Design and Safety Aspect of Lead and Lead-Bismuth Cooled Long-Life Small Safe Fast Reactors for Various Core Configurations

Transcription

1 Journal of NUCLEAR SCIENCE and TECHNOLOGY, 32[9], pp (September 1995). Design and Safety Aspect of Lead and Lead-Bismuth Cooled Long-Life Small Safe Fast Reactors for Various Core Configurations Zaki SU'UD õ and Hiroshi SEKIMOTO Research Laboratory for Nuclear Reactors, Tokyo Institute of Technology* (Received October 25, 1994), (Revised March 20, 1995) Design and safety aspects of long-life small safe fast reactors using liquid lead or leadbismuth coolant with metallic or nitride fuel are discussed. Neutronic analyses are performed to investigate the effect of core height to diameter ratio (H/D) on design performance of the proposed reactors. All reactors are subjected to the constraint of 12 years operation without refueling and shuffling with constant 150MWt reactor power and also to the requirement of maximum excess reactivity during burnup to be less than 0.1%ƒ k. The results show that the pancake design with H/D of -2/3 gives the most negative coolant void coefficient under the requirements for excess reactivity. Modified designs with the central region axially fulfilled with fertile material are proposed to improve the coolant void coefficient. Thermal-hydraulic analysis results show the possibility to operate the reactors up to the end of life without changing their orifice pattern, necessary pumping power for the proposed design smaller than the conventional large sodium cooled FBR, and the natural circulation contribution of 25-40% at the normal operating condition. The reactivity feedback coefficients are also estimated and appeared to be negative for all the components including the coolant density coefficient. KEY WORDS: liquid metals, lead, lead-bismuth, metallic fuels, nitride fuel, excess reactivity, coolant void coefficient, balance core, pancake core, tall-slender core, height to diameter ratio, long-life fast reactors, reactor safety, pressure drop, driving head, coolant outlet temperature, reactivity feedback I. INTRODUCTION A series of studies on long-life small safe fast reactors using liquid lead or lead-bismuth as coolant and employing metallic or nitride fuel are being conducted(1)-(3). As mentioned in the previous study(2)(3), the goal is to obtain some designs of small reactors suitable for relatively isolated area, for which we set the following basic requirements; long time (10 years or more) operability without refueling or fuel shuffling, compactness, and inherent safety characteristics. In this paper the term inherently safe means that the reactor can survive UTOP (unprotected rod run out transient over power), ULOF (unprotected loss of flow), and other severe accidents without reactor scram or help of operator, and without causing coolant boiling, cladding melting, or fuel pellet melting. Since liquid metal cooled fast reactors generally have small reactivity coefficients, it is necessary to minimize excess reactivity during burnup to below ƒàeff to survive the reactivity accident caused by complete withdrawal of control rods. Enhancing natural circulation is also necessary to survive the loss of flow accident caused by the total failure of all primary pumps. The longer lifetime of the proposed design is basically desirable. However, the lifetime of 12 years is chosen from the consideration of material * O-okayama, Meguro-ku, Tokyo 152. Present address: Dept. of Phys., õ Bandung Inst. of Technol., Jl. Ganesha 10, Bandung, INDONESIA. 8

2 Vol. 32, No. 9 (Sep. 1995) 835 limitation especially relating to corrosion by the coolant. To anticipate the TOP accident, which is initiated by the simultaneous withdrawal of all control rods, and combinations of this accident with other type of accident such as LOF and LOHS accidents, in the present study the excess reactivity is chosen to be less than -0.1%ƒ k or about 22.5 cents of reactivity over the whole lifetime of 12 years without refueling or fuel shuffling, and the coolant void coefficient is required to be negative over the whole life-time(3). In the previous studies(3), it has been confirmed that the above requirements can be achieved by using the core referred as balance design shown in Fig. 1(a), with the axial configuration approximately equal to the radial one using 239Pu as the main fissile material. However in this design a large peaking factor of axially integrated power density in the central axial region occurs and may cause some problems in the thermal-hydraulic design. Fig. 1 Overview of original and modified cores In the present study, we investigate the effects of height to diameter ratio (H/D) on the design performance and we also use a plutonium isotopic composition for "equilibrium cycle"*. We also propose an alternative core design by charging the blanket material in the central axial region (see Fig. 1(b)) so that axially integrated power density for the central region becomes smaller than the other region. Coolant volume fraction in this region is then adjusted to fit the power production in this region at the end of life (EOL). With these methods we can expect a design with both better coolant void coefficient and smaller power peaking factor at the EOL. The results of thermal-hydraulic calculations for various types of design are given. And finally, four reactivity compo- * The term "equilibrium cycle" is related to the condition in which dn/dt=0 for each produced actinide in the core. In this cycle natural uranium is supplied to the reactor. All of the actinides are recycled into the reactor, and the end products of the heavy-isotope decay series (lead and bismuth) and the fission products are taken out from the reactor. The detail description of this cycle and analytical methods for studying this system are given in Ref.(4). The isotopic composition of plutonium in the discharged fuel can be estimated to be equal to the composition in the reactor. This plutonium isotopic composition is shown in Table 1. 9

