CHAPTER 24 LIQUID METAL FUEL REACTOR DESIGN STUDY* 24-1. CompARISON oF Two-FLuip AND SINGLE-FLUID LMFR DeEsignNs In Chapter 18, the two-fluid and the single-fluid externally cooled LMFR concepts were discussed in a general way. It was pointed out that the two- fluid design has the better breeding possibilities but is somewhat more complex than the single-fluid reactor. In this chapter a complete design study of a two-fluid full-sized LMTR reactor is deseribed and discussed, and a shorter discussion of a single-fluid design study follows. This does not mean that one design is necessarily favored over the other. In fact both of these designs are being studied very extensively. 24-2. Two-Fruip Rracror DESIGN 24-2.1 General description. The two-fluid externally cooled LMFR concept consists of a relatively small core surrounded, for the most part, by a blanket containing fertile material. The core is composed of high-density, impervious graphite through which vertical channels are drilled to allow circulation of the fuel coolant. The fuel in the core is dissolved U233 or U233 dissolved and suspended in liquid bismuth. The fluid fuel also acts as coolant for the core system. The required coolant to moderator ratio is obtained by proper size and spacing of the fuel coolant channels. The blanket is constructed of high-density graphite through which flows a liquid bismuth slurry containing the bred U?2*3 fuel and thorium, the fertile material. In this study, thorium is assumed to be suspended in bis- muth as thortum bismuthide, although thorium oxide particles could be used. The blanket is wrapped around the core as completely as possible for good neutron economy. An important economic consideration is the degree of end blanketing which can be achieved while keeping coolant velocities below the allowable limit. Several blanket designs were in- vestigated, but a complete study for obtaining the best end blanket design has not yet been carried out. *This chapter is based on studies made by Babeock & Wilcox Company for the USAEC, BAW-1046, March 1958, and on a 17 company report BAW--2, June 30, 1955, for which Brookhaven National Laboratory contributed information and sup- plementary design studies. 866 24-2] TWO-FLUID REACTOR DESIGN 867 24-2.2 General specifications. Unless otherwise noted, the specifications listed below are common to all calculations performed in this design. Total power 825 mw (thermal) 315,000 kw (electrical) Coolant to moderator ratio in core, Vpi/Vc 1.22 Coolaut to moderator ratio in blanket, Vaury/Ve — 0.50 Core-blanket barrier material graphite Blanket thickness 3.0 ft Blanket slurry composition: Bismuth 90 w/o Thorium, as ThzBis 10 w/o C'oolant inlet temperature 750°F Coolant outlet temperature 1050°F Nuclear ealeulations utilizing latest cross sections and multigroup diffu- sion theory indicate that the values 1.22 and 0.50 listed above are close to the optimum. The =everal factors which dictated the choice of a bismuth-to-carbon volume ratio merit some attention. There are some losses of neutrons due to capture in graphite. Hence, one would wish to use only enough graphite to 0.04 0.02 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24 [ I | T I I Noa/Ng. ~ 15 x 10 4 23" 7B ] ] ] | ] ] 1 1 0.25 0.50 075 1.00 1.25 1.50 175 2.00 2.25 Bismuth te Carbon Volume Ratio F1a. 24-1. Breeding gain vs. bismuth-to-carbon volume ratio in core. Breeding Ratio 1.1 0.9 l ] L 2.0 2.5 3.0 3.5 4,0 Blanket Thickness in Ft. Fia. 24-2. Breeding vs. blanket thickness for slurry-to-carbon volume ratio = 1.00 and bismuth to carbon volume ratio in core = 1.00. Breeding Ratio i i ! [ i 1.04 0.50 0.75 1.00 1.25 1.5¢ 175 Slurry to Carbon Volume Ratio Fia. 24.3. Breeding vs. slurry-to-carbon volume ratio in blanket for bismuth to carbon volume ratio = 1.00 and blanket thickness = 3 ft. 24-2] TWO-FLUID REACTOR DESIGN 869 24-2.3 End blanket effects. A series of nuclear calculations were per- formed to determine the effects of end blanket design upon breeding ratio and eritical fuel concentration. Two extreme blanket designs were con- s1dered. In the most optimistie case, a spherieal core, equivalent to a 61-in.