 1500 K for fuel, 900 K for cladding, and 600 K for coolant.
The fuel has a burnup of 41 GW•d/tHM and it is assumed to have undergone 30 years of cooling.
Correspondingly, Pu/Np/Am fraction is then equal to 83/5/12. Depleted uranium (0.3% 235U) is used in the analyses.
In order to reach reasonable calculation times in MCB, we have chosen to adjust the system parameters (fissile enrichment) such that keff is one at BOC rather than at EOC. Our calculations thus somewhat underestimate the reactivity burnup swing since the U/TRU fraction would have to be decreased in the latter case. The composition of the actinide vector is that of spent LWR UOX fuel (see Table 2).
TABLE 2. PLUTONIUM AND MINOR ACTINIDE VECTOR CORRESPONDING TO THE LWR UOX SPENT NUCLEAR FUEL WITH BURNUP 41 GW•d/tHM AFTER 30 YEARS OF COOLING
4.4. Neutronic and burnup performance
As already mentioned, CERMET AnO2-92Mo was chosen as fuel for SFR and LFR burner cores.
Volume fraction of 92Mo is kept at 50% due to the reason of fuel fabricability and thermal stability during irradiation. In order to reach considerable TRU consumption rates, TRU fraction should be at least 40-50% (owing to favourable neutronic characteristics). A comparison of neutronic and burnup characteristics of LFR and SFR burners is given in Table 3.
TABLE 3. NEUTRONIC AND BURNUP PERFORMANCE OF SFR AND LFR BURNERS:
DOPPLER AND COOLANT TEMPERATURE REACTIVITY FEEDBACK FOR LFR AND SFR BURNER CORES CORRESPOND TO THE INCREASE OF FUEL AND COOLANT TEMPERATURES BY 100 K. BeO MODERATOR PINS WERE USED. THE TWO LFR DESIGNS DIFFER ONLY REGARDING THE POWER LEVEL. THE FUEL CYCLE PERFORMANCE VALUES CORRESPOND TO THE 1st YEAR OF THE START-UP MODE.
First, we note that the TRU fraction in the fuel had to be higher for an LFR than SFR in order to attain criticality at BOL despite more than twice the actinide inventory in the LFR. This is due to the better neutron economy of the SFR tight lattice.
Burnup reactivity swing and fuel burnup are approximately inversely proportional to initial actinide inventories. A slight departure from the proportionality can be ascribed to different breeding characteristic of SFR and LFR due to different TRU fractions. The SFR is loosing reactivity twice as fast as the LFR, so its fuel has to be reprocessed more often. However, this also means that the SFR has a larger actinide burnup rate than the LFR. Reactivity coefficients are somewhat better for the SFR, where Doppler is about 50% stronger than the coolant temperature reactivity coefficient.
Due to the large neutron mean free paths (4.2 cm in Pb, 12 cm in Na), the impact of moderator pins on the local power peaking is limited and pin-to-pin local power peaking factor remains below 1.1 at BOL.
Concerning TRU consumption, both 600 MW(e) and 900 MW(e) LFRs perform better than 600 MW(e) SFR (Table 4). While a 600 MW(e) SFR can annually transmute only about 260 kg of TRUs, an LFR of the same power level incinerates over 300 kg/a. Understandably, TRU consumption is higher in an up-rated LFR (900 MW(e)) and equals to about 450 kg, of which 315 kg is plutonium and 134 kg MAs. Observe that due to the self-production of plutonium the destruction rate of MAs in the fuel is in fact higher than what would correspond to their share in the initial load.
TABLE 4. TRANSMUTATION PERFORMANCE OF SFR AND LFR BURNERS EMPLOYING URANIUM-BASED AnO2-92Mo CERMETs. AN LFR ANNUALLY CONSUMES ABOUT 300 kg OF TRUs THAT IS ROUGHLY THE ANNUAL PRODUCTION OF A 1.1 GW(e) LWR WITH A FUEL BURNUP OF 41 GW•d/tHM. ALL 242Cm WAS ASSUMED TO DECAY TO 238Pu IN THE SPENT FUEL. THE FUEL CYCLE PERFORMANCE VALUES CORRESPOND TO THE 1st YEAR OF THE START-UP MODE.
