- Email: [email protected]
ENHS reactor power level enhancement possibilities
Available online at www.sciencedirect.com
Progress in Nuclear Energy 50 (2008) 279e285 www.elsevier.com/locate/pnucene
ENHS reactor power level enhancement possibilities Arnaud Susplugas, Ehud Greenspan* Department of Nuclear Engineering, University of California, Berkeley, CA 94720, USA
Abstract The ENHS thermal hydraulic optimization code was modified and applied to search for the maximum attainable power from a wide range of ENHS design options subjected to the following constraints: maximum permissible hot channel coolant outlet temperature of 600 C, clad inner temperature of 650 C and primary coolant temperature rise of either 150 C or 90% of the theoretical limit for accelerated corrosion rate. The TeH optimization variables include the intermediate heat exchanger number of channels, channel width and elevation; diameter of the riser and diameter of a flow-splitting shroud in the riser. It was found possible to increase the attainable power from the nominal 125 MWth up to 311 MWth for the reference core, 400 MWth for a reference-like core having equilibrium composition fuel and 372 MWth for a flattened power core with 9 plutonium concentration zones. A power level exceeding 400 MWth may be achieved by flattening the power distribution of the equilibrium core or using nitride fuel with enriched nitrogen rather than metallic fuel. With forced circulation it is possible to operate the flattened power core at up to 532 MWth corresponding to 223 MWe. Ó 2007 Elsevier Ltd. All rights reserved. Keywords: Encapsulated Nuclear Heat Source; ENHS; Power; Specific power; Thermal-hydraulics; LeadeBismuth coolant; Natural circulation; Forced circulation
1. Introduction The power level of the Encapsulated Nuclear Heat Source (ENHS) reactor (Greenspan et al., in press) and, hence the power density and the specific power are limited by the coolant flow rate attainable using natural circulation. The coolant mass flow rate is that for which the total friction losses balance the hydrostatic head established due to coolant temperature difference. A major contributor to the coolant friction is the flow through the core. For a given coolant flow rate and core height, the larger the pitch-to-diameter (P/D) ratio the core can be designed to have, the smaller are the core friction losses. The other major contributor to the coolant friction is the flow through the Intermediate Heat Exchanger (IHX). Although a preliminary economic analysis of the reference ENHS reactor design concluded that it is economically viable (Stewart et al., 2002), an increase in the power density and in the specific power will improve its economic viability.
The reference ENHS core uses metallic fuel of the IFR type. The heavy metal (HM) of this fuel is made of plutonium extracted from LWR fuel discharged at 50 GWD/tHM and cooled for 10 years. The required P/D ratio is 1.361 and the corresponding attainable power is 125 MWth when the coolant average outlet temperature is 500 C (Sienicki, 2003). Several alternative core designs having larger P/D ratios have been identified over the last few years (Hong et al., 2005; Hong and Greenspan, 2004, 2005; Monti et al., in press). One objective of this work is to quantify the power level enhancement made possible by the increase in the P/D ratio as well as by an increase in the temperature drop across the core from 100 K to 150 K and in certain cases, to w180 K. Power enhancement due to flattening of the power density across the core is also quantified. A second objective is to explore farther power enhancement possibilities by better optimizing the Intermediate Heat Exchanger (IHX) and the riser designs. A third objective of this work is to estimate the power
1
* Corresponding author. E-mail address: [email protected] (E. Greenspan). 0149-1970/$ - see front matter Ó 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.pnucene.2007.11.001
The reference P/D ratio quoted in some of the ENHS publications is 1.34 rather than 1.36. This difference is due to use of a different composition of the initial fuel loading.
280
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
level the ENHS can operate at when using forced rather than natural circulation. The design modifications analyzed to meet each of the objectives and corresponding power gains attained are described in, respectively, Sections 3e5 following a brief description, in Section 2, of the methodology and assumptions.
2. Methodology 2.1. ENHS modelled Fig. 1 shows simplified schematic views of the ENHS reactor and the flow path of the primary and intermediate coolant systems. The PbeBi primary coolant enters the core from the bottom after passing through the core support plate, flows along the fuel rods and fission gas plena, both 125 cm long, into the riser. At the top of the riser it turns over into the downcomer, through the upper slots in the inner structural walls, where it enters the IHX and flows in-between the IHX rectangular channels, then passes down through the shield section (located below the IHX, at the core level) and back to the coolant cavity below the core support plate through the lower slots in the inner structural wall. The intermediate coolant, also PbeBi, gets into the IHX channels at their bottom through slots (not shown in the figure) in the outer structural wall, flows up and exits the IHX through slots (not shown in the figure) in the outer structural wall from where it gets into 8 steam generators located in the intermediate coolant pool and arranged around the ENHS module, flows down out of the steam generators to the bottom of the pool where it turns over and flows upwards on the inner side of the flow partition
(aimed at eliminating PbeBi stagnation at the bottom of the pool) into the IHX through the lower slots. 2.2. Thermal-hydraulic model The thermal-hydraulic simulation of the ENHS was done using an expanded version of the ENHS Thermal Hydraulic Calculation (ETHC) code developed for the primary loop by Okawa (2006) on the basis of the simplified analytic model, based on first principles, developed by Sienicki (2003). The expanded code is designed to optimize the ENHS for maximizing electricity generation. For this purpose the code was expanded to also simulate the intermediate PbeBi loop and steam generators. New search algorithms were implemented to optimize geometrical in addition to the thermal-hydraulic parameters, such as coolant flow rates and temperatures in primary, intermediate and steam loops. A number of optional features were added to the code including using a riser of reduced diameter combined with use of a wider IHX channels; insertion of a flow partition shroud into the riser; varying the elevation of the IHX by reducing its length but keeping its uppermost location fixed; increasing the maximum permissible coolant temperature rise in the core to 90% of the maximum value beyond which enhanced corrosion is expected (see Section 3.4). The upgraded ETHC code was benchmarked against Sienicki (2003) and was found satisfactory. Details can be found in Sienicki (2003), Okawa and Greenspan (2006) and Susplugas and Greenspan (2005). Following is a summary of assumptions used for the thermal-hydraulic analysis. The riser as well as the reference IHX length is 13 m, making the separation distance between the thermal centers of the heat source (core) and heat sink (IHX) 8.725 m. The primary
Outer structural wall Riser Upper slots
Steam generators
Secondary coolant
Inner structural wall
IHX Channel Downcomer
Primary coolant Heat exchangers Peripheral control assembly Central control assembly
Core
Lower Slots
Core
Fig. 1. Simplified view of the reference ENHS module.
