Core Design of a High Temperature Reactor Cooled and Moderated by Supercritical Light Water

Transcription

1 GENES4/ANP2003, Sep , 2003, Kyoto, JAPAN Paper 1041 Core Design of a High Temperature Reactor Cooled and Moderated by Supercritical Light Water Akifumi YAMAJI 1*, Yoshiaki OKA 2 and Seiichi KOSHIZUKA 2 1 Department of Quantum Engineering and System Science, The University of Tokyo, Tokyo, , Japan 2 Nuclear Engineering Research Laboratory, The University of Tokyo, Tokai, Naka, Ibaraki, , Japan An SCLWR-H equilibrium core is designed by three-dimensional core calculations. A new square fuel assembly is designed for a reduced local power peaking factor. The fuel assembly burnup calculations and core burnup calculations are carried out in the similar way adopted in BWR core designs. 24 cycles of core burnup calculations are carried out to design an equilibrium core. A fuel load pattern and a control rod pattern scheme are designed so that the reactivity and power distributions are controlled throughout the cycle. In order to maximize the average coolant core outlet temperature, it is important to keep the radial power distribution as flat as possible. The radial power peaking factor of the designed core is The coolant core flow rate distribution is adjusted to the radial power distribution by the inlet orifices. The coolant flow rate for the fuel channels enclosed by the orifice is determined by the maximum cladding surface temperature criterion. The average coolant core outlet temperature is 528 when the coolant flow rate is adjusted to every quarter of the fuel assembly. KEYWORDS: core, fuel assembly, fuel rod, design, power distribution, coolant flow rate distribution, three dimensional, nuclear reactor, concept I. Introduction One of the main goals in designing the core of SCLWR-H is to make its average coolant temperature at the outlet of the core as high as possible. In order to design a core with high average coolant core outlet temperature, the power distribution of the core needs to be controlled to match the coolant flow rate distribution throughout a cycle. The design of such a core requires three-dimensional core burnup calculations representing each fuel assembly separately. In the past design of the SCLWR-H 1), core calculations were carried out in simplified two-dimensional R-Z geometry. The macro cross-section sets, used in the core calculation, were prepared by cell burnup calculations. The linear heat generation rate was assumed to be constant throughout a cycle. The effects of control rods were not considered in evaluating radial and axial power peaking factors. The local power peaking factor was evaluated for a fresh fuel separately from the core calculation, and the effect of burnable poisons was not considered. In this study, a new SCLWR-H core is designed with a three-dimensional core burnup calculation similar to the way a BWR core is deigned. The macro-cross section sets for the core burnup calculation are prepared by fuel assembly burnup calculations, including fuel rods with burnable poisons. The linear heat generation rate is calculated from the core power distribution at each time step of the core burnup calculation. A control rod pattern scheme and a fuel load pattern are designed and considered in evaluating the radial and axial power peaking factors. Local peaking factors are evaluated for the relevant plane of the maximum power fuel assembly throughout a cycle. The past hexagonal fuel assembly design 1) had a large local power peaking factor, because the neutron moderation by the hexagonal water rods were not uniform. The local power peaking factor needs to be reduced to achieve high average outlet temperature. In this study, a new square fuel assembly with a reduced local power peaking factor is designed. The sub-channel is also more uniform than the hexagonal fuel assembly, so it is more suitable for achieving high average outlet temperature. Three-dimensional core design of SCPR, using a modified BWR core simulator, was briefly reported 2). The preliminary results are obtained with BWR design criteria, but comprehensive core design needs to be studied. The result of the transient analysis of the University of Tokyo 3) has shown that the maximum cladding surface temperature at the rated power can be raised from 620 to 650 at rated power. The design target of this study is preliminary set to an average coolant core outlet temperature greater than 500. The aim of this study is to show an example of such core with three-dimensional core calculations. II. Design Features 1. Fuel Rod Design Features Fuel rod design follows that of LWR. The Fuel rod consists of enriched UO 2 pellets and Ni-alloy (Inconel718) claddings. It is a self-contained type for which collapse should be prevented. The fuel rod is internally pressurized and the cladding thickness is determined to avoid failures due to buckling collapse, stress rupture, plastic deformation and creep at both normal operation and abnormal transient * Corresponding author, Tel , Fax yamaji@utnl.jp

