Fuel Management Effects on Inherent Safety of Modular High Temperature Reactor

Transcription

1 Journal of NUCLEAR SCIENCE and TECHNOLOGY, 26[7], pp. 647~654 (July 1989). 647 Fuel Management Effects on Inherent Safety of Modular High Temperature Reactor Yukinori HIROSEt, Peng Hong LIEM, Eiichi SUETOMI, Tohru OBARA and Hiroshi SEKIMOTO Research Laboratory for Nuclear Reactors, Tokyo Institute of Technology* Received January 13, 1989 Analysis was performed on the effects of fuel loading schemes and fuel materials on the inherent (passive) safety characteristic of the modular pebble-bed type high-temperature reactor against depressurization accident involving loss of He forced circulation Two extreme fuel loading schemes, the infinite-velocity multipass and Once-Through-Then-Out (OTTO), were evaluated for both U and Th fuel cycles. The results of the analysis show that the maximum core temperatures attained following the accident were much lower for the infinite-velocity multipass scheme than for the OTTO scheme. For both schemes, the Th cycle showed slightly higher maximum peak temperatures, compared to U cycle. KEYWORDS: fuel management, modular high-temperature reactor, reactor safety, inherent safety, depressurization accident, loss of helium forced circulation, reactor fueling, OTTO, infinite-velocity multipass, fuel cycle, equilibrium condition, core temperature I. INTRODUCTION As an inherently safe reactor, the modular high-temperature reactor (HTR) was proposed through development of the pebble-bed reactor. The inherent or passive safety for the HTR can be stated as the capability of the reactor to limit the maximum fuel temperature to the values for which the fission product release will increase the fission product inventory in the primary circuit by factors, but not by order of magnitude(1). The well known TRISO particles with silicon-carbide (SiC) layer comply with this requirement for temperatures up to 1,600dc. In other words, providing the modular HTR design with passive safety property means limiting the maximum fuel temperature below 1,600dc during the accidents. One essential determinant of maximum fuel temperatures for a pebble-bed reactor is the in-core fuel management. This involves numerous parameters such as choice of fuel loading scheme, fuel material etc. In the present paper, the above two parameters were considered. Concerning the first, two extreme fuel loading schemes, infinitevelocity multipass and Once-Through-Then- Out (OTTO), were considered. For the fuel material parameter, low enriched U fuel and Th with 233U fuel, were compared. II. FUEL LOADING SCHEMES AND ACCIDENT CONDITIONS In pebble-bed reactors, fuel balls are loaded from the top of the reactor, flow through the core and are discharged from the bottom. The fuel ball loading schemes commonly used for the pebble-bed reactor may be classified into multipass and OTTO schemes. For the multipass scheme, the fuel balls flow through the reactor numerous times before being * O-okayama, Meguro-ku, Tokyo 152. Present address: Nippon Atomic t Industry Group Co., Ltd., Ukishima-cho, Kawasaki-ku, Kawasaki