3 836 J. Nucl. Sci. Technol., nents: Doppler coefficient, coolant density coefficient, core radial expansion coefficient and fuel axial expansion coefficient are calculated and discussed in order to give general view of safety performance. Table 1 Reactor design main parameters and requirement II. CALCULATION MODEL AND METHODS Figure 1(a) shows the configuration of long-life small safe reactors proposed in the previous study(3), and Fig. 1(b) shows a modified configuration proposed in the present study. Design parameters are given in Tables 1 and 2. The main idea in the proposed design is as follows. The internal region usually contributes positively to the coolant void coefficient, therefore minimizing the content of coolant in this region can expectedly reduce the coolant void coefficient. In the previous design the internal region filled with fertile material is surrounded by the core. At the EOL, the axially integrated power density has in general its peak at the center for the previous design though at the Table 2 Sample design main parameters Note: 1. Coolant: Pb: lead, Pb-Bi: lead (44.5%)-bismuth (55.5%) 2. Fuel: Mt: metal, N: nitride 3. Core type: Org: original design, Mod: modified design 4. Pu-composition: Eq: from equilibrium composition, Pu9: only 239Pu present 5. Axial reflector width: 6.0cm, radial reflector width: 10cm 6. ƒ RI, ƒ ZI, ƒ Ro, ƒ Zo are shown in Fig. 1 10

4 Vol. 32, No. 9 (Sep. 1995) 837 BOL the peak position is located in the inner core (see Fig. 1(a)). Therefore from the thermal-hydraulic point of view it is impossible to reduce coolant volume fraction in this region. In the present design the central axial region is completely fulfilled by fertile material as shown in Fig. 1(b) so that the coolant volume fraction in this region can be reduced to 60-75% of the other part of the core. Along burnup, though the plutonium accumulates in the central region, the peak of axially integrated power density does not reach the central part up to the EOL. The original core to be discussed in the present paper is slightly different from the one in the previous study(3). In the previous study the coolant volume fraction in the central axial region is taken to be 30%, and in the present study it is taken to be 31-32% in order to reduce the large peaking factor in the central axial region at the EOL. For the modified design the coolant volume fraction in this region is taken to be 20-27% depending on thermal-hydraulic consideration. The coolant volume fraction for other regions is generally taken to be 35% for both original and modified designs. In this paper the plutonium isotopic composition from equilibrium cycle (fuel type Eq) is generally used, but plutonium containing only 239Pu (fuel type Pu9) is also studied. But we have confirmed that the composition for type Eq is nearly similar to the plutonium composition from the spent fuel of the respective core. The diffusion-burnup calculation is performed for a two-dimensional r-z geometry. The burnup calculation includes 29 heavy metal nuclides from 234U to 248Cm. The detail calculation method is described in Refs. (5) and (6). The cell-average group cross sections are obtained by cell calculations using SLAROM code with JFS-3-J3R cross section The thermal-hydraulic calculation method is described in Ref. (6) in detail, and in this paper only a general description is given. The model for steady state thermal-hydraulic calculation is shown in Fig. 2(a) and the flow diagram is shown in Fig. 2(b). Two-dimensional r-z geometry diffusion calculation is performed to get the power distribution. Then for each value of total coolant flow-rate across Fig. 2 Model for steady state thermal-hydraulic calculation 11