2 conditions. The maximum pressure difference between inside and outside of the fuel rod throughout its life is conservatively determined to be 22.80MPa from the past transient analysis of University of Tokyo 4). The fuel rod designed in this study has a diameter 10.2mm with a Ni-alloy (Inconel718) cladding thickness 0.63mm. The maximum design temperature of the cladding is 800 at which temperature the buckling collapse of the cladding becomes limiting. The cladding thickness is designed with 10% thickness margin. The strength against a buckling collapse can be reinforced with ribbed claddings. As for the cladding materials, austenitic stainless steel, ferrite martensitic steel as 12Cr steels, ODS (oxide dispersion strengthened) steels and many others are also the candidates. The choice of materials and fuel design criteria have direct impacts on the performance of the core, such as the coolant outlet temperature and fuel enrichment. Experimental developments of cladding materials are currently proceeding at various organizations. The selection of the cladding will be made from the results. 2. Features To Be Considered In The Core Design There are three important features, which are considered in designing the core. The first feature is that the coolant flow rate is as low as about 1/8 of that in BWR. To maintain a sufficient cooling of the fuel rods at such a low flow rate, coolant flow speed needs to be maximized by minimizing the gap between fuel rods. The fuel rod gap in the design is 1.0mm, assuming grid spacers can be manufactured. The second feature is the large axial density change of the coolant in the core. The coolant density at the inlet is about 0.8g/cc whereas it is less than 0.1g/cc at the outlet. To prevent a large axial reactivity difference, and to maintain sufficient neutron moderations, especially at the top of the core, many square water rods are introduced in the fuel assembly. The moderator flow direction in the water rods is downward. A core with descending flow in water rods was originally proposed to prevent a thermal fatigue of the control rod guide tube at the top of the core 5). The flow scheme in the core is shown in Fig.1. Part of the inlet water is directed to the upper dome and then, descends down the control rod guide tubes and water rods to the bottom dome. It is then, mixed with the rest of the inlet water and flows up the fuel channels to the outlet. In this flow scheme, the RPV temperature can be kept as low as the inlet coolant temperature. The average coolant core outlet temperature can be kept high, since there is no mixing between the hot coolant and the relatively cold moderator in the water rods at the outlet of the core. The axial water density distributions are shown in Fig.2. The axial change in volume averaged water density is smaller than that in BWR. The third feature is that the plant system is once-through direct cycle, and the average coolant outlet temperature needs to be kept as high as possible. To maximize the average coolant core outlet temperature, the radial power peaking factor needs to be kept as low as possible and the coolant flow rate distribution needs to be adjusted to the radial power distribution. It is also necessary to make the outlet temperature of the sub-channels of the fuel assemblies as flat as possible. III. Core Design Method 1. Design Criteria The core is designed with the following criteria: 1. maximum liner heat rate at rated power; 39kw/m (maximum fuel centerline temperature 1930 ) 2. maximum cladding surface temperature 650 for Ni-alloy cladding at rated power 3. negative coolant void reactivity 4. core shut down margin, greater than 1.0% K The maximum cladding surface temperature is conservatively determined for avoiding oxidation corrosions of the claddings during normal operations. In supercritical FPP, Ni-alloy is a candidate material for the super-heater tubes operating at coolant temperatures from 650 to 700. Even austenitic stainless steels have been used as super-heater tubes operating at coolant temperature 610. Considering that the temperatures of the super-heater tubes themselves are higher than the coolant temperature, the present design criterion of 650 is conservative for Ni-alloy. The cladding temperature limit is also determined