2 J. Nucl. Sci. Technol., finally discharged. But for the OTTO scheme, the fuel balls transit the core only once. The OTTO scheme may attain a nearly exponential axial power profile with the aim of fuel temperature flattening. For both multipass and OTTO schemes, the ball flow velocities affect the nuclide density distributions, neutron flux, burn-up, and in turn the power and temperature profiles. With slower ball flow velocities, the power peak shifts to the upper part of the core. This effect is greater for the OTTO scheme. If the velocities become infinite for the multipass scheme, the power peak becomes broader and occurs at the center. For the usual multipass scheme, the ball velocity is slow and the power peak appears in the upper part of the core showing the characteristics between the OTTO and infinite-velocity multipass schemes. From the above, investigations were carried out on the OTTO and infinite-velocity multipass fuel loading schemes. The inherent safety characteristics were evaluated for each scheme during the depressurization accident involving loss of He forced circulation after the equilibrium condition. Furthermore, two kinds of fuel material-low enriched UO2 and Th02 with 233UO2-were also investigated for each fuel loading scheme. It should be noted that for the infinitevelocity multipass loading scheme, the ball flow velocities are large enough that they can be assumed to be infinite. For this loading scheme the nuclide densities can be treated to be spatially uniform. The accuracy of this assumption was investigated at the start of the present study. For the OTTO loading scheme, all nuclides are treated space-dependently, since the ball flow velocities are relatively slow. The details of the OTTO cycle burn-up simulation are presented in Ref. (2). After the accident, the reactor becomes subcritical by its negative temperature coefficient reactivity feedback, but the cooling power of the He goes down through the depressurization of the He coolant. Decay heat is removed only by conduction inside and between fuel balls, radiative cooling, natural convection of He, and conduction through the graphite reflector into surrounding sections. III. CALCULATIONAL PROCEDURE Based on the conditions stated in the preceding chapter, the following calculations were performed for both OTTO and infinitevelocity multipass schemes. 1. Equilibrium Condition of Core Search of the equilibrium condition of the core (density distribution for each nuclide, neutron flux distribution for each energy group, and power distribution) is necessary to obtain information about the initial conditions prior to the accident. Considering the fuel ball motion through the core, the simulated burn-up equation for the OTTO cycle can be written as follows(2): (1) where Ni(s): Atomic density of isotope i at position s s: Distance measured along fuel ball stream line v: Ball speed i: Decay constant l of isotope i i,a,g: Absorption cross section of s isotope i for energy group g : Probability that decay of isotope i' produces isotope i bi'->i,g : Probability that neutron absorption in isotope i' produces isotope i : Yield of isotope i due to fission in isotope i' g(s); Neutron p flux in energy group g at position s. For the infinite-velocity multipass scheme, the first term of Eq. (1) is removed, and Ni(s) becomes independent of s. The neutron transport problem was treated as an g-z two-dimensional four group diffusion problem. The group constants and their selfshielding factors as a function of temperature and atomic density were prepared using a part of VSOP code system(3) (ZUT-DGL, THERMOS ai'->i gi'->i,g

3 Vol. 26, No. 7 (July 1989) 649 and GAM). 2. Steady State Temperature Distribution After the nuclide densities and neutron flux distributions for the equilibrium cycle were calculated, thermal-hydraulic calculation was performed to obtain the temperature distribution in the core for this equilibrium condition, which was used as an initial condition in the accident analysis. In the normal condition, He coolant enters the core from the upper part of the core, flows through the voids among fuel balls, and exits from the bottom part of the core. Considering the spherical shape of the fuel and the void fraction of the core, an effective coolant flow was modelled. In this model, within the core the coolant flows through the virtual flow channels without cross-flow, and in a particular virtual flow channel the mass flow is assumed to be constant. For the momentum equation, Ergun equation(4), which is based on Blake-Kozeny equation for laminar flow and Burke-Plummer equation for turbulent flow, was used (2) where P: Coolant pressure G: Mass flow rate for virtual channel : Coolant mass density m r : Coolant viscosity g: Gravitational acceleration : Void fraction of ecore Dp: Fuel ball diameter. In the core the heat transfer processes involved are (1) forced convective heat transfer between fuel balls and He coolant, (2) radiative heat transfer between the adjacent fuel balls, (3) heat conductions inside the fuel balls and (4) heat convections in the He coolant. Considering these heat transfer processes the overall effective heat transfer coefficient between He coolant and fuel balls was obtained and used for the energy balance equation. The energy balance equation in the fuel can be expressed as (3) where ke: Effective conductivity inside fuel ball h: Forced convective heat transfer coefficient between fuel balls and He coolant Q: Power density Tb: He bulk temperature Ts: Fuel ball surface temperature. For the coolant, the energy balance equation can be expressed similarly as where kb,r: Coolant effective conductivity in radial direction Coolant effective kb,z: conductivity in axial direction. 3. Temperature Distributions and Peaks during Accident (4) Using the equilibrium condition as the initial condition, the fission products and other radiative nuclides distributions, and the corresponding decay heat distribution after a depressurization accident involving loss of He forced circulation were calculated. For the fission-products yield calculation, JNDC FP Decay and Yield Data were used(5). After the depressurization accident, the He flow velocity was assumed to be practically zero and He pressure was also assumed to be equal to the atmospheric pressure. The temperature for the mixture of He coolant and fuel balls was calculated from the He temperature and the surface temperature of the with balls. where Tmx: Mixture temperature p: Mass density r of ball rb : Mass density of He Cpp: Specific heat for constant pressure of ball Cpb : Specific heat for constant pressure of He (5) -3-