5 838 J. Nucl. Sci. Technol., the core the following calculation is performed. First coolant temperature distribution and flow-rate distribution across the core are calculated by an iteration method until pressure drop in each channel becomes equal for all coolant channel in the core. Then thermalhydraulic calculation across the steam generator (SG) is performed. The average coolant outlet temperature across the core is considered to be equal to the hot pool temperature and to the inlet temperature for SG primary side calculation. Thermal-hydraulic calculation of SG can be performed using both analytical and numerical methods. Some iterations are employed to converge the primary side coolant temperature, secondary side coolant temperature and wall temperature. Then the primary side outlet temperature becomes equal to the cool pool and core inlet temperatures. The core and SG calculations are repeated until both hot pool and cool pool temperatures converge. The temperature across the fuel pin is calculated after the coolant temperature distribution across the core is obtained. The coolant void coefficient is calculated by assuming that all coolant in the core is voided but coolant in any other regions (reflector and shielding) is in the normal condition. Doppler and coolant density coefficients are calculated using a perturbation theory, while core radial expansion and fuel axial expansion coefficients are calculated using direct multigroup diffusion calculations. These calculation are performed after the temperature profile in the core is obtained by steady state thermal-hydraulic calculations. CALCULATION RESULTS 1. Neutronics Figures 3(a) and (b) show the effective neutron multiplication factor (keff) and coolant void coefficient change along burnup for the previous design (Fig. 1(a)) with fuel type Eq. And Figs. 4(a) and (b) show the results for the modified design (Fig. 1(b)). It is shown that all of these designs satisfy the excess reactivity criteria, i.e. _??_0.1%ƒ k, and have negative coolant void coefficients over the whole lifetime of 12 years. The modified design gives the better coolant void coefficient. Fig. 3 keff and coolant void coefficient change during burnup for original design Fig. 4 keff and coolant void coefficient change during burnup for modified design 12

6 Vol. 32, No. 9 (Sep. 1995) 839 The average and peak burnup level for each design at the EOL are shown in Table 2. The modified design shows a lower average burnup than the original design, while the peak burnup is almost same. The modified design contains a larger amount of fertile material. The average burnup is calculated for all fissile and fertile material in the core. Figure 5(a) shows the neutron spectrum of case N (Table 2) during the normal and voiding conditions for the core and blanket regions at the BOL. The neutron spectrum shift caused by the coolant voiding is small for lead and lead-bismuth cooled reactors. The relatively important shift occurs in the high energy part as shown in Fig. 5(b) attributed to the loss of inelastic scattering with the coolant. Fig. 5 Neutron spectra in core and inner blanket during normal and voiding conditions at BOL for case N Figure 6 shows the components of reactivity change caused by 5% coolant voiding for case N as the function of burnup calculated using the perturbation theory. It is shown that the outer core gives a larger leakage reactivity contribution compared to the inner core and internal region. Note that in this case the thickness of the outer core is smaller than the inner core. The spectral shift component for the outer core and internal region is small compared to the inner core. The leakage component for the outer core decreases with burnup since the neutron distribution shifts toward the center of core. This neutron spatial distribution change also tends to reduce the spectral shift component in the outer core and increase in the inner core. The accumulation of the fission product may also be considered as a reason for increasing spectral shift component in the inner core. Figures 7(a) and (b) show keff and coolant void coefficient, respectively, along burnup for cases N and Q. Case Q is similar to case Fig. 6 Components of reactivity change for 5% coolant voiding for case N 13

7 840 J. Nucl. Sci. Technol., Fig. 7 keff and coolant void coefficient change during burnup for cases N and Q N except the fuel type. These figures show that both designs satisfy keff and coolant void coefficient design criteria, and their coolant void coefficient values are approximately the same. For all cases given in Table 2, the excess reactivity constraint is satisfied. The coolant void coefficient for lead-bismuth cooled nitride fueled reactors with fuel type Eq for both original and modified designs and for different core shapes: balance (H/D=1/1), pancake (H/D=-1/2, -2/3) and tall-slender (H/D= 3/2, -2/1), are shown in Figs. 8(a) through - (d). For the original design, decreasing H/D from 1/1 to 2/3 improves the coolant void coefficient though further decrease to 1/2 worsens it (Fig. 8(a)), and increasing H/D from 1/1 to 3/2 also improves it though further increase to 2/1 worsens it in certain areas of burnup time (Fig. 8(b)). For the modified design, decreasing H/D from 1/1 to 2/3 improves the coolant void coefficient although further decrease to 1/2 worsens it (Fig. 8(c)), but increasing H/D from 1/1 to 3/2 and 2/1 worsens it (Fig. 8(d)). Usually both pancake and tall-slender designs are supporsed to give us better coolant void coefficient, but in the present study the con- Fig. 8 Coolant void coefficient change during burnup for cases D, I and J, cases J, K and L, cases H, M and N, and cases N, O and P 14