to prevent fuel cladding failures during abnormal transients. It has been raised from the 620 used in the past design 1), because the recent transient analysis of SCLWR-H shows that there is a margin for the maximum cladding surface temperature criterion at all transient scenarios 6). Furthermore, the designed temperature for the Ni-alloy cladding (Inconel718) is currently limited by the buckling collapse of the cladding at 800. However, the cladding strength against buckling collapse can be reinforced by ribbed claddings. In SCR, there is no such criterion as minimum critical heat flux ratio of LWR. The cladding temperature does not rise sharply during heat deterioration and the deterioration disappears in the down-stream. As a result, the heat deterioration phenomenon during abnormal transients does not lead to a burnout. The supercritical water can be treated as a single-phase flow. These characteristics enable the direct numerical calculations of the heat transfer coefficients to evaluate fuel integrities during abnormal transients 4). It should be noted that the above values are determined for the conceptual development purpose. The maximum cladding temperatures need to be determined from experiments in relation with water chemistry and transient analysis of a designed core. 2. Neutronic Calculation All neutronic calculations are carried out using SRAC code system of JAERI (Japan Atomic Energy Research

3 Institute). The macro cross sections required for a core burnup calculation are prepared for 10 different water densities, depending on the axial positions of the fuels in the core. The water density distribution is obtained assuming a cosine power distribution in the axial direction. Each cross section is prepared through three different burnup calculations as described below: 1) nuclides production/decay calculations for enriched uranium fuel rods 2) nuclides production/decay calculations for a fuel rod with burnable poisons (Gd 2 O 3 ) 3) fuel assembly burnup calculation All production/decay calculations are carried out using 107 energy groups based on JENDL3.3 nuclear data library. The 107-energy group cross section sets are collapsed to 10 (5 fast and 5 thermal) energy groups at the end of the calculations. In the fuel assembly burnup calculation, non-burnable materials, such as the moderator in water rods and the fuel assembly structural materials are treated heterogeneously in a 1/4-symmetrical geometry. The cross sections for each fuel rod is interpolated by burnups, and homogenized 2energy-group (1 fast and 1 thermal) cross section sets are obtained for the core burnup calculation. All of these cross section sets are prepared for two cases. The case with control rods withdrawn, and the other case with control rods inserted into the assembly. The core burnup calculation refers to either of the above two cases at each of 40 axial planes to evaluate control rod positions. It is also an interpolating calculation by the burnups, in the same way that the fuel assembly burnup calculation is. The cross section sets used at this stage is homogenized, so the information regarding the heterogeneous nature of the fuel assembly is not evaluated in the core burnup calculation. In order to design an equilibrium core, the core burnup calculation needs to be repeated until the power distributions and effective multiplication factors are converged for a given control rod pattern scheme and a fuel load scheme. These schemes are designed so that the effective multiplication factor of the core is kept constant at critical throughout a cycle, with reasonable core power distributions. In this study, all evaluations are carried out on the equilibrium core obtained after 23 cycles of core burnup calculations. The equilibrium core refers to the 24 th core burnup calculation unless stated otherwise. The three-dimensional evaluation of the power distributions with a control rod pattern scheme is introduced for the first time in the SCLWR-H core development. The accuracy in the evaluation of the core has been greatly improved from the past study by using more detailed models. The featuring differences are summarized in Fig Thermal Hydraulic Calculation In the core burnup calculation, each fuel assembly is divided into 6 by 6 by 40 meshes in the x, y, and z directions respectively. A point power distribution is obtained for each of these meshes. A single channel thermal hydraulic calculation is applied to the power distributions at BOC, MOC and EOC to evaluate the average coolant outlet temperature. Orifices are attached to the inlet of each fuel assembly to fix the coolant flow rate during a cycle. The flow rate is