4 J. Nucl. Sci. Technol., Tpa: Average temperature of ball Tpb: Bulk temperature of He k: Heat conductivity of ball. Finally, the heat transfer equation to be solved was expressed as the following: The reactor configuration used in the calculation is shown in Fig. 1. The main design parameters, as well as the respective calculation results for the equilibrium condition for each fuel loading scheme and fuel material, are shown in Table 1. The core diameter is much smaller than the usual HTR designs in order to enable both reactor-shutdown by only reflector rods and decay-heat removal without any active cooling system. For both fuel loading schemes, the same fissile enrichments and moderation ratios were used. The burnup and conversion ratios for the infinitevelocity multipass scheme are relatively higher than for the OTTO scheme for both U and Th cycles. The Th fuel cycle gave slightly higher values of burnup and conversion ratio. (6) where k'e: Effective conductivity of core during accident q: Decay heat production during accident IV. CALCULATED RESULTS AND DISCUSSION Fig. 1 Reactor configuration Table 1 Reactor main design parameters and calculated result for equilibrium condition Before comparing the two extreme fuel loading schemes, we discuss the effects of the spatial distribution of the nuclide densities resulted from the finiteness of fuel ball velocity for the multipass scheme. Even for finitevelocity multipass scheme, the densities of large half-life nuclides do not depend strongly on the spatial distribution of neutron flux, when the ball velocity is high enough. For short half-life nuclides, however, the dependency on the spatial distribution of the neutron flux increases and these nuclide densities actually exhibit spatial distributions. Since these spatial distributions are essential in the present analysis, they were investigated for some important short half-life nuclides 135Xe, 149Sm, Np and 233Pa, and compared with 239 the results for the infinite-velocity multipass scheme. Discrepancies of not more than 1, 3 and 1% respectively were detected in group neutron fluxes, densities of short half-life nuclides, and burn-up values. For the accident analysis concerned, these will not contribute significantly to the final results.

5 -5- Vol. 26, No. 7 (July 1989) 651 The power distributions for each fuel loading scheme and fuel cycle are shown in Fig. 2(a), (b) and Fig. 3(a), (b). The power distributions of the OTTO scheme showed peaks on the upper part of the core. In the OTTO scheme, since the fuel balls transit the core only once, the upper part of the core contains more fresh fuel and the lower part contains highly burned-up fuel. For this fuel distribution, the fission reactions mostly occur in the upper part of the core, and determine the power density profiles with a sharp peak in this part. These effects appear to an extreme degree in the present design, since the core radius is small and axial coupling of the neutron flux distribution is weak. The power distribution of the infinite-velocity multipass scheme becomes broader and its peak locates close to the center. Using the neutronic calculation results for the equilibrium condition, the thermal hydraulic calculation were done. These results are shown in Table 2. All of the maximum temperatures of He and fuel ball occur at the center bottom of the core. Under these maximum temperatures, the reactor can be operated safely concerning the fission product release. The maximum He temperatures are higher for the Th cycle than for the U cycle. It is attributed to the larger outlet temperature Fig. 2(a),(b) Power density in equilibrium condition for OTTO U and Th fuel cycles