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
coolant inlet temperature is 450 C. A no-reheat Rankine cycle having 120 bar and 482 C inlet steam conditions and 34 C (0.056 bar) condenser was chosen for the reference energyconversion (EC) system (Greenspan, in press). The feedwater inlet temperature is 242 C and the intermediate loop inlet temperature is 0.6 Tinlet (feedwater) þ 0.4 Tinlet (primary). This value is a compromise between the performance and dimensions of the IHX and steam generator; the numerical values 0.4 and 0.6 were set to match the reference case temperatures. The intermediate coolant outlet temperature is 30 C below, and the steam outlet temperature 60 C below the core outlet temperature. The code accounts for pressure drop due to primary coolant flow through the fuel rods and fission gas plena, 5 grid spacers in the core, sudden expansion at core exit, flow through riser, contraction at entrance and expansion at exit from the upper slots, contraction at entrance to the IHX, flow through the IHX, expansion at exit from IHX, flow through radial shield section, contraction at entrance and expansion at exit from the lower slots, and contraction at entrance and expansion at exit from the core support plate. Similar pressure loss components are accounted for the intermediate coolant flow: friction in the IHX and in the steam generators, including through 5 grid spacers, and sudden expansion/extraction pressure drops and entrance and exit of ENHS module, IHX and steam generators. The constraint imposed on the design are maximum primary coolant temperature of 600 C, maximum clad temperature of 650 C, maximum metallic fuel temperature of 1040 C and maximum temperature rise across the core of 150 C. At a later part of the study this DT constraint has been relaxed, as described in Section 3.4.
281
study assumes a maximum permissible coolant temperature rise across the core of 150 C, straight riser, no internal shroud but optimizes the IHX elevation. The results are summarized in Fig. 2. The maximum thermal power, reached at P/D w 1.64, is 240 MWth, while the maximum electrical power, obtained for P/D w 1.62, is approximately 98 MWe. The attainable power decreases for larger P/D ratios because of a reduction in the clad-to-coolant heat transfer coefficient that is due to a reduction in the coolant velocity and, therefore, in the coolant turbulence close to the fuel rods. The optimal IHX elevation, not shown in the figure, drops close to linearly with P/D from 7 m at around P/D of 1.22 to zero at P/D w 1.64 and stays at zero for larger P/D ratios. The reference core (P/D ¼ 1.36) power is now predicted to be 177 MWth versus 125 MWth of the original design (Sienicki, 2003). Of the 52 MWth difference, 42 MWth are due to the fact that the original design did not try to reach the temperature constraints adopted for the present work and used different correlation for the pressure drop through the grid spacers (Susplugas and Greenspan, 2005) while 10 MWth are due to IHX elevation optimization; the IHX length was reduced by 476 cm thus decreasing the IHX pressure drop and increasing the thermal center separation height. Alternatively, if the power is to be maintained at 125 MWth, it is possible to reduce the IHX riser (and, hence, ENHS module length) by 6.75 m using a 20 cm IHX elevation. Fig. 3 shows the variation of selected temperatures with P/D at the maximum attainable power of Fig. 2. It is found that what limits the attainable power is the primary coolant outlet temperature when P/D < w1.45 and maximum clad temperature for larger P/D ratios.
2.3. Optimization strategy For a given core, the upgraded ETHC code searchers for that IHX configuration that can remove the maximum power from the core subjected to the temperature constraints defined above. The design variables include the core power, the number of IHX channels, their thickness and elevation. Additional optional design variables include the riser inner diameter and the shroud diameter. Temperatures in the intermediate loop and in the energy conversion system are derived so as to maximize the electrical energy generation.
3.2. Power attainable from selected ENHS cores Twelve specific core designs were analyzed accounting for the geometry and power density distribution of the specific core. All cores, except where noted otherwise, use metallic PuUZr(10) fuel and 9862 fuel rods that are 1.56 cm in outer diameter, 125 cm in length and of uniform initial composition. The 12 different cores considered are: Optimal IHX Elevation
3. Results 300
3.1. Effect of core P/D ratio on the attainable power
250
Power (MW)
A number of alternative cores were designed for the ENHS over the last few years. These cores feature the reference fuel rod diameter, D, but different lattice pitch, P. If P/ D increases, the pressure drop in the core decreases and the attainable power goes up. A parametric study was first undertaken to quantify the attainable power dependence on the core P/D assuming that the relative power density distribution in all cores is the same as in the reference core. The
Optimization with constant Maximum Delta-T across core (150 °C)
200 Attainable Electrical Power Attainable Thermal Power
150 100 50 0 1.2
1.3
1.4
1.5
1.6
1.7
1.8
P/D Fig. 2. Attainable power as a function of the core pitch-to-diameter (P/D) ratio.
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
282 700
Temperature (°C)
650 600 550 500 450 Tmax CLADDING (out) Tmax CLADDING (in) Tinlet Tmax COOLANT
400 350 300 1.1
1.2
1.3
1.4
1.5
1.6
1.7
1.8
1.9
P/D Fig. 3. Selected core temperatures as a function of the core P/D for the peak power designs of Fig. 2.
1. The reference core; design I in Hong et al. (2005), 2. Core fueled with PuNeUN using natural N; design IIa in Hong et al. (2005), 3. Core fueled with PuNeUN using enriched N; design IIb in Hong et al. (2005), 4. Core using Na for the primary and intermediate coolants; design IIIa in Hong et al. (2005), 5. Core using Na primary and PbeBi intermediate coolants; design IIIb in Hong et al. (2005), 6. Core with 3 radial Pu concentration levels (Hong and Greenspan, 2004), 7. Core with 3 radial and 3 axial plutonium concentration levels (Hong and Greenspan, 2004), 8. Core using 3 different fuel diameters of identical composition (Hong and Greenspan, 2004), 9. Small core with no minor actinides (MA) e S-1 in Hong and Greenspan (2005), 10. Small core with 15w/o MA and 115 cm fuel length e SM-1 in Hong and Greenspan (2005), 11. Small core with 30w/o MA and 115 cm fuel length e SM-2 in Hong and Greenspan (2005), 12. Small core with 50w/o MA and 115 cm fuel length e SM-3 in Hong and Greenspan (2005).
The number of fuel rods in core 9 is 1926 and 2094 in cores 10e12. The outer fuel diameter in these cores is 2.50 cm. Table 1 gives the attainable power as well as specific power and selected temperatures of ENHS reactors using these cores assuming maximum permissible coolant temperature rise across the core of 150 C, fixed (rather than optimized; as in the previous section) IHX elevation, straight riser and no internal shroud. The maximum power of 220 MWth is provided by the core that uses 9 Pu concentration zones for power flattening. Closely following is the core having 3 radial fuel rod diameters of equal composition. The SM-3 core offers comparable specific power to the best full size cores, when measured in terms of the total HM inventory. The ‘‘equilibrium core’’ given in line 13 of Table 1 is similar to the reference core except that it uses TRU that has been recycled many times through ENHS cores (Monti et al., in press). Its P/D ratio is close to that offering the maximum possible power (Fig. 1), without using power flattening. Superimposing power flattening on the equilibrium core is likely to enable increasing its power level and specific power up to approximately (212/167*229¼) 291 MWth and 16.6 MWth/tHM or 136 MWth/tTRU. These are nearly 2.3 times larger than of the original reference ENHS design (Greenspan, in press).