determined to match the maximum power of the fuel assembly so that the maximum cladding surface temperature and fuel centerline temperature criteria are satisfied. In the single channel thermal hydraulic analysis code (SPROD), the geometry is modeled by a single channel shown in Fig.4. As for the water rods, the heat transfer area and ratio of the number of fuel rods to the number of water rods are considered. In the code, the radial heat transfer and conductance in each of the 40 planes are calculated first. Then, the axial heat transport is calculated. The calculation is repeated until the coolant and moderator temperature distribution is converged. The core pressure (25.0MPa) is at the supercritical pressure of water. The heat transfer rate from the claddings to the coolant is calculated using Koshizuka-Oka-Kitoh s correlation 4). 4. Consideration for a Neutronic and Thermal Hydraulic Coupling The neutronic calculation code (SRAC code system) does not have a function, which couples the neutronics and the thermal hydraulics. The water densities in any single burnup calculation are kept constant. In the SCLWR-H core, the coolant flow rate is adjusted to each fuel assembly power, keeping the power/flow rate ratio constant. Moreover, the neutronics are mostly governed by the neutron moderation by the moderator in water rods, whose density change is relatively small in the core. These features suggest that assumption of constant water density throughout the cycle may not be a bad estimate. However the current calculation method cannot evaluate the density feed back effects due to the change in power distributions during a cycle. Hence, there is a tendency to over estimate the power peaking factors. This may also have an effect on the average coolant outlet temperature. A qualitative evaluation with coupled calculation must be done in the future.. Core Design 1. Fuel Assembly Design The cross section of the designed fuel assembly is shown in Fig.5(a). 300 fuel rods are arranged in a square lattice. The square boxes, regularly arranged inside the assembly, are the water rods in which supercritical water descends down the core for neutron moderations. The 16 central water rods are equipped with control rod guide tubes. The 16 control rods constituting a cluster unit are to be inserted

4 from the top of the core as in PWR. The design is more suitable for achieving a high average coolant outlet temperature than the previous hexagonal fuel assembly design 1) shown in Fig.5(b). Square water rods separate the array of fuel rods. The sub-channel flow is restricted by the water rods. Rectangular water rods are provided at the outer part of the fuel assembly and serve as a channel box restricting the cross flows between fuels and assemblies. The sub-channel areas and neutron moderation are more uniform compared to the previous design. The relative fuel rod power peaking of a fuel assembly was 1.29 in the past design when all fuel rods are fresh 5.22wt% enriched U 235. It is reduced to 1.05 in the present design, when all fuel rods are fresh 6.47wt% enriched U 235. The enrichment splits of fuel rods within the assembly are not necessary. The fuel assembly design is summarized in Table 1. The fuel rod is axially divided into three parts and the U 235 fuel enrichment in each of these three parts is 6.10wt%, 6.60wt% and 6.10wt% respectively. The average enrichment is 6.28wt%. The average fuel discharge burnup is 45GWd/t. Compared to BWR or PWR the fuel enrichment is high, but the fuel cycle cost is low due to the high thermal efficiency. The fuel claddings and all of the fuel assembly structural materials are made from Ni-alloy (Inconel718). The thermal neutron absorption cross section of Ni-alloy is much larger (more than 10 times) than that of Zircaloy, but it cannot be used at such a high temperature and pressure of SCLWR-H. For a burnup reactivity compensation, burnable poisons (Gd 2 O 3 ) are mixed with the fuel as in BWR. The chemical shim as in PWR cannot be used due to the potential negative coolant density coefficient. Among the 300 fuel rods, 10wt% Gd 2 O 3 is mixed in 24 fuel rods. 10wt% is the highest commercial use value in BWR. The burnable poison design is determined to satisfy two design targets. The first target is to minimize the excess reactivity at the beginning of the cycle. The other is that the infinite multiplication factor of the fuel decreases monotonously with the burnup. An infinite multiplication factor change with a burnup for a typical SCLWR-H fuel assembly is shown in Fig.6(a) for a comparison with the schematic feature of a BWR fuel shown in Fig.6(b). The BWR-type design cannot be adopted