In Fig. 1, we compare the transmutation performance of a 600 MW(e) LFR burner with a 600 MW(e) LFR self-breeder (without blankets). The LFR self-breeder employed (U,TRU)O2 mixed oxide fuel (without an inert matrix), average TRU fraction in fuel was 22.6%. We note that in the start-up mode, self-breeder and burner perform similarly concerning the consumption of minor actinides (67 kg/a vs.
88 kg/a for the burner), but while the self-breeder still generates some Pu (14 kg/a), Pu is consumed in the burner (215 kg/a).
It is to be noted that neither of these designs used in the present study were optimized with respect to the burnup reactivity swing performance and fuel management scheme. Particularly, burnup reactivity swing could be reduced for an SFR by increasing fissile inventory through enlarging pin diameter and/or number of core channels. Note, however, that these changes enhance coolant temperature coefficient and void worth and has to be accompanied by a reduction of the MA fraction in the fuel.
This would lead to a lower MA consumption rate.
FIG. 1. Amount of annually consumed transuranics in a 600 MW(e) (U,TRU)O2–92Mo fuelled LFR burner compared to a 600 MW(e) self-breeder employing (U,TRU)O2 fuel (see Domain-I, JRC/IE contribution). All 242Cm was assumed to decay to 238Pu in the spent fuel.
4.5. Safety aspects of waste burners
With regard to safety aspects of TRU waste burners, most of the aspects presented in Domain-I by JRC/IE also hold here. The only important difference is the lower core height for the LFR TRU waste burner (1 m vs. 2 m for Th-fuelled burners/breeders and U-fuelled self-breeders), which leads to better natural circulation in the ULOF case (Fig. 2).
FIG. 2. Temperature evolution during a ULOF accident for three LFR cores: (Th,TRU)O2– fuelled 600 MW(e) burner/breeder, (U, TRU)O2–92Mo fuelled 600 MW(e) TRU waste burner, and (U, TRU) O2–92Mo fuelled 900 MW(e) TRU waste burner. No feedbacks are considered in these STAR-CD calculations.
4.6. Conclusions
In this paper, we first indicated that fuels containing minor actinides and inert matrices could be fabricated and reprocessed. Then, we compared the neutronic, TRU consumption, and safety performance of LFR and SFR burners employing uranium-based CERMET fuels with 92Mo matrices.
Minor actinides were homogeneously mixed into the fuel.
Comparing SFR and LFR cores, the SFR core is notably smaller than that of LFR. The reason is twofold. First, lead has lower capability to remove heat from the reactor core (mainly due to lower permissible velocities), which consequently require higher P/Ds for LFR. Second, lead is an excellent neutron reflector, which provides more freedom to the designer to choose the core geometry. For instance, flatter core geometry can be used that offers better safety performance without loosing too many neutrons. In-core moderators employed in the core sub-assemblies were used in order to improve the safety coefficients (Doppler and coolant temperature reactivity). BeO pins were considered for both SFRs and LFRs.
The main conclusion from these calculations, however, is that the large scale burnup of plutonium is counterproductive in a time when a nuclear renaissance is starting and fissile material will become more and more important.
REFERENCES TO CHAPTER 4
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[2] WIDER, H.U., et al., Comparative Analysis of a hypothetical Loss-of-Flow accident in an irradiated LMFBR core using different computer models for a common benchmark problem, European Commission, EUR 11925 (1989).
[3] ADAMOV, E.O., White book of nuclear power, N.A. Dollezhal Research Development Institute of Power Engineering, Moscow, Russian Federation (2001).
[4] CETNAR, J., et al., MCB — a continuous energy Monte Carlo Burnup code, Proc. Fifth International Information Exchange Meeting, OECD/NEA, Mol, Belgium (1989).
[5] LAIDLER, J.J., et al., Development of pyro-reprocessing technology, Prog. Nucl. Energy Vol. 31, 1/2 (1997) pp. 131-140.
[6] LEE, Y.E., et al., Decision-making and nuclear energy policy: application of environmental tool to nuclear fuel cycle, Energy Policy, Vol. 30, 13 (2002).
[7] PONCELET, F., Research at COGEMA: benefits and a future outlook of the nuclear fuel cycle, paper presented in Intl Conf ATALANTE 2004, Nîmes, France (2004).
[8] SERP, J., et al., Electrochemical behaviour of plutonium ion in LiCl-KCl eutectic melts, Electroanal. Chem, Vol. 561, 143 (2003).
CHAPTER 5. DOMAIN-III: HYBRID SYSTEM (ADS) WITH FERTILE FUEL