3.3. Riser and IHX optimization The effect of three design variations in the riser and IHX on the attainable power was investigated for the split plutonium concentration core (#7) as well as for the reference core with fuel composition of the first core (#1) and of the equilibrium core (#13). The first design variation is optimization of the elevation of the IHX while keeping its upper level fixed (as done in Section 3.1 but not in Section 3.2), thus increasing the thermal centers separation distance. The optimization variables are the IHX elevation and the number and width of the IHX channels. The resulting attainable powers are given in Table 2. A second design variation, illustrated in Fig. 4 (left), is reducing the riser diameter from 2.46 m to about 1.5 m while
Table 1 Attainable power and specific power from different ENHS designs
1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13.
Core type
P/D
Peak/av power
Power MWth
MWth/tHM
MWth/tTRU
Av. out T ( C)
Peak temp., C Clad
Fuel
Reference PuNeUNnatural PuNeU15N Na/Na Na/PbeBi 3 Radial Pu 3 Radial/3 axial 3 Fuel diameters Small core, S-1 SM-1 SM-2 SM-3 Equilibrium
1.36 1.21 1.45 1.16 1.20 1.36 1.36 Varies 1.07 1.10 1.19 1.40 1.53
1.773 1.893 1.859 1.98 1.98 1.51 1.34 1.42 1.68 1.707 1.647 1.743 1.773
167 78 198 77 114 206 221 212 10.2 17 37 100 229
9.54 4.06 10.31 4.40 6.51 11.77 12.62 11.30 1.17 1.94 4.23 11.43 13.08
78.2 32.8 78.8 37.1 54.3 89.4 92.9 92.7 10.0 13.5 23.3 41.3 107.3
551 546 548 523 530 569 575 573 564 563 559 498 536
630 617 649 612 617 639 644 629 609 614 629 650 650
709 645 748 639 656 715 721 701 624 640 706 910 763
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285 Table 2 Attainable power and selected characteristics of the reference, equilibrium and flattened power cores using wide IHX, flow shroud, IHX elevation and maximum DT Characteristic
Reference core (#1)
Equilibrium core (#13)
Flat power core (#7)
Power prior to optimization, MWth Power with IHX elevation, MWth Power with narrow riser, MWth Power with shroud, MWth Max. attainable power, MWth Estimated electric power, MWe Module Diameter, m IHX elevation, m Primary coolant DT, C Specific power, MWth/tHM Specific power, MWth/tTRU Riser diameter Coolant inlet temperature, C Average outlet temperature, C Maximum coolant temperature, C Maximum clad temperature, C Maximum fuel temperature, C
167 177 209 247 311 128 3.57 5.92 181 17.8 145.7 1.46 402 542 583 650 801
229 230 306 360 400 162 3.84 2.20 169 22.9 187.4 1.52 364 498 533 650 858
221 229 280 302 372 154 3.57 5.39 180 21.3 156.4 1.46 398 558 577 650 791
correspondingly increasing the IHX width. The resulting attainable powers are given in Table 2. The third design variation, illustrated in Fig. 4 (right), involves insertion into the riser a concentric cylindrical shroud of similar shape but smaller diameter. By providing an effectively separate riser for the inner high power density part of the core the coolant circulation driving head in the central riser increases and the coolant flow rate through the high power density region increases as well. The thermal-hydraulic analysis was done assuming that the shroud also splits the core into two radial channels. The shroud inner diameter in the core region is 165 cm; the diameter has not been optimized for each particular design. The resulting attainable powers are compared in Table 2.
Straight Riser
Narrow Riser
Vertical Shroud
Core
Core
Core
Fig. 4. Schematic view of the reduced riser diameter design (left) and of the cylindrical inner shroud (right).
283
3.4. Optimization of operating temperature So far a constraint of 150 C was imposed on the maximum permissible primary coolant temperature rise in the core. This constraint is imposed for the purpose of maintaining the corrosion rate of the clad by the PbeBi coolant at a sufficiently low level. A close examination of the Ellingham diagram for oxides revealed that whereas for low coolant temperatures the 150 C limit is justified, for the temperature range 450e 600 C presently considered for the ENHS the maximum temperature rise across the core for which the structure is protected from corrosion by the coolant by means of formation of a protective Fe3O4 oxide layer without the coolant becoming contaminated by lead oxide (PbO) is significantly higher. A conservative upper bound used is w90% of the maximum theoretical value (Susplugas and Greenspan, 2005): DTmax ð90%Þ ¼ 171 þ 0:225 ðTMax Hot Channel 540Þ
ð1Þ
Table 2 gives the attainable power for the reference, equilibrium and flattened cores along with selected design and performance characteristics. All these systems have a 13 m high riser; the riser diameter is an optimization variable. The module outer diameter is 3.57 m for the reference and flat power core and 3.84 m for the equilibrium core. The equilibrium and flattened cores power and specific powers are larger than those of the original reference ENHS design (125 MWth) by more than a factor of 3. Even more substantial power increase is expected from flattening the power of the core that uses equilibrium TRU composition. 3.5. Power attainable with forced circulation Even though an important design goal of the ENHS is natural circulation cooling, it is useful to know what power level an ENHS like reactor could operate at when using forced circulation, while maintaining sufficient natural circulation capability for safe decay heat removal in case of loss of forced pumping accident. Two types of pumps have been examined for forced circulation e a lift pump and an electro-magnetic pump. The liftpump involves injection of non-condensable gas into the primary coolant above the core and/or above the IHX. The electro-magnetic pump circulates the conductive Pbe Bi coolant using the Lorentz force and thus is free of moving parts and, therefore, of the need for seals and bearings that are needed for mechanical pumps. This simplifies maintenance and improves reliability. It is assumed located in the riser, even though it would have been better to locate it below the IHX where the primary coolant temperature in the lowest. Unfortunately, doing so would have required a substantial increase in the module dimensions. It is estimated that an EM pump, that is compatible with the ENHS module dimensions and does not endanger the reactor safety by constricting too much the primary coolant flow channel in the riser, could be designed to provide up to 200 kW of mechanical power. A design arrived at using a 150 kW EM-pump in the primary
284
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
circuit and a 30 kW EM-pump in the intermediate circuit. This pumping power split is estimated to be the optimal for the ENHS reactor. Using these EM pumps it is estimated that the reference ENHS reactor power can increase from 177 to 360 MWth. An ENHS module featuring a lift pump in the primary loop and a 200 kW EM pump in the intermediate loop was also designed. The lift pump in the primary enables to use a narrow riser and, hence, a wider IHX. The module diameter is increased by 20 cm, but the overall length is reduced by 6 m. The attainable power is estimated 530 MWth or 220 MWe. A simplified safety analysis showed that in case of pump outage, this design can still remove up to 60% of the nominal power by natural circulation. Table 3 summarizes selected characteristics of this design.