for the large SCLWR-H fuel assembly (about 4 times larger than the BWR fuel assembly) for the following reasons. Firstly, the radial power peaking factor becomes large due to the large reactivity differences between the first-cycle fuels and the second-cycle fuels. Secondarily, the average coolant outlet temperature becomes low due to the power peak shifts from the second-cycle fuel at the BOC to the first-cycle fuel at the EOC. 2. Fuel Load Pattern The fuel load pattern is similar to that in PWR. The core consists of 120 fuel assemblies of the first to the third cycle fuels and a fourth-cycle fuel assembly at the center of the core. The fuel assembly pitch is 29.62cm and the equivalent diameter of the core is 3.68m. The out-in fuel load pattern for the three-cycle equilibrium core is shown in Fig.7. The first cycle fuels have the highest reactivity and most of them are loaded at the periphery of the core. The rest of the first-cycle fuels are all surrounded by the third cycle-fuels to minimize the power peaks. 3. Reactivity Control The reactivity control is similar to that in BWR. However, at BOC, the excess reactivity to be controlled by control rods (about 4 to 5% K) is relatively higher than that of BWR(about 2% K) and gradually decreases to about 1% K towards EOC. This is because the use of burnable poisons is limited to keep the radial power distribution flat and outlet temperature high, as explained earlier. Evaluations concerning the reactivity control have been carried out to confirm that the excess reactivity of the core is controlled throughout a cycle and the core can be brought to a cold shut down at all time with a sufficient shut down margin. All evaluations are carried out for the equilibrium core. A cluster of control rods is designed. The16 control rods constituting the cluster are to be inserted into the water rods as shown in Fig.5(a). The control rod is made from 70%TD natural boron carbide (B 4 C), and has a diameter 12.4mm. The fresh core has no Xe-poison, and the excess reactivity is about 6.60% K. The Xe-poison reaches equilibrium after about one day of operation (equivalent to about 30MWd/t) and the excess reactivity falls to about 4.66% K. The excess reactivity at EOC is intentionally kept a little higher than zero (about 1% K) so that some control rods are still inserted at EOC. These control rods are needed for the power distribution control as well as the plant control 6). The power distribution control is described in the next section of this paper. The hot/cold reactivity difference of the fresh core (Xe-free) is 5.3% K. The core can be brought to a cold shut down at all time. The cold shut down margin for a fresh core is 4.2% K. It is evaluated assuming all control rod clusters, except for the four control rod clusters with the highest reactivity worth, are inserted into the core. The design criterion of 1% K is satisfied. The highest reactivity worth for one cluster is 0.44% K at hot operating condition, and 0.62% K at cold shut down condition. The reactivity characteristics of the core is summarized in Table Power Distribution Control The power distribution control is a crucial part of the core design, because it has a direct influence on the average coolant outlet temperature. Compared to BWR and PWR, the role of control rods in controlling the power distribution is much more important. Two important points need to be considered.

5 Firstly, a large axial power peak near the top of the core must be avoided so that the fuel centerline temperature does not exceed the design criterion This is a typical point in a high temperature core. As the coolant rises up the core, its temperature rises and the heat transfer coefficient from the fuel rod to the coolant decreases. The fuel centerline temperature cannot be kept below the criterion of 1930 if a large power peak exists at the top of the core. The second point is that the radial power distribution has to be kept not only flat, but also as constant as possible. This is because the coolant flow rate is fixed to the maximum power of the fuel assembly throughout a cycle. The average coolant outlet temperature cannot be kept high, if there is a large radial power swing. This is a difficult target in SCLWR-H core design, but reasonably attainable. In PWR, the reactivity control is mostly done by the chemical shim mixed in the coolant. There is very little burnable poison mixed in the fuels themselves, so that the reactivity differences between each cycle of fuels are kept almost constant throughout a cycle. The radial power distribution is governed by the fuel load pattern and active control of the radial power distribution is