8 Vol. 32, No. 9 (Sep. 1995) 841 straint on the excess reactivity makes these reactor cores larger and worsens the void coefficient for extreme designs. As a result, the modified design of pancake core with H/D about 2/3 gives the best results of coolant void coefficient. The calculated results of average and maximum burnup are shown in Table 2. The average burnup defined in this paper considers blanket fuel (in region C1) also, although the conventional definition does not consider it. Then the present definition gives smaller value of the average burnup than the conventional definition. The average burnup is proportional to the average power density, and inversely proportional to the core size, since the total power and reactor life are fixed. The balance core shows the largest average burnup for both original and modified designs, since it gives the smallest core size. The difference of coolant void coefficient between the original and modified cores is generally attributed to the following aspects. The first aspect is that the modified core has slightly larger size than the original one due to replacement of fissile material by fertile material at the BOL in the upper and lower part of the central axial region. This aspect in general worsens the coolant void coefficient of the modified core compared to the original core. The second aspect is that the coolant volume fraction in the central axial region can be set much lower than that of the other regions in the modified core compared to that in the original core due to thermal-hydraulic consideration. This factor causes significant improvement of the coolant void coefficient in the modified core compared to the original core. In general, for the balance and pancake cores the modified designs give better coolant void coefficient than the original designs. And for the tall-slender shape the original design gives better results. For the tall-slender core with modified design, significant reduction of coolant volume fraction in the central region can not be performed due to faster conversion of the central axial region from blanket to active core. 2. Thermal-hydraulics In order to investigate the feasibility of the present design from the thermal-hydraulic point of view the results of steady state thermal-hydraulic analysis are given for mainly some of lead-bismuth cooled modified designs. And some of the results for lead cooled reactors are also given. In general the height of chimney is set to be 2m and in other cases it will be mentioned explicitly. Figures 9(a) and (b) show the distribution of axially integrated power density at the BOL and EOL for cases H, N and O. There is a large pattern change from BOL to EOL especially at the central axial part of the core. However we fix the orifice pattern during burnup in order to keep simplicity of operation. The possibility of this concept will be discussed in the following part. Figures 10(a) and (b) show steady state calculated results for the modified design, cases F and H, at the BOL. For each case coolant average outlet temperature, coolant peak outlet temperature, coolant inlet tem- Fig. 9 Axially integrated power distribution of cases H, N and O at BOL and EOL 15

9 842 J. Nucl. Sci. Technol., Fig. 10 Pressure drop, driving head and coolant inlet-outlet temperature for various total flow-rate values of case F (Pb coolant, Mod, H/D=2/3), and case H (Pb-Bi coolant, Mod, H/D=2/3) at BOL perature, total pressure drop across the core, and driving head are shown as functions of the total coolant flow-rate. The larger total coolant flow-rate makes the outlet temperature lower, but the pressure drop across the core also increases. The difference between the total pressure drop and driving head gives an estimate for the necessary pumping power. Note that this pressure drop consists of ones across rod bundles, orifice block, inlet and outlet assembly plenums. In the present study, for lead-bismuth cooled reactor, the total flow-rate is chosen to be 5,500kg/s, while for lead cooled reactor it is chosen to be 7,000kg/s. As shown in the above figures, the necessary pumping power is relatively small compared to that of conventional large sodium cooled FBR which is usually about 0.3MPa or more. This is caused by the low coolant flow-rate across rod bundle, less than 1m/s, and also relatively large pin pitch value. And relatively smaller pumping power is obtained for lead-bismuth cooled reactors than that for lead cooled reactors in the present designs due to the smaller total coolant flow-rate. Design using metallic fuel and nitride fuel using the same type of coolant gave similar results. For comparison, the results for original design of case D are given and shown in Fig. 11. These results show that from thermal-hydraulic point of view, in case of the pancake designs with Fig. 11 Pressure drop, driving head and coolant inlet-outlet temperature for various total flow-rate values of case D (Org, H/D=2/3) at BOL 16