- diameter eylinder, was surrounded by a 3-ft spherical blanket. The pessi- mixtic caleulations assumed a cylindrical core with a diameter of 61 in., height equal to 1.5 times the diameter, a 3-ft radial blanket, and no end blanket. Critical values of fuel concentrations and breeding ratio were caleulated for four power fractions in the blanket for cach design. All caleulations were performed for hot, clean conditions with an average temperature of 900°I. A two-group, multiregion code was used to solve the diffusion equations, and a 37-group spectral code was used to determine the two-group nuclear constants. The results of these calculations are taubulated in Table 24-1. The breeding ratio is decreased 0.20 to 0.25 by completely eliminating the end blankets. This is due primarily to the added neutron leakage out the ends of the core, despite the fact that the core height 1s Inereased. Although the eritical mass of fuel in the core is higher without end blankets, the fuel concentration is somewhat lower due to the increased core volume. TasLe 24-1 CriTicaniTy Cancurations rorR Two-I'wuip LMEFR WITH AND WITHOUT IKND BLANKETS Vay/Npi X 10° Ratio of . Blanket ‘ \ Breeding : . (use ———— | blanket power ratio thickness, (ieometry -~ Core | Blanket | to total power ' ft I - 509 152 0.0665 1.053 3.0 Full blanket IT 530 034 0.205 1.051 3.0 ? ” 1 461 | 1600 0.445 1.039 3.0 7 ” IV 7 436 2100 0.515 1.033 3.0 7 " V 403 1050 0.272 0.80 3.0 No end blanket VI 366 2100 0.425 0.82 3.0 v ” VII 347 2808 0.492 0.83 3.0 v ! VIIT' 403 | 1050 0.272 — 4.0 v " The actual core and blanket design is between the two extremes assumed in these caleulations. The blanket can be extended beyond the end bound- aries of the core, and a graphite reflector can cover the ends of the core except for the coolant inlet and outlet. Cooling becomes a serious design 870 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24 0 C ] | L shielding ' Control Rod/ET ] Control Rod Drive ] /L Core Qutlet Nozzle ///// /// Blanket C])uflef T 7 S R A B R R R SRR R R R R R RS NN pSSSSE TR R R R R R R AR TN . A R T ] | Fuef Passage Blanket Inlet Nozzle Core Inlet Nozzle FiG. 24-4. Two-region, externally cooled liquid metal fuel reactor. problem, if the end reflector is replaced with blanket material. The design in Fig. 24-4 is a substantial improvement over no end blanket or reflector. However, further improvement in breeding ratio could be achieved with even better end blanket designs. 24-2.4 Power level in the blanket. For a given geometry, coolant-to- moderator ratio, and thorium concentration in the blanket, specification of the fraction of total fissions generated in the blanket establishes a unique set of values for fuel concentration in the blanket, fuel concentration in the core, and fissions generated in the core. For simplicity, the power generated in a region is assumed directly proportienal to the fissions in that region. The data in Table 24-1 indicate that breeding ratio changes very little with large changes in the fraction of total power generated in the blanket. This increase in blanket power results in an Increased ratio of resonance to thermal absorptions, a phenomenum which tends to offset the additional fast neutron leakage out of the blanket as blanket power increases. 24-21 TWO-FLUID REACTOR DESIGN 871 An economic analysis of the effects of changing the blanket power frac- tion was performed to determine the optimum core-blanket power split under equilibrium operating conditions. The parameters affecting this choice are (1) fission-product poison levels in the blanket, (2) fission- product poison levels in the core, and (3) chemical processing costs, Frission-product poisons in the blanket. The chemical processing of the blanket slurry accomplishes two things: (1) The removal of bred U?3? from the blanket system at a rate necessary to maintain the 17232 concentration in the blanket slurry at some equilibrium value corresponding to the desired blanket power fraction. (2) The removal of fission products from the blanket slurry. If the blanket processing cycle 1s determined by the minimum removal rate of 17238 for steady-state operation, a corresponding poison level in the blanket 1= automatically set. If the