not necessary. Almost all control rods are withdrawn when the core is at a rated power. In BWR, large amount of burnable poisons are mixed in fresh fuels. Therefore, the reactivity profiles of the fresh fuels differ from the fuels in the second or third cycle. Hence, the differences in reactivity between the first cycle fuels and the rest of the fuels change with time (burnup). However, the radial power distribution of the core does not vary much throughout a cycle. This is because BWR fuel assemblies are relatively small, and all fresh fuels are surrounded by non-fresh fuels. One fuel assembly is not large enough to characterize the radial power distribution. Hence, the radial power distribution of BWR is also governed by the fuel load pattern. Only a small number of control rods are inserted in the core at a rated power. In SCLWR-H, the fuel assembly is similar in dimensions as that of PWR (about 4 times larger than the BWR fuel assemblies). The fuel assembly is large enough to characterize the radial power distribution, and the fresh fuels contain large amount of burnable poisons. As a result, there would be a large radial power swing with the burnup of burnable poisons in the fresh fuels. It is difficult to keep the radial power distribution flat throughout a cycle, without the presence of an active power distribution control. The radial power distribution is actively controlled by control rods. The design target is to keep the radial power as flat as possible throughout a cycle. The control rod pattern scheme is shown in Fig.8. It is adjusted to keep the core reactivity at critical, and controlling the power distribution within an acceptable range during a cycle operation of the core. The control rod pattern is adjusted at every 1.1GWd/t apart from the first and the last part of the cycle, where the reactivity change of the core is relatively large with respect to the burnup. The rapid reactivity change at the start of operation is due to the effect of Xe-poison accumulations, whereas it is due to the absence of burnable poisons near the last part of the operation cycle. The normalized axial power distributions at BOC, MOC and EOC are shown in Fig.9. It is the relative power distribution of the average power in each of the 40 planes of the core. The control rods are inserted from the top of the core and gradually withdrawn. The power peak at the top of the core at EOC is prevented by the shallow-inserted control rods (see Fig.8). The normalized radial power distributions at BOC, MOC, EOC are shown in Fig.10. The figure shows the normalized relative power of the fuel assemblies in 1/4-symmetrical geometry. The maximum fuel assembly power is 1.27 relative to the core average. 5. Peaking Factors and Core Power The core power depends on the average linear heat generation rate for a given geometry, and the average linear heat generation rate depends on the total power peaking factor of the core. Each of the three peaking factors, radial, axial and local power peaking factors are evaluated throughout a cycle. The definitions of these peaking factors are derived from the BWR core design as below. 1) Radial: ratio of the maximum fuel assembly power to the average fuel assembly power in the core 2) Axial: ratio of the maximum plane power to the average plane power in the maximum power fuel assembly 3) Local: ratio of the maximum fuel rod power to the average fuel rod power at the maximum power plane of the maximum power fuel assembly. The radial power peaking factor does not vary much, and its highest value is Meanwhile, the axial power peaking factor varies from 1.14 to 1.50 depending on the position of the control rod in the maximum power fuel assembly. The local power peaking factor varies from 1.05 to 1.17 depending on the amount of burnable poison contained in the relevant plane of the fuel assembly. The total power peaking factor at any given time is obtained by the product of radial, axial and the local power peaking factors at that time. The total peaking factor, and therefore the maximum linear heat generation rate is largely determined by the axial power peaking factor. Fig.11 shows the burnup profile of the axial power peaking factor and the maximum linear heat generation rate. The total power peaking factor reaches its maximum value of 2.17 at EOC, and this is where the maximum linear heat generation rate reaches the design criterion of 39kW/m. Hence, the average linear heat generation rate is 18.0kW/m, which is about the same as in BWR. The average linear heat generation rate can be raised by first, raising the maximum heat generation rate, and second, by reducing the total power peaking factor.