needs be re-visited as well for designs featuring smaller diameter risers and wider IHX. Particular attention needs to be given to the design of the peripheral control elements drive mechanisms. The use of narrow riser will not interfere with the core loading as it is planned to be done from the bottom of the ENHS module. Further design optimization, including flattening of the power distribution of cores having the equilibrium fuel composition, is likely to identify designs offering even higher power levels than reached in this work. Most promising is expected to be nitride fueled cores using nitrogen enriched with 15N. Thermal-hydraulic optimization of ‘‘stepped geometry’’ cores (Okawa and Greenspan, 2004) is also desirable. 5. Conclusions
4. Discussion Although indicative of the power attainable from different ENHS designs, the actual power that could be achieved is likely to be somewhat different from the values quoted in this paper due to the feedback the thermal-hydraulic and neutronic performance of ENHS reactors. In particular, for a given fuel composition, the P/D ratio required for providing a nearly burnup-independent keff depends on the power density (Monti et al., in press). Consequently, it is necessary to perform at least one additional iteration of neutronic analysis to determine the P/D and power distribution using the attainable power levels quantified in the present work, to be followed by a revised thermal-hydraulic analysis. Thorough safety analysis is also required before the power attained from the different core and module designs could be reliably determined. The mechanical design of the ENHS Table 3 Selected characteristics of an ENHS reactor using the reference core and forced circulation by a lift pump in the primary and an EM pump in the intermediate coolant loops Characteristic
Pu wt% split (3 radial, 3 axial)
Attainable power, MWth Attainable power, MWe Estimated EC efficiency, % Steam outlet temperature, C Riser height, m Module Diameter, m Riser diameter, m Shroud in the riser Primary loop Pump Intermediate loop Pump Primary pumping power, KW Intermediate pumping power, KW Average outlet temperature, C Maximum coolant temperature, C Cladding Maximum temperature, C DT in hot channel Number of IHX channels Removable power if primary pump is out, MWth
532 223 40.8 464 7 3.77 1.46 No Lift pump EM pump 200 30 524 549 650 150 424 318 (60%)
Assuming a maximum permissible hot channel coolant outlet temperature of 600 C, clad inner temperature of 650 C and primary coolant temperature rise of 150 C, refining the correlation for pressure drop through grid spacers and optimizing for the IHX elevation it was found that the maximum attainable power from the reference ENHS is 177 MWth and that the core pitch-to-diameter (P/D) ratio giving the maximum power of 240 MWth, for a power distribution as of the reference core, is 1.62. The attainable power decreases for larger P/D ratios because of a reduction in the clad-to-coolant heat transfer coefficient that is due to a reduction in the coolant velocity. The attainable power is limited by the primary coolant outlet temperature when P/D < w1.45 and by the maximum clad temperature for larger P/D ratios. Of the 12 ENHS designs featuring different cores analyzed, the maximum power of 221 MWth is offered by the core having flattened power distribution achieved using split plutonium concentration in 3 radial and 3 axial zones, followed closely by the core made of 3 radial zones having fuel rods of increasing diameters. A comparable high power of 229 MWth is attainable from a reference-like ENHS that uses the equilibrium core composition for which the required P/D is 1.53. Flattening the power of the equilibrium core is expected to result in significantly higher power level. By reducing the riser diameter and correspondingly increasing the IHX width and by introducing a flow-splitting shroud into the riser it is possible to increase the attainable power up to 247, 302 and 360 MWth for, respectively, the reference core, the 9-zone flattened power core and the unflattened equilibrium core. By relaxing the maximum coolant temperature drop across the core from 150 C to 90% of the theoretical maximum limit the corresponding attainable power levels are 311, 400 and 372 MWth. Even higher power level may be achieved by flattening the power distribution of the equilibrium core and by using nitride fuel with enriched nitrogen rather than metallic alloy fuel. With forced circulation in the form of a lift pump for the primary coolant and an electro-magnetic pump for the intermediate coolant it is possible to operate the flattened core at up to 532 MWth corresponding to 223 MWe. In case
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
of pump trip accident, natural circulation could remove approximately 60% of the maximum power, thus providing an effective passive decay heat removal capability. Additional iteration with the neutronic, safety and structural analyses is required to more accurately determine the attainable power. Nevertheless, the obtained results indicate that the power and specific power of the ENHS reactor can be at least 3 times higher than of the original reference design. This is likely to significantly improve the ENHS economic viability. Acknowledgment This work was supported by the Lawrence Livermore National Laboratory under contract number B535246. References Greenspan, E., and the ENHS design team, Innovations in the ENHS reactor design and fuel cycle. In: The Proceedings of the Second COE-INES International Symposium on Innovative Nuclear Energy Systems, INES-2, Yokohama, Japan, 26e30 November, in press. Greenspan, E., ENHS general information, technical features, and operating characteristics. In: The IAEA Status Report on Innovative Small and Medium-Sized Reactors (SMRs), in press.
285
Hong, S.G., Greenspan, E., 2004. Power flattening options for the ENHS (Encapsulated Nuclear Heat Source) core. In: Proceedings of the International Symposium on Innovative Nuclear Energy Systems, Tokyo, Japan, October 31eNovember 4. Hong, S.G., Greenspan, E., Kim, Y.I., 2005. The Encapsulated Nuclear Heat Source (ENHS) reactor core design. Nuclear Technology 149, 22. Hong, S.G., Greenspan, E., 2005. Use of minor actinides for reduced power ENHS reactor core design. In: Proceedings of GLOBAL’05, Tsukuba, Japan, October 9e13. Monti, L., Greenspan, E., Sumini, M., Fratoni, M., Rocchi, F., Multi-recycling in the ENHS. In: The Proceedings of the Second COE-INES International Symposium on Innovative Nuclear Energy Systems, INES-2, Yokohama, Japan, 26e30 November, in press. Okawa, T., Greenspan, E., 2004. Alternative ENHS core design using stepped geometry core. Transactions of American Nuclear Society 91. Okawa, T., Greenspan, E., 2006. Effect of fuel type on the attainable power of the Encapsulated Nuclear Heat Source reactor, Proceedings of the ICAPP2006, Reno, NV, 4e7 June. Sienicki, J.J., 2003. Updated thermal hydraulic analysis for recent ENHS design improvements, International Congress on Advances in Nuclear Power Plants, ICAPP 2003, Cordoba, Spain. Stewart, J.S., Lamont, A.D., Rothwell, G.S., Smith, C.F., Greenspan, E., Brown, N., Barak, A., 2002. An Economic Analysis of Generation IV Small Modular Reactors, Lawrence Livermore National Laboratory Report UCRL-ID-148437. Susplugas, A., Greenspan, E., 2005. Implementation of the ENHS Thermal Hydraulic Optimization Code for Recent ENHS Design Improvements. University of California, Department of Nuclear Engineering. Internal Report UCB-NE-5104.