blanket chemical processing cyele 1= determined by the poison level and is less than the cycle determined by the above criteria, the bred fuel removed from the blanket must be fed hack into both core and blanket to maintain steady-state fuel concentra- tions. In this analysis the blanket processing eyele i all eases was assumed to be based on the minimum removal rate to maintain steady-state U233 concentrations without feeding fuel into the blanket system. (e mical processing cycle for blanket slwrry. The chemical processing was a==umed to be performed continuously on the reactor site. Unless other- wise specified, the fluoride volatility process is utilized as described in Article 24-3.16. The chemical processing cycle for the blanket may be caleulated 3] from the equation _ ZW Mg [V 4 (Z1a/ Z.) (0 a)] s s — (07 T'p = blanket processing cycle, days, Tg where Z, = removal efhciency for uranium = 0.25, Z13 == removal efficiency for protactinium = 0.04, M, = mass of fuel in blanket system, kg, h, a = ratio of Pa®33 to U233 in blanket, 3 = kg of fuel burned per Mwd = 1.05(1 4+ «23), P, = total power, 825 Mw, BR = breeding ratio, Pg = blanket power, Mw, 872 LIQUID METAL FULL REACTOR DESIGN STUDY [cHAP. 24 and Z Tg ¥ 0% (efh)d™ + Y13 Qe where 0% (eff) = an effective absorption cross section to account for resonance and thermal absorption in 17233 3 o5 = average thermal flux over the blanket system, Y13 = decay constant for Pa®33, The poison level in the blanket depends upon T, and Tg 1= a function of M J, b/a, breeding ratio, and power fraction in the blanket. All these arviables are interrelated. The ratio b,/a 1s a function of Ty, but Tpis a slowly varying function of b/a due to the low value of Z,3/Z, (0.16). Breeding ratio is a slowly varying function of fission-product levels in the blanket due to the heavy loading of fuel and thorium in that regiow. The breeding ratio is sensitive to the poison level, and thus to the chemical processing rate, in the core fuel solution. An iterative calculation procedure was required to arrive at optimum values of T, fission-product poison level in the blanket, and the power fraction in the blanket. For a given chemical processing rate in the blanket, the fission-product poison level was determined from the data in KAPL 1226 [4]. Relative poisoning, RP, is defined as the absorptions in fission products per thermal fission in fuel, while the fission-product poison fraction is the absorptions in fission products per total absorption in fuel. Xenon and samarium are treated separately and are not included in the term fission products. The burnup, F, in a region is defined as the atoms of fuel fissioned per atom present in the region. The burnup F at time T in the blanket is caleulated from . 0.866 T(Pg/1IP) F= - ;B Mas Using this relation, the relative polsoning in the blanket was determined for ench processing cycle from a graph of R versus I7 [4]. The RP curve used is based upon high cross sections of all fission products with the excep- tion of u low value for Zr¥s, Xenon in the blanket. Xenon is removed from the blanket by the degasser. Although the removal rate of fission-product gases cannot be determined until experimental information becomes available, a poison fraction of 0.01 was assumed for Xe!33, 24-2] TWO-FLUID REACTOR DESIGN 873 Samartum n the blanket, The removal rate of samarium by chemical processing wus neglected. The steady-state ratio of Z50/Z3 using ap- propriate thermal absorption cross sections, is determined by the relation 8 5 Som= 1.42 X 10716¢ + 0.0126, where ¢ = average thermal flux in the region of interest. Fisston-product poisons in the core. The level of fission products, FP, other than xenon and samarium, in the core 1s determined by the chemical processing cyele for the core fuel solution. The steady-state value of FP poisons in the core should be established by an economie balance between the value of improved breeding ratio and inereased chemical processing costs, The relationship between the core processing cycle, T, and the rela- tive poizon, RP, in the core may be expressed as d(RP) _RP dF ~ F and 0.866 T.