6 There exists a fuel with maximum linear heat generation rate 60kW/m in BWR, but further analysis and experiments are necessary to determine the allowable maximum linear heat generation rate in SCLWR-H. The maximum value of the total power peaking factor 2.17 is obtained at EOC, when the maximum power fuel assembly is the first-cycle fuel. First-cycle fuels tend to have large local peaking factors (around 1.20), because of the burnable poisons mixed in some of the fuel rods. The total power peaking factor can be reduced to around 2.0 by modifying the control rod patterns, so that the maximum power fuel assembly is not in its first cycle. When maximum linear heat generation rate is 39kW/m, reducing the total power peaking factor from 2.17 to 2.0 would increase the average linear heat generation rate from 18.0kW/m to 19.5kW/m. The core power is 2740MW thermal, and the average core power density is 61.5kW/l when the average linear heat generation rate is 18.0kW/m. The power density is in between of the values in BWR and PWR. 6. Average Coolant Core Outlet Temperature Average coolant core outlet temperature is one of the most important core parameters. The coolant core outlet temperature can be raised to a high temperature by reducing the core flow rate. However, the fuel integrity must be ensured. In order to achieve a high average outlet temperature, the coolant flow rate needs to be adjusted to the radial power distribution. The average coolant core outlet temperature is 496 when the coolant flow rate is adjusted for each fuel assembly. The maximum coolant outlet temperature of each fuel assembly exceeds 580, but the average coolant outlet temperature of the fuel assembly is greatly reduced, especially for the fuels loaded at the periphery of the core. This is because the change in radial power distribution within the fuel assembly is large. The fuel load patterns and control rod pattern scheme need to be optimized to improve the radial power distribution. The optimization of the core design remains for future studies. Meanwhile, several approaches can be thought to raise the average core outlet temperature of the designed core. One approach is to adjust the coolant flow rate more precisely to the power distribution. When the coolant flow rate is adjusted for each quarter of the fuel assembly, the average coolant core outlet temperature is 528 and exceeds the design target of 500. The other approach is to cool the fuels at the periphery of the core by a downward flow. The main cause of the reduction in the average core outlet temperature is the relatively cold outlet coolant from the fuel assemblies at the periphery of the core. So, cooling the peripheral fuels with a downward flow is effective in raising the average core outlet temperature. By cooling the peripheral fuels with a downward flow and adjusting the flow rate for each fuel assembly, the average core outlet temperature can be raised to 565. Raising the maximum cladding temperature criterion from 650 to 700 can raise the average core outlet temperature from 496 to 526 when the coolant flow rate is adjusted for each fuel assembly. High outlet temperature is beneficial for the concept, but the simplicity of the concept must be considered at the same time. The fuel load pattern and the control rod pattern scheme should be optimized before considering other approaches. The above values are obtained based on a single channel thermal hydraulic analysis code. More detailed consideration for the outlet temperature needs to be done with a sub-channel analysis. The thermal efficiency is 44.4% when the average outlet temperature is 528, giving the electrical output of 1220MW.. Core Characteristics Summary and Discussion The core characteristics are summarized in Table 3. The core pressure (25MPa) is at a supercritical pressure of water. The fuel rod has a diameter 10.2mm with a Ni-alloy (Inconel718) cladding of thickness 0.63mm. The thickness of the cladding is determined to withstand a maximum pressure difference of 22.80MPa at 800. The average fuel enrichment is 6.28wt% for a discharge burnup 45GWd/t. The enrichment is relatively high, because of the large neutron absorption by the Ni-alloy. The cladding material for SCR has to be developed with experiments. The average linear heat generation rate is 18.0kW/m when the maximum linear heat generation rate is 39kW/m. The average linear heat generation rate can be increased by reducing the total