Progress in Nuclear Energy 50 (2008) 279e285 www.elsevier.com/locate/pnucene
ENHS reactor power level enhancement possibilities Arnaud Susplugas, Ehud Greenspan* Department of Nuclear Engineering, University of California, Berkeley, CA 94720, USA
Abstract The ENHS thermal hydraulic optimization code was modified and applied to search for the maximum attainable power from a wide range of ENHS design options subjected to the following constraints: maximum permissible hot channel coolant outlet temperature of 600 C, clad inner temperature of 650 C and primary coolant temperature rise of either 150 C or 90% of the theoretical limit for accelerated corrosion rate. The TeH optimization variables include the intermediate heat exchanger number of channels, channel width and elevation; diameter of the riser and diameter of a flow-splitting shroud in the riser. It was found possible to increase the attainable power from the nominal 125 MWth up to 311 MWth for the reference core, 400 MWth for a reference-like core having equilibrium composition fuel and 372 MWth for a flattened power core with 9 plutonium concentration zones. A power level exceeding 400 MWth may be achieved by flattening the power distribution of the equilibrium core or using nitride fuel with enriched nitrogen rather than metallic fuel. With forced circulation it is possible to operate the flattened power core at up to 532 MWth corresponding to 223 MWe. Ó 2007 Elsevier Ltd. All rights reserved. Keywords: Encapsulated Nuclear Heat Source; ENHS; Power; Specific power; Thermal-hydraulics; LeadeBismuth coolant; Natural circulation; Forced circulation
1. Introduction The power level of the Encapsulated Nuclear Heat Source (ENHS) reactor (Greenspan et al., in press) and, hence the power density and the specific power are limited by the coolant flow rate attainable using natural circulation. The coolant mass flow rate is that for which the total friction losses balance the hydrostatic head established due to coolant temperature difference. A major contributor to the coolant friction is the flow through the core. For a given coolant flow rate and core height, the larger the pitch-to-diameter (P/D) ratio the core can be designed to have, the smaller are the core friction losses. The other major contributor to the coolant friction is the flow through the Intermediate Heat Exchanger (IHX). Although a preliminary economic analysis of the reference ENHS reactor design concluded that it is economically viable (Stewart et al., 2002), an increase in the power density and in the specific power will improve its economic viability.
The reference ENHS core uses metallic fuel of the IFR type. The heavy metal (HM) of this fuel is made of plutonium extracted from LWR fuel discharged at 50 GWD/tHM and cooled for 10 years. The required P/D ratio is 1.361 and the corresponding attainable power is 125 MWth when the coolant average outlet temperature is 500 C (Sienicki, 2003). Several alternative core designs having larger P/D ratios have been identified over the last few years (Hong et al., 2005; Hong and Greenspan, 2004, 2005; Monti et al., in press). One objective of this work is to quantify the power level enhancement made possible by the increase in the P/D ratio as well as by an increase in the temperature drop across the core from 100 K to 150 K and in certain cases, to w180 K. Power enhancement due to flattening of the power density across the core is also quantified. A second objective is to explore farther power enhancement possibilities by better optimizing the Intermediate Heat Exchanger (IHX) and the riser designs. A third objective of this work is to estimate the power
1
* Corresponding author. E-mail address: [email protected] (E. Greenspan). 0149-1970/$ - see front matter Ó 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.pnucene.2007.11.001
The reference P/D ratio quoted in some of the ENHS publications is 1.34 rather than 1.36. This difference is due to use of a different composition of the initial fuel loading.
280
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
level the ENHS can operate at when using forced rather than natural circulation. The design modifications analyzed to meet each of the objectives and corresponding power gains attained are described in, respectively, Sections 3e5 following a brief description, in Section 2, of the methodology and assumptions.
2. Methodology 2.1. ENHS modelled Fig. 1 shows simplified schematic views of the ENHS reactor and the flow path of the primary and intermediate coolant systems. The PbeBi primary coolant enters the core from the bottom after passing through the core support plate, flows along the fuel rods and fission gas plena, both 125 cm long, into the riser. At the top of the riser it turns over into the downcomer, through the upper slots in the inner structural walls, where it enters the IHX and flows in-between the IHX rectangular channels, then passes down through the shield section (located below the IHX, at the core level) and back to the coolant cavity below the core support plate through the lower slots in the inner structural wall. The intermediate coolant, also PbeBi, gets into the IHX channels at their bottom through slots (not shown in the figure) in the outer structural wall, flows up and exits the IHX through slots (not shown in the figure) in the outer structural wall from where it gets into 8 steam generators located in the intermediate coolant pool and arranged around the ENHS module, flows down out of the steam generators to the bottom of the pool where it turns over and flows upwards on the inner side of the flow partition
(aimed at eliminating PbeBi stagnation at the bottom of the pool) into the IHX through the lower slots. 2.2. Thermal-hydraulic model The thermal-hydraulic simulation of the ENHS was done using an expanded version of the ENHS Thermal Hydraulic Calculation (ETHC) code developed for the primary loop by Okawa (2006) on the basis of the simplified analytic model, based on first principles, developed by Sienicki (2003). The expanded code is designed to optimize the ENHS for maximizing electricity generation. For this purpose the code was expanded to also simulate the intermediate PbeBi loop and steam generators. New search algorithms were implemented to optimize geometrical in addition to the thermal-hydraulic parameters, such as coolant flow rates and temperatures in primary, intermediate and steam loops. A number of optional features were added to the code including using a riser of reduced diameter combined with use of a wider IHX channels; insertion of a flow partition shroud into the riser; varying the elevation of the IHX by reducing its length but keeping its uppermost location fixed; increasing the maximum permissible coolant temperature rise in the core to 90% of the maximum value beyond which enhanced corrosion is expected (see Section 3.4). The upgraded ETHC code was benchmarked against Sienicki (2003) and was found satisfactory. Details can be found in Sienicki (2003), Okawa and Greenspan (2006) and Susplugas and Greenspan (2005). Following is a summary of assumptions used for the thermal-hydraulic analysis. The riser as well as the reference IHX length is 13 m, making the separation distance between the thermal centers of the heat source (core) and heat sink (IHX) 8.725 m. The primary
Outer structural wall Riser Upper slots
Steam generators
Secondary coolant
Inner structural wall
IHX Channel Downcomer
Primary coolant Heat exchangers Peripheral control assembly Central control assembly
Core
Lower Slots
Core
Fig. 1. Simplified view of the reference ENHS module.
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
coolant inlet temperature is 450 C. A no-reheat Rankine cycle having 120 bar and 482 C inlet steam conditions and 34 C (0.056 bar) condenser was chosen for the reference energyconversion (EC) system (Greenspan, in press). The feedwater inlet temperature is 242 C and the intermediate loop inlet temperature is 0.6 Tinlet (feedwater) þ 0.4 Tinlet (primary). This value is a compromise between the performance and dimensions of the IHX and steam generator; the numerical values 0.4 and 0.6 were set to match the reference case temperatures. The intermediate coolant outlet temperature is 30 C below, and the steam outlet temperature 60 C below the core outlet temperature. The code accounts for pressure drop due to primary coolant flow through the fuel rods and fission gas plena, 5 grid spacers in the core, sudden expansion at core exit, flow through riser, contraction at entrance and expansion at exit from the upper slots, contraction at entrance to the IHX, flow through the IHX, expansion at exit from IHX, flow through radial shield section, contraction at entrance and expansion at exit from the lower slots, and contraction at entrance and expansion at exit from the core support plate. Similar pressure loss components are accounted for the intermediate coolant flow: friction in the IHX and in the steam generators, including through 5 grid spacers, and sudden expansion/extraction pressure drops and entrance and exit of ENHS module, IHX and steam generators. The constraint imposed on the design are maximum primary coolant temperature of 600 C, maximum clad temperature of 650 C, maximum metallic fuel temperature of 1040 C and maximum temperature rise across the core of 150 C. At a later part of the study this DT constraint has been relaxed, as described in Section 3.4.
281
study assumes a maximum permissible coolant temperature rise across the core of 150 C, straight riser, no internal shroud but optimizes the IHX elevation. The results are summarized in Fig. 2. The maximum thermal power, reached at P/D w 1.64, is 240 MWth, while the maximum electrical power, obtained for P/D w 1.62, is approximately 98 MWe. The attainable power decreases for larger P/D ratios because of a reduction in the clad-to-coolant heat transfer coefficient that is due to a reduction in the coolant velocity and, therefore, in the coolant turbulence close to the fuel rods. The optimal IHX elevation, not shown in the figure, drops close to linearly with P/D from 7 m at around P/D of 1.22 to zero at P/D w 1.64 and stays at zero for larger P/D ratios. The reference core (P/D ¼ 1.36) power is now predicted to be 177 MWth versus 125 MWth of the original design (Sienicki, 2003). Of the 52 MWth difference, 42 MWth are due to the fact that the original design did not try to reach the temperature constraints adopted for the present work and used different correlation for the pressure drop through the grid spacers (Susplugas and Greenspan, 2005) while 10 MWth are due to IHX elevation optimization; the IHX length was reduced by 476 cm thus decreasing the IHX pressure drop and increasing the thermal center separation height. Alternatively, if the power is to be maintained at 125 MWth, it is possible to reduce the IHX riser (and, hence, ENHS module length) by 6.75 m using a 20 cm IHX elevation. Fig. 3 shows the variation of selected temperatures with P/D at the maximum attainable power of Fig. 2. It is found that what limits the attainable power is the primary coolant outlet temperature when P/D < w1.45 and maximum clad temperature for larger P/D ratios.
2.3. Optimization strategy For a given core, the upgraded ETHC code searchers for that IHX configuration that can remove the maximum power from the core subjected to the temperature constraints defined above. The design variables include the core power, the number of IHX channels, their thickness and elevation. Additional optional design variables include the riser inner diameter and the shroud diameter. Temperatures in the intermediate loop and in the energy conversion system are derived so as to maximize the electrical energy generation.
3.2. Power attainable from selected ENHS cores Twelve specific core designs were analyzed accounting for the geometry and power density distribution of the specific core. All cores, except where noted otherwise, use metallic PuUZr(10) fuel and 9862 fuel rods that are 1.56 cm in outer diameter, 125 cm in length and of uniform initial composition. The 12 different cores considered are: Optimal IHX Elevation
3. Results 300
3.1. Effect of core P/D ratio on the attainable power
250
Power (MW)
A number of alternative cores were designed for the ENHS over the last few years. These cores feature the reference fuel rod diameter, D, but different lattice pitch, P. If P/ D increases, the pressure drop in the core decreases and the attainable power goes up. A parametric study was first undertaken to quantify the attainable power dependence on the core P/D assuming that the relative power density distribution in all cores is the same as in the reference core. The
Optimization with constant Maximum Delta-T across core (150 °C)
200 Attainable Electrical Power Attainable Thermal Power
150 100 50 0 1.2
1.3
1.4
1.5
1.6
1.7
1.8
P/D Fig. 2. Attainable power as a function of the core pitch-to-diameter (P/D) ratio.
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
282 700
Temperature (°C)
650 600 550 500 450 Tmax CLADDING (out) Tmax CLADDING (in) Tinlet Tmax COOLANT
400 350 300 1.1
1.2
1.3
1.4
1.5
1.6
1.7
1.8
1.9
P/D Fig. 3. Selected core temperatures as a function of the core P/D for the peak power designs of Fig. 2.
1. The reference core; design I in Hong et al. (2005), 2. Core fueled with PuNeUN using natural N; design IIa in Hong et al. (2005), 3. Core fueled with PuNeUN using enriched N; design IIb in Hong et al. (2005), 4. Core using Na for the primary and intermediate coolants; design IIIa in Hong et al. (2005), 5. Core using Na primary and PbeBi intermediate coolants; design IIIb in Hong et al. (2005), 6. Core with 3 radial Pu concentration levels (Hong and Greenspan, 2004), 7. Core with 3 radial and 3 axial plutonium concentration levels (Hong and Greenspan, 2004), 8. Core using 3 different fuel diameters of identical composition (Hong and Greenspan, 2004), 9. Small core with no minor actinides (MA) e S-1 in Hong and Greenspan (2005), 10. Small core with 15w/o MA and 115 cm fuel length e SM-1 in Hong and Greenspan (2005), 11. Small core with 30w/o MA and 115 cm fuel length e SM-2 in Hong and Greenspan (2005), 12. Small core with 50w/o MA and 115 cm fuel length e SM-3 in Hong and Greenspan (2005).
The number of fuel rods in core 9 is 1926 and 2094 in cores 10e12. The outer fuel diameter in these cores is 2.50 cm. Table 1 gives the attainable power as well as specific power and selected temperatures of ENHS reactors using these cores assuming maximum permissible coolant temperature rise across the core of 150 C, fixed (rather than optimized; as in the previous section) IHX elevation, straight riser and no internal shroud. The maximum power of 220 MWth is provided by the core that uses 9 Pu concentration zones for power flattening. Closely following is the core having 3 radial fuel rod diameters of equal composition. The SM-3 core offers comparable specific power to the best full size cores, when measured in terms of the total HM inventory. The ‘‘equilibrium core’’ given in line 13 of Table 1 is similar to the reference core except that it uses TRU that has been recycled many times through ENHS cores (Monti et al., in press). Its P/D ratio is close to that offering the maximum possible power (Fig. 1), without using power flattening. Superimposing power flattening on the equilibrium core is likely to enable increasing its power level and specific power up to approximately (212/167*229¼) 291 MWth and 16.6 MWth/tHM or 136 MWth/tTRU. These are nearly 2.3 times larger than of the original reference ENHS design (Greenspan, in press).