6 J. Nucl. Sci. Technol., Fig. 3(a),(b) Power density in equilibrium condition for infinitevelocity multipass U and Th fuel cycles Table 2 Thermal hydraulic calculated result for equilibrium condition mismatch among coolant channels for the Th cycle caused by the steeper radial power profile. The accident analysis was begun with the calculation of decay heat production of fission products, actinides and other radiative nuclides subsequent to the accident. The results for the OTTO scheme, Fig. 4(a), (b), gave distributions of the decay heat production in similar

7 -7- Vol. 26, No. 7 (July 1989) 653 Fig. 4(a),(b) Heat production after reactor shutdown for OTTO U and Th fuel cycles profiles, with the power distributions in the equilibrium condition. For the infinite-velocity multipass scheme, since the nuclide densities were uniform, the calculated results of timedependent decay heat production shown in Fig. 5 are considered valid through all core regions. Fig. 5 Heat production after reactor shutdown for infinite-velocity multipass U and Th fuel cycles Using the temperature distributions in the equilibrium condition as initial conditions, and the time dependent decay heat production of fission product after the accident, the axial temperature profiles were calculated as shown in Fig. 6(a), (b) and Fig. 7(a), (b). In these figures, the axial temperature profiles in the equilibrium condition, which served as initial conditions, are also shown in dashed lines. These initial temperature profiles are more flat for the OTTO scheme than the infinite-velocity multipass scheme as we expected. The maximum temperatures attained for each loading scheme and fuel cycle are summarized in Table 3. At the beginning of the accident, the temperature peaks for the OTTO scheme occurred in the upper part of the core. On the other hand, the peaks for the infinite-velocity multipass scheme occurred at the lower part of the core. Both peaks moved toward the center of the core with the passage of time. Fig. 6(a),(b) Axial temperature profiles after accident for OTTO U and Th fuel cycles

8 J. Nucl. Sci. Technol Fig. 7(a),(b) Axial temperature profiles after accident for infinitevelocity multipass U and Th fuel cycles Table 3 Calculated result of depressurization accident analysis The maximum temperature of the OTTO scheme exceeded 1,800dc in 23.5 h for U and 22 h for Th cycles respectively. The maximum temperatures of the infinite-velocity multipass scheme were much lower than the 1,600dc, further, their temperature transients were much slower. The differences in the maximum peak temperatures between the two schemes exceeded 500dc; considerably large from the reactor design point of view. Clearly, the distributions of the fission product as well as the decay heat production for the two fuel loading schemes, can explain the above results. The distributions of the fission product and actinide for the infinite-velocity multipass scheme do not produce any peak, so the heat production density is distributed more uniformly throughout the core. However, in the usual multipass fuel loading scheme, the fuel ball velocity is not infinite, and if velocity is very slow the fission products' distribution becomes similar to the distribution for the OTTO fuel loading scheme; the peaks of the distribution will then shift to the upper part of the core. Slight differences in the maximum temperatures were found between U and Th fuel cycles. They were attributed to the different radial power profile in the equilibrium condition which determined the decay heat productions. In conclusion, the multipass fuel loading scheme is relatively safer than the OTTO scheme for the depressurization accident, though the OTTO scheme aims to obtain the higher gas outlet temperature with lower maximum fuel temperature at operation condition. By use of the infinite-velocity multipass loading scheme, a reactor can avoid the 1,600dc by a safety margin of about 300dc. ACKNOWLEDGMENT The authors are very grateful to Dr. M. Aritomi of Tokyo Institute of Technology and Dr. T. Watanabe of Kawasaki Heavy Industries, Ltd., for their advice in performing the thermal analyses. -REFERENCES- (1) REUTLER, H., et al.: Nucl. Technol., 62, 22 (1983). (2) SEKIMOTO, H., et al.: J. Nucl Sci. Technol., 24[10], 765 (1987). (3) TEUCHERT, E., et al.: Jul-1649, (1980). (4) BIRD, R. B., et al.: "Transport Phenomena", 196 (1960), John Wiley & Sons. (5) IHARA, H., et al.: JAERI-M 9715, (1981).