power peaking factor and raising the maximum linear heat generation rate. However, more analysis and experiments are necessary for determining the allowable maximum linear heat generation rate. The power density is 61.5kW/l, when the average linear heat generation rate is 18.0kW/m, giving the thermal output of the core 2740MW. The average coolant core outlet temperature is 528 for the maximum cladding surface temperature 650 when the coolant flow rate is adjusted to the power for each of the quarter of the fuel assembly. The optimization of the design remains for future studies. The thermal efficiency at the average outlet temperature 528 is 44.4%, giving the electrical output of the core 1220MW.. Conclusions An SCLWR-H core is designed by three-dimensional core calculations. The reactivity control is successfully demonstrated using burnable poisons and control rods. A control rod pattern scheme is designed for controlling the power distribution so that the average coolant core outlet temperature is maximized under the design criteria. The average coolant core outlet temperature is 528 when the coolant flow rate is adjusted for each quarter of the fuel assembly. It is 565

7 when the peripheral fuels are cooled by downward coolant flow and the coolant flow distribution is adjusted for each fuel assembly. The preliminary design target of 500 is achieved. Nomenclature None Acknowledgment The authors thank all members involved in the SCLWR-H development project. In particular, thank Mr. S.Sakurai, TOSHIBA Engineering Co. for his inputs on the SCLWR-H core design. moderated by supercritical light water" Proc. GENES4/ANP2003, Kyoto Research Park, Kyoto, Japan, Sept , ) S. Koshizuka, Y. Oka, et al., "Computational Analysis of Deterioration Phenomena and Thermal Hydraulic Design of SCR" Proc. SCR2000 Symposium, The University of Tokyo, Tokyo, Japan, Nov. 6-9, ) S. Tanaka, et al., "Core Design of Supercriticalpressure Light Water Reactor" Proc. ICONE-4, vol.2, pp , ASME (1996). 6) Y. Ishiwatari, Y. Oka, S. Koshizuka, et al., "Plant control of high temperature reactor cooled and moderated by supercritical light water", Proc. GENES4/ANP2003, Kyoto Research Park, Kyoto, Japan, Sept , 2003 References 1) K. Dobashi, et al., "Conceptual Design of a High Temperature Power Reactor Cooled and Moderated by Supercritical Light Water" Ann. Nucl. Energy, vol.25, pp (1998). 2) S. Sakurai, N. Yoshida, S. Siga, et al., "Development Project of Supercritical-water Cooled Power Reactor Plant Concept (3) " J. Nucl. Sci. Technol., 41[3], 326 (2003). 3) Y. Ishiwatari, Y. Oka, S. Koshizuka, et al., "Safety analysis of high temperature reactor cooled and Fig.1: Water flow scheme in SCLWR-H core Fig.2: Axial water density distribution

8 Fig.3: Neutronic calculation models modifications Fig.4: Single channel thermalhydraulic calculation model (a) New design Fig.5: SCLWR-H fuel Assembly (b) Past design K-inf Burnup (a) Typical SCLWR-H fuel (b) Typical BWR fuel Fig.6: Infinite multiplication change of fuels with respect to burnups Fig.7: Fuel load pattern (1/4 symmetrical geometry) Fig.9: Normalized axial power distribution

9 Fig.8: Control rod pattern scheme Fig.10: Normalized radial power distribution Fig.11: Burnup profile of the axial power peaking factor and maximum linear heat generation rate

10 Table 1: Fuel assembly design Axial design U 235 fuel rod Gd 2 O 3 fuel rod Number of CRs U 235 enrichment (wt%) Number Gd 2 O 3 concentration (wt%) Number per unit 3 splits Upper Middle Bottom Table 2: Core reactivity characteristics Excess reactivity (% K) Shutdown margin (% K) Maximum CR cluster reactivity worth (% K) BOC EOC BOC Hot operating Cold shutdown condition condition Table 3: Core characteristics Core SCLWR-H BWR PWR Pressure [atm] Fuel rod diameter / Cladding Thickness [mm] 10.2/0.63 (Ni-alloy) 12.0/0.9 (Zircaloy) 9.5/064 (Zircaloy) U 235 fuel enrichment (top/mid/bot/ave) 3.0(average) 3.5(average) [wt%] 6.1/6.6/6.1/6.28 Number of 121/ /62 193/264 FAs/Number of fuel rods per FA Average discharge burnup [GWd/t] Number of CRs 121(Cluster-type) 185(cross-shaped) 53(Cluster-type) Total Peaking (axial/radial/local) (1.46/1.27/1.17) (1.40/1.40/1.24) Inlet / Outlet 280/ / /325 temperature [ ] Thermal efficiency [%] Thermal / Electrical 2740/ / /1180 power [MW] Active height / 4.20/ / /3.4 Equivalent diameter [m] Max / Average linear / /17.9 heat generation rate [kw/m] Average power density [kw/l] Coolant flow rate [kg/sec]