3.3. Riser and IHX optimization The effect of three design variations in the riser and IHX on the attainable power was investigated for the split plutonium concentration core (#7) as well as for the reference core with fuel composition of the first core (#1) and of the equilibrium core (#13). The first design variation is optimization of the elevation of the IHX while keeping its upper level fixed (as done in Section 3.1 but not in Section 3.2), thus increasing the thermal centers separation distance. The optimization variables are the IHX elevation and the number and width of the IHX channels. The resulting attainable powers are given in Table 2. A second design variation, illustrated in Fig. 4 (left), is reducing the riser diameter from 2.46 m to about 1.5 m while
Table 1 Attainable power and specific power from different ENHS designs
1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13.
Core type
P/D
Peak/av power
Power MWth
MWth/tHM
MWth/tTRU
Av. out T ( C)
Peak temp., C Clad
Fuel
Reference PuNeUNnatural PuNeU15N Na/Na Na/PbeBi 3 Radial Pu 3 Radial/3 axial 3 Fuel diameters Small core, S-1 SM-1 SM-2 SM-3 Equilibrium
1.36 1.21 1.45 1.16 1.20 1.36 1.36 Varies 1.07 1.10 1.19 1.40 1.53
1.773 1.893 1.859 1.98 1.98 1.51 1.34 1.42 1.68 1.707 1.647 1.743 1.773
167 78 198 77 114 206 221 212 10.2 17 37 100 229
9.54 4.06 10.31 4.40 6.51 11.77 12.62 11.30 1.17 1.94 4.23 11.43 13.08
78.2 32.8 78.8 37.1 54.3 89.4 92.9 92.7 10.0 13.5 23.3 41.3 107.3
551 546 548 523 530 569 575 573 564 563 559 498 536
630 617 649 612 617 639 644 629 609 614 629 650 650
709 645 748 639 656 715 721 701 624 640 706 910 763
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285 Table 2 Attainable power and selected characteristics of the reference, equilibrium and flattened power cores using wide IHX, flow shroud, IHX elevation and maximum DT Characteristic
Reference core (#1)
Equilibrium core (#13)
Flat power core (#7)
Power prior to optimization, MWth Power with IHX elevation, MWth Power with narrow riser, MWth Power with shroud, MWth Max. attainable power, MWth Estimated electric power, MWe Module Diameter, m IHX elevation, m Primary coolant DT, C Specific power, MWth/tHM Specific power, MWth/tTRU Riser diameter Coolant inlet temperature, C Average outlet temperature, C Maximum coolant temperature, C Maximum clad temperature, C Maximum fuel temperature, C
167 177 209 247 311 128 3.57 5.92 181 17.8 145.7 1.46 402 542 583 650 801
229 230 306 360 400 162 3.84 2.20 169 22.9 187.4 1.52 364 498 533 650 858
221 229 280 302 372 154 3.57 5.39 180 21.3 156.4 1.46 398 558 577 650 791
correspondingly increasing the IHX width. The resulting attainable powers are given in Table 2. The third design variation, illustrated in Fig. 4 (right), involves insertion into the riser a concentric cylindrical shroud of similar shape but smaller diameter. By providing an effectively separate riser for the inner high power density part of the core the coolant circulation driving head in the central riser increases and the coolant flow rate through the high power density region increases as well. The thermal-hydraulic analysis was done assuming that the shroud also splits the core into two radial channels. The shroud inner diameter in the core region is 165 cm; the diameter has not been optimized for each particular design. The resulting attainable powers are compared in Table 2.
Straight Riser
Narrow Riser
Vertical Shroud
Core
Core
Core
Fig. 4. Schematic view of the reduced riser diameter design (left) and of the cylindrical inner shroud (right).
283
3.4. Optimization of operating temperature So far a constraint of 150 C was imposed on the maximum permissible primary coolant temperature rise in the core. This constraint is imposed for the purpose of maintaining the corrosion rate of the clad by the PbeBi coolant at a sufficiently low level. A close examination of the Ellingham diagram for oxides revealed that whereas for low coolant temperatures the 150 C limit is justified, for the temperature range 450e 600 C presently considered for the ENHS the maximum temperature rise across the core for which the structure is protected from corrosion by the coolant by means of formation of a protective Fe3O4 oxide layer without the coolant becoming contaminated by lead oxide (PbO) is significantly higher. A conservative upper bound used is w90% of the maximum theoretical value (Susplugas and Greenspan, 2005): DTmax ð90%Þ ¼ 171 þ 0:225 ðTMax Hot Channel 540Þ
ð1Þ
Table 2 gives the attainable power for the reference, equilibrium and flattened cores along with selected design and performance characteristics. All these systems have a 13 m high riser; the riser diameter is an optimization variable. The module outer diameter is 3.57 m for the reference and flat power core and 3.84 m for the equilibrium core. The equilibrium and flattened cores power and specific powers are larger than those of the original reference ENHS design (125 MWth) by more than a factor of 3. Even more substantial power increase is expected from flattening the power of the core that uses equilibrium TRU composition. 3.5. Power attainable with forced circulation Even though an important design goal of the ENHS is natural circulation cooling, it is useful to know what power level an ENHS like reactor could operate at when using forced circulation, while maintaining sufficient natural circulation capability for safe decay heat removal in case of loss of forced pumping accident. Two types of pumps have been examined for forced circulation e a lift pump and an electro-magnetic pump. The liftpump involves injection of non-condensable gas into the primary coolant above the core and/or above the IHX. The electro-magnetic pump circulates the conductive Pbe Bi coolant using the Lorentz force and thus is free of moving parts and, therefore, of the need for seals and bearings that are needed for mechanical pumps. This simplifies maintenance and improves reliability. It is assumed located in the riser, even though it would have been better to locate it below the IHX where the primary coolant temperature in the lowest. Unfortunately, doing so would have required a substantial increase in the module dimensions. It is estimated that an EM pump, that is compatible with the ENHS module dimensions and does not endanger the reactor safety by constricting too much the primary coolant flow channel in the riser, could be designed to provide up to 200 kW of mechanical power. A design arrived at using a 150 kW EM-pump in the primary
284
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
circuit and a 30 kW EM-pump in the intermediate circuit. This pumping power split is estimated to be the optimal for the ENHS reactor. Using these EM pumps it is estimated that the reference ENHS reactor power can increase from 177 to 360 MWth. An ENHS module featuring a lift pump in the primary loop and a 200 kW EM pump in the intermediate loop was also designed. The lift pump in the primary enables to use a narrow riser and, hence, a wider IHX. The module diameter is increased by 20 cm, but the overall length is reduced by 6 m. The attainable power is estimated 530 MWth or 220 MWe. A simplified safety analysis showed that in case of pump outage, this design can still remove up to 60% of the nominal power by natural circulation. Table 3 summarizes selected characteristics of this design.
needs be re-visited as well for designs featuring smaller diameter risers and wider IHX. Particular attention needs to be given to the design of the peripheral control elements drive mechanisms. The use of narrow riser will not interfere with the core loading as it is planned to be done from the bottom of the ENHS module. Further design optimization, including flattening of the power distribution of cores having the equilibrium fuel composition, is likely to identify designs offering even higher power levels than reached in this work. Most promising is expected to be nitride fueled cores using nitrogen enriched with 15N. Thermal-hydraulic optimization of ‘‘stepped geometry’’ cores (Okawa and Greenspan, 2004) is also desirable. 5. Conclusions
4. Discussion Although indicative of the power attainable from different ENHS designs, the actual power that could be achieved is likely to be somewhat different from the values quoted in this paper due to the feedback the thermal-hydraulic and neutronic performance of ENHS reactors. In particular, for a given fuel composition, the P/D ratio required for providing a nearly burnup-independent keff depends on the power density (Monti et al., in press). Consequently, it is necessary to perform at least one additional iteration of neutronic analysis to determine the P/D and power distribution using the attainable power levels quantified in the present work, to be followed by a revised thermal-hydraulic analysis. Thorough safety analysis is also required before the power attained from the different core and module designs could be reliably determined. The mechanical design of the ENHS Table 3 Selected characteristics of an ENHS reactor using the reference core and forced circulation by a lift pump in the primary and an EM pump in the intermediate coolant loops Characteristic
Pu wt% split (3 radial, 3 axial)
Attainable power, MWth Attainable power, MWe Estimated EC efficiency, % Steam outlet temperature, C Riser height, m Module Diameter, m Riser diameter, m Shroud in the riser Primary loop Pump Intermediate loop Pump Primary pumping power, KW Intermediate pumping power, KW Average outlet temperature, C Maximum coolant temperature, C Cladding Maximum temperature, C DT in hot channel Number of IHX channels Removable power if primary pump is out, MWth
532 223 40.8 464 7 3.77 1.46 No Lift pump EM pump 200 30 524 549 650 150 424 318 (60%)
Assuming a maximum permissible hot channel coolant outlet temperature of 600 C, clad inner temperature of 650 C and primary coolant temperature rise of 150 C, refining the correlation for pressure drop through grid spacers and optimizing for the IHX elevation it was found that the maximum attainable power from the reference ENHS is 177 MWth and that the core pitch-to-diameter (P/D) ratio giving the maximum power of 240 MWth, for a power distribution as of the reference core, is 1.62. The attainable power decreases for larger P/D ratios because of a reduction in the clad-to-coolant heat transfer coefficient that is due to a reduction in the coolant velocity. The attainable power is limited by the primary coolant outlet temperature when P/D < w1.45 and by the maximum clad temperature for larger P/D ratios. Of the 12 ENHS designs featuring different cores analyzed, the maximum power of 221 MWth is offered by the core having flattened power distribution achieved using split plutonium concentration in 3 radial and 3 axial zones, followed closely by the core made of 3 radial zones having fuel rods of increasing diameters. A comparable high power of 229 MWth is attainable from a reference-like ENHS that uses the equilibrium core composition for which the required P/D is 1.53. Flattening the power of the equilibrium core is expected to result in significantly higher power level. By reducing the riser diameter and correspondingly increasing the IHX width and by introducing a flow-splitting shroud into the riser it is possible to increase the attainable power up to 247, 302 and 360 MWth for, respectively, the reference core, the 9-zone flattened power core and the unflattened equilibrium core. By relaxing the maximum coolant temperature drop across the core from 150 C to 90% of the theoretical maximum limit the corresponding attainable power levels are 311, 400 and 372 MWth. Even higher power level may be achieved by flattening the power distribution of the equilibrium core and by using nitride fuel with enriched nitrogen rather than metallic alloy fuel. With forced circulation in the form of a lift pump for the primary coolant and an electro-magnetic pump for the intermediate coolant it is possible to operate the flattened core at up to 532 MWth corresponding to 223 MWe. In case
A. Susplugas, E. Greenspan / Progress in Nuclear Energy 50 (2008) 279e285
of pump trip accident, natural circulation could remove approximately 60% of the maximum power, thus providing an effective passive decay heat removal capability. Additional iteration with the neutronic, safety and structural analyses is required to more accurately determine the attainable power. Nevertheless, the obtained results indicate that the power and specific power of the ENHS reactor can be at least 3 times higher than of the original reference design. This is likely to significantly improve the ENHS economic viability. Acknowledgment This work was supported by the Lawrence Livermore National Laboratory under contract number B535246. References Greenspan, E., and the ENHS design team, Innovations in the ENHS reactor design and fuel cycle. In: The Proceedings of the Second COE-INES International Symposium on Innovative Nuclear Energy Systems, INES-2, Yokohama, Japan, 26e30 November, in press. Greenspan, E., ENHS general information, technical features, and operating characteristics. In: The IAEA Status Report on Innovative Small and Medium-Sized Reactors (SMRs), in press.
285
Hong, S.G., Greenspan, E., 2004. Power flattening options for the ENHS (Encapsulated Nuclear Heat Source) core. In: Proceedings of the International Symposium on Innovative Nuclear Energy Systems, Tokyo, Japan, October 31eNovember 4. Hong, S.G., Greenspan, E., Kim, Y.I., 2005. The Encapsulated Nuclear Heat Source (ENHS) reactor core design. Nuclear Technology 149, 22. Hong, S.G., Greenspan, E., 2005. Use of minor actinides for reduced power ENHS reactor core design. In: Proceedings of GLOBAL’05, Tsukuba, Japan, October 9e13. Monti, L., Greenspan, E., Sumini, M., Fratoni, M., Rocchi, F., Multi-recycling in the ENHS. In: The Proceedings of the Second COE-INES International Symposium on Innovative Nuclear Energy Systems, INES-2, Yokohama, Japan, 26e30 November, in press. Okawa, T., Greenspan, E., 2004. Alternative ENHS core design using stepped geometry core. Transactions of American Nuclear Society 91. Okawa, T., Greenspan, E., 2006. Effect of fuel type on the attainable power of the Encapsulated Nuclear Heat Source reactor, Proceedings of the ICAPP2006, Reno, NV, 4e7 June. Sienicki, J.J., 2003. Updated thermal hydraulic analysis for recent ENHS design improvements, International Congress on Advances in Nuclear Power Plants, ICAPP 2003, Cordoba, Spain. Stewart, J.S., Lamont, A.D., Rothwell, G.S., Smith, C.F., Greenspan, E., Brown, N., Barak, A., 2002. An Economic Analysis of Generation IV Small Modular Reactors, Lawrence Livermore National Laboratory Report UCRL-ID-148437. Susplugas, A., Greenspan, E., 2005. Implementation of the ENHS Thermal Hydraulic Optimization Code for Recent ENHS Design Improvements. University of California, Department of Nuclear Engineering. Internal Report UCB-NE-5104.