(P./P) F=—— My, where D d(;;‘}') 1s the slope of the curve RP versus F [4], Mgy = total mass of U23? in the core system. The xenon and samarium poisons in the core are determined as described for the blanket. [eonomie optimization. An optimization study was performed to de- ternune the most economic power split between core and blanket systems qind fission-product poison level for the core during equilibrium operation. The tuel cost items which vary with these two parameters are (1) bismuth Heventory. (29 fuel inventory, (3) fuel burnup, (4) thorium amortization, 5 thortum burnup, and (6) chemical processing. Nuclear calculations specitied the fuel concentrations for both core and blanket and breeding ritin=. These values were then used to determine the chemical processing cvele for the blanket and the pertinent costs. 874 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP, 24 2800 2600 - 2400 — 2200 | 2000 - 1800 1600 |— Ngq/Np; in Blanket x 10® 5 ~ = g8 8 8 | o o L o o - 4 £ w — — 2 U ) _ & - 100 ] 5 -— 3 [— T ~—RP = 03— 501 | ] I ] I ] I L1 ] Q.10 20 30 .40 .50 .60 Pg/Py Fig. 24-10. Chemical processing cycles vs, blanket power, based on a blanketed sphere with total reactor power of 825 Mw and the removal efficiencies of Z,, = 0.25, Zy3=0.04, ZB = 0.10, ZEp = 1.00. 5x]06 T T T 1 T T 1 T T 1 T T T 1 B Total Annual Cost - ‘“_: | o "; 1061 Capital Equipment ] £ — - g — i 3 — 7 ° - . ] s 6 — Operating Cost 3 T [— - o — £ _ c L - i I _ — | Building Cost 10° ! [ bt P Lt bl ! 1 3 6 10 30 60 100 200 Plant Liquid Throughput, Cu Ft Per Day Fic. 24-11. Annual fluoride volatility processing cost vs. plant throughput for 825-Mw-two-fluid LMFR. 24- 2] TWO-FLUID REACTOR DESIGN 881 6 810 T T T T T T T T T T T T 6 |- e 4 rs A - P - S i Total Annual Cost /’ 7 D 2 - ] £ - gl | § Capital investment Return a 5 - 3 10 N e — E s | -7 Operating Cost j a4 L B o 6 -7 — AX]OS 1 I L I L 1 L IJIllllJlli 1 I L 1 L J 1 llll]lll]l 1 3 6 10 30 60 100 Plant Throughput, kg Therium Per Day Fie. 24-12. Annual aqueous processing costs vs. plant throughput for 825-Mw- two-fluid LMFR. ("hemical processing costs. The chemical processing cycle time for the blanket 1= determined by the Pg/I’ ratio and the breeding ratio, as dis- cu==ed m previous paragraphs. The processing rate for the core system is determined by the method also deseribed previously; see Fig. 24-10. The totul throughput to the fluoride volatility chemical separations plant is ~imply: Vcs Vbs T T Ty Throughput (ft3/day) = The wnnmual processing charges based on fluoride volatility can be read directly trom Iig. 24-11, a plot of annual charges versus plant through- put. A=~ & matter of comparison, the chemical processing charges were also computed for each case, assuming on-site aqueous processing methods. The capacity and eost of an aqueous processing plant are determined by the wnmount of thorium per day which must be processed. The core solu- tion provessing does not enter into the cost unless the ratio of fuel to thorium present= criticality problems in the process equipment., This situation is likely 1o oceur for the higher power levels in the blanket. This analysis did not take this possibility into account, however, and annual aqueous processing costs were taken directly from Fig. 24-12. This design plant capacity 1= 33 kg day of thorium feed. Results of optimization. 'The bismuth inventory is slightly greater for the case of I’g I’; = 0.10 than for the other two cases, because of the added primary system volume. Fuel inventory charges are not very sensitive to 882 LIQUID .40 METAL FUEL REACTOR DESIGN STUDY 36} 32} 28— 24— 20— Relative Fuel Costs, Mills/KWH -.04+ -.08 | l l l l Thorium Charges RP=.157] RP= .09 RP=.037] Fuel inventory ; Bi Invemory/ RP = .15 A0 .20 .30 .40 .50 .60 Pg/Py [cHAP. 24 Fig. 24-13. Relative fuel costs vs. blanket power for two-fluid LMFR based on a fully blanketed sphere operating at 825-Mw with a plant factor of 809,. 2.00 1.801 1.601— 1.40— 1.201— 1.00H— Chemical Processing costs, Mills/KWH 0.80 |— 0.60— 0.40 Aqueous On-Site Fluoride Volatility RP=0.09 RP=0.15 0.10 .20 .30 A0 .50 .60 Pg/Py I'1g. 24-14. Chemical processing costs vs. blanket power. The cycle times are based on a blanketed spherical reactor with a total heat power of 825-Mw. 24-2] TWO-FLUID REACTOR DESIGN 883 2.4 | I l 1 l 2.2 — 20 | — = X " 28 | ] 7 S On-Site Aqueous Tg 16 Lo Processing B [V .qaa RP =.03 I @ 14— RP =.15. RP = .09 RP =15 Fluoride RP = .09 12 — Vo|a?i|i.ry RP =.03 Processing o | 1 1 | 1 0 10 .20 .30 .40 50 .60 PB/Pi Fi6. 