10 Vol. 32, No. 9 (Sep. 1995) 8 43 H/D of -2/3, comparable results are obtained for the original and modified designs. The results at the EOL for case H are shown in Fig. 12. For pancake designs the change of the power density distribution during burnup does not cause any problems in the thermal-hydraulic aspect though the orifice pattern is not changed during burnup. Figures 13(a) and (b) show the results for balance and tall-slender cores using the modified design at the BOL. These results show that the increase of H/D significantly increase the necessary pumping power though it is still below that for the large conventional sodium cooled FBR. Finally the effect of chimney and the level of natural circulation in the present design are shown in Fig. 14. In this figure power to total flow-rate ratio is kept constant by adjusting the reactor power value for each Fig. 12 Pressure drop, driving head and coolant inlet-outlet temperature for various total flow-rate values of case H (Mod, H/D=2/3) at EOL point of total coolant flow-rate value. The cross point of total core pressure drop and driving head shows the natural circulation limit. Figure 14 shows the longer chimney enhances the natural circulation. In general for the cases A through H the natural circulation level for the length of chimney of 2m are 25-40% of the total coolant flow-rates. The use of chimney also slightly reduces peaking factor. Fig. 13 Pressure drop, driving head and coolant inlet-outlet temperature for various total flow-rate values of case N (Mod, H/D=1/1) and case O (Mod, H/D=3/2) at BOL 17

11 844 J. Nucl. Sci. Technol., lead-bismuth cooled metallic fueled core its contribution is comparable to the Doppler effect, and it becomes more important in the ULOF accident. N. CONCLUSION The long-life small safe fast reactors using lead or lead-bismuth as coolant and using metallic or nitride fuels presented in this paper have some important characteristics: unnecessary refueling or fuel shuffling during the lifetime of 12yr, small maximum excess reactivity (0.1%ƒ k), negative coolant void coefficient during whole burnup period, and Fig. 14 Pressure drop, driving head and coolant inlet-outlet temperature for various total flow-rate values and various chimney values of case H (Mod, H/D=2/3) at BOL 3. Reactivity Coefficient Doppler, coolant density, core radial expansion and fuel axial expansion coefficients of the modified pancake design with H/D of 2/3 (case E to H) are shown in Table -3. All of these reactivity feedbacks are. negative. It will be very useful for achieving the inherent safety capability toward various types of accidents. In general, for nitride fueled core, the core radial expansion coefficient and Doppler coefficient are the most dominant feedbacks, while for metallic fueled cores, the core radial expansion coefficient and fuel axial expansion coefficient are the most important components. The coolant density coefficient is generally small but for the peak burnup of about 9-10% HM. For pancake and balance cores, the modified design gives better coolant void coefficient than the original design, but for tall-slender core the original design gives slightly better results. The pancake shape with H/D of 2/3 gives the most negative coolant void- coefficient in the present design. The coolant void coefficient generally takes the least negative value near the EOL. The present results show that either plutonium composition (pure 239Pu or the equilibrium composition) can be used to satisfy the present design constraints. From thermal-hydraulics analysis, it can be concluded that the fixed orifice pattern does not cause any problems to the long-life small safe reactor proposed in the present design though the power density distribution changes along burnup. The necessary pumping power is also small relative to the conventional large sodium cooled LMFBR though lead or lead-bismuth is used as the coolant. In general, with the chimney of -2m long, the natural circulation level is about 25-40% Table 3 Reactivity feedback coefficient 18

12 Vol. 32, No. 9 (Sep. 1995) 845 of the total flow-rate for the pancake design with H/D of -2/3. The longer chimney enhances natural circulation level. This relatively large value of natural circulation becomes very useful to assure safety against accidents especially that include loss of primary pumping power. The Doppler, coolant density, core radial expansion and fuel axial expansion reactivity feedback coefficients are all negative. For the nitride fueled reactors the core radial expansion and Doppler coefficients are the most important components, while for the metallic fueled reactors the core radial expansion and fuel axial expansion coefficients are the most important components. REFERENCES (1) ZAKI, S., SEKIMOTO, H.: "Potential of Small Nuclear Reactors for Future Clean and Safe Energy Sources", 225 (1992), Elsevier, Amsterdam. (2) ZAKI, S., SEKIMOTO, H.: Proc. ICENES'93 Conf., Makuhari-Japan, 316 (1993). (3) SEKIMOTO, H., ZAKI, S.: Nucl. Technol., 109[3], 307 (1995). (4) SEKIMOTO, H., TAKAGI, N.: J. Nucl. Sci. Technol., 28[10], 941 (1991). (5) ZAKI, S., SEKIMOTO, H.: Nucl. Eng. Des., 140, 251 (1993). (6) ZAKI, S.: Design study on lead and lead-bismuth cooled long-life small fast reactors, Doctor Dissertation, Dept. of Nucl. Eng., TIT, (1995), (in Japanese). (7) NAKAGAWA, M., TSUCHIHASHI, K.: JAERI M- 1294, (1984). 19