24-15. Relative fuel costs vs. blanket power for a blanketed spherical re- actor operating at a total power of 825 Mw with a plant factor of 809. the relative poison level in the core, but they increase sharply with an in- crease In power level (Fig. 24-13). Thorium charges increase linearly with blanket svstem slurry volume, and fuel burnup charges increase as Pg/P, increases, as shown in Fig, 24-13. Chemical processing costs drop rapidly as the power fraction in the blanket increases. The increased processing rate required to maintain a stendy-state fission-product relative poison level in the core of 0.03 results i u processing cost much higher than required for RP values greater than 0.09. The aqueous processing costs appear to become essentially equal to fluoride volatility costs at a value of 509, for Pg/P,. Further analysis would be required to determine the validity of the aqueous processing cost curve for low throughput and high N23/Np2 ratios encountered in the cases of high blanket power. The chemical processing costs are tabulated in Table 24—4 and shown graphically in Fig. 24-14, The results of the economic comparisons are summarized in Table 24-5 and are graphed in Fig. 24-15. (RP on the graphs refers to the relative poison level of the fission products in the core.) Figure 24-15 shows that for all values of RP a minimum fuel cost occurs for a Pg/P; of approxi- mately 0.33. 24-2.5 Selection of a reference design. The optimization study indi- cated that the most economic reactor design should produce one-third of TaABLE 244 Cuemicar, ProcessiNng Costs Two-Fruip LMFR Slurry U-Bi flow flow to C.hem— Fluoride | Fluoride Th/day Blanket | rate to : ical . L to Aqueous | Aqueous i chemi- volatil- | volatil- . o process | chemi- | RP T, plant |. - __4+ | chemi- | process- | process- Case | PB/P: ! cal 1ty costs, ity costs,| TB(Z,=1) . : cycle, cal |(core)| days thru- | 7, . cal ing cost, | ing cost, plant, $/yrx | mills/ b days | plant, ft2/day put, 10-6 kwh plant, [$/yrx10" 6mills/kwh ft‘g/d&y, Vcs/T ’ ft3/day, M()Q/ TB Vos/ T ‘ I1(a)0.10 16.26 | 45.8 0.03 75. 25.3 71.1 3.03 1.37 65 325 4 .20 1.90 (b) 16.59 | 44.9 | 0.09 | 557 3.44 | 48.3 2.57 1.17 66 320 4.15 1.88 (¢) 17.08 | 43.6 | 0.15 | 1116 1.71 | 45.3 2.51 1.14 68 310 4.10 1.86 Il (a) |0.3333] 122.3 8.05 | 0.03 60 20.9 29.0 2.15 0.974 489 57.1 2.10 0.951 (b) 129 7.64 | 0.09 | 446 2.81 | 10.5 1.65 0.747 516 54 1 2.08 0.937 (c) 136.3 7.23 1 0.15 1 900 1.39 8.62 1.58 0.716 525 53.1 2.06 0.933 11T (a) { 0. 50 444 5 2.77 10.03 56.5| 17.9 20.7 1.95 0.883 1778 19.6 1.38 0.625 (b) 513 2.40 | 0.09 | 432 2.34 1 4.74 1.40 0.634 2052 17.0 1.32 0.598 (¢) 547 2.25 |1 0.15 | 854 1.18 3.43 1.33 0.602 2188 15.9 1.30 0.589 788 AdALS NOISHd YOLOVAY THAd TVIHW dindIl ¥z "dVHD] TaprLem 21 D Revartive IFven Cost vor Two-'Lumn ENMEFR (WrtH PorsoNs-BLANKETED SprisRy) Bismuth Fuel Thorium Fluoride | Total costs inel, A ueoUs Total costs incl. il;\'en— im’eil- Fuel ven- Thorium | volatility fluoride vol. : ;roc - AQUeOUS Proc- y ' ' burnup burnup, | process- pProc. ' essing tase | ’g/P tory tory L1 tory, ol Case Bre C :Ifi)’,, 3l o >(<)Il‘\0’”3 C3x10°3 o 21]}0,3 CsX1073 | ing costs, |~ - o (;HI(A)L'? e | S S| S/yr | Cpx1078 Cix1073 ) mills R0 00 x1078 ], mills T T o $ar v “'kwh o $/yr “kwh T(a);0.10 440 353 —116.5 133 12.3 3031 3852 1.75 4200 5022 2.28 (b) 440 303 —31.4 133 12.1 2570 3477 1.58 4150 5057 2.29 {¢) 440 369 98 .6 133 11.8 2510 3562 1.61 4100 5152 2.33 IT (a) | 0.3333 7 399 —31 .4 176 12.1 2150 3077 1.39 2100 3027 1.37 (h) 71 407 31.4 176 11.9 1650 2647 1.20 2070 3067 1.39 (¢) 371 429 98. 6 176 11.8 1580 2666 1.21 2060 3146 1.44 III (a) | 0.50 371 678 89,8 220 11.8 1950 3311 1.50 1380 2741 1.25 (b) 371 731 184.5 220 11.6 1400 2018 1.32 1320 2838 1.29 (¢) 371 766 247 220 11.4 1330 2945 1.33 1300 2915 1.32 NDISEd MOLOVAY dINTId-OML [Z-%2 G8% 8806 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24 the total power in the blanket system and that the relative poison in the core due to fission products should be approximately 0.09. However, sev- eral effects must be considered in relating the optimum reactor to the ae- tual operating reactor. A geometry more realistic than the fully blanketed sphere must be considered in establishing new specifications; effects of higher uranium isotopes, Pa losses, and control rods on breeding ratio must be taken into account; and a new chemical processing cycle for the blanket, along with a new fission-product poison level in the blanket, must be cal- culated based upon the adjusted breeding ratio. Geometry effects. The inability to wrap a blunket around the ends of the core requires an adjustment to the parameters for the reference design hased on the caleulations with a full blanket. The axial leakage out of a bare ended core and a blanket with a height 1.5 times its diameter was cal- culated to be 0.18 neutron per absorption in fuel. An extension of the blanket length and the addition of partiat end graphite reflectors are esti- mated to reduce the end leakage to one-half this value. The total neutron leakage, both fast und thermal, out of the partially blanketed reactor is estimated at 0.17 neutron per absorption in fuel. The added length of core and blanket will slightly increase the critical mass, but the required Nas/Npi ratio will decrease slightly. In order to be conservative in the fuel inventory costs, however, the critical values of Nas/Nyg; for the fully blanketed sphere are assumed for both core and blanket. Breeding ratio. Higher uranium isotopes. The higher uranium isotopes, primarily U283 U235 and U2, continue to build up in both the core and blanket fuels throughout reactor life, since they cannot be separated in the _chemical plant. The relative poison due to these isotopes, however, rises rapidly at first with the buildup of U2 but increases very slowly there- after. The return from U235 fissions almost balances for losses to U#3% and U236 [4]. An average poison fraction of 0.01 for the reactor is used for the reference design. ‘ Protactinium losses. The equilibrium Pa233 concentration can be com- puted from the relationship N =N, using an effective thermal absorption cross section of Pa**3 based on the calculated neutron spectrum in the blanket. The relative absorptions of the Pa233 are very small (0.005), but they are included. Control rods. The self-regulating properties of an LMFR have not been established at this time. An allowance of 0.01 in relative absorptions is included to account for the possibility of using a regulating rod and a small 24-2] TWO-FLUID REACTOR DESIGN REFERENCE DESIGN SPECIFICATIONS SPECIFICATIONS FOR HQUILIBRIUM OPERATION Core: Thermal power Flectric power Diameter, inches Height, inches Fuel pi/Ve Noy/Npi Mass of U233 in system, kg Total volume of fuel, ft3 Breeding ratio, over-all Chemical processing cycle, days Volume flow rate through chemical plant, ft3/day Mas= flow rate through chemical plant, g 17233 /day Average thermal flux in active core Average thermal flux in core system Blanlt: Thermal power Flectrie power Thickness, ft [ I'C .\hmy content: Thorinm (as ThsBis) 10% wt Bi~muth 90% wt Noy N (atom I‘thiO) 1190 x 10~¢ Miss of U299 1n system, kg Mis= of thorinm in system, kg Total volume of fuel, ft? Chemieal processing eyele, days Volume flow rate through chemical plant, {t%/day Mas= flow rate through chemical plant, kg of Th/day 887 550 Mw 210,000 kw 61 91.5 233 1.22 600 X 1076 234 1255 0.86 446 2.81 525 1.6 X 1019 6.4 X 1018 275 Mw 105,000 kw 3 0.5 328 27,900 985 200 4.91 140 amount of =shim control for normal operation. Safety rods are included in the reference design but do not affect neutron economy. Fission-praduct potsons. The adjustment of breeding ratio to correspond to the effect= outlined above changes the required chemical processing cyele for the blanket system. This change in T'g also changes the equilibrium 88K LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24 value of fission products in the blanket. Proper adjustments result in a blanket processing cycles of 200 days (assuming Z, = 0.25) and a fission- product poison fraction in the blanket of 0.039 (RP in blanket = 0.15). Neutron balance. The neutron losses proportional to one absorption in U233 are listed below: Absorptions in: U233 1.000 Th 0.860 C 0.025 Bi 0.050 Xeld 0.010 Sm 149 0.017 Figsion products 0.073 Higher isotopes 0.010 Control rod 0.010 Pa233 0.005 Leakage 0.170 Total 2.230 214-3. SysTEMSs DESIGN 24~-3.1 General. Systems design covers all of the reactor plant external to the reactor, except for chemical processing. The reactor plant includes the steam generator, but not the steam system or its auxiliaries. The principal purpose of the systems is to transport heat from the reactor and generate steam. They also provide supporting functions, such as shield cooling, uranium addition, ete. The primary system consists of six heat transport loops, each consisting of a pump, a heat exchanger, check valve, and interconnecting piping. The hot-leg temperature is 1050°F; the cold-leg temperature 750°F. In each of the intermediate heat exchangers, heat 1s transferred from the bismuth to the intermediate fluid, sodium. There are six intermediate heat transport loops, each containing a pump, steam generator, and interconnecting piping. The hot-leg temperature is 1010°F; the cold-leg temperature 680°F. Steam is produced at 2100 psia, 1000°F, Selection of the above parameters was a problem involving consideration of the steam plant as well as the reactor plant. The primary system temperatures were first fixed by using the largest AT considered likely to prove practical. The temperature approach of the intermediate heat exchanger was set at 40°F, resulting in a sodium hot-leg temperature of 1010°F. To provide the close approach necessary for steam temperature stability, the steam temperature was set at 1000°F. A steam pressure of 2100 psig was picked to correspond with 1000°F. 24-3] SYSTEMS DESIGN 889 Shifting the sodium cold-leg temperature redistributes heat-transfer surface between the intermediate heat exchanger and the steam generator. However, it seems desirable to favor making the intermediate heat ex- changer small to cut down on fuel inventory. For this reason, the sodium cold-leg temperature was established at 680°F, 24~-3.2 Plant arrangement. Plant arrangement starts with positioning the primary system relative to the reactor, and this is determined by seven principal considerations: (1) reactor design, (2) plant operation, (3) main- tenance, (4) operational limitations of major components, (5) structural integrity of piping, (6) economies, and (7) safety. A preliminary analysis of the two reactor concepts, single-fluid and two-fluid, resulted in the decision to use three external loops for the single- fluid and six for the two-fluid reactor. For both these alternates the main- tenance philosophy selected was that of removal and replacement by hori- zontal transfer of a complete primary loop upon failure of any major component in the loop [5]. Thus, for arrangement purposes, the primary loop= assume the shape of a rail-mounted horizontal containment vessel, or capsule, sized to contain all loop components. The height of the capsules relative to the reactor is dictated by an economic balance between height or elevation costs and pump net positive suction head. The arrangement for the two-fluid reactor with six primary loops is shown in Figs, 24-16 and 24-17. [ plan, the primary loops were located radially around the reactor, Fig. 24-16. A minimum length of interconnecting pipe between the reactor and the loops was used because of high fuel inventory costs. This latter con=ideration ruled out shielding of any appreciable thickness between the reactor and the loops. Maintenance access doors and other shielding aronnd the outside of the loops was sized for source conditions 6 to 8 hr after shutdown of the reactor to permit access by maintenance personnel at that time into the annular area. With the primary loop arrangement established, the next problem was locution of the intermediate system. Since this system is the connecting link between the primary systems and the steam turbines, it must be locuted between them. The turbine is above ground level for gravity drain- aue of condenser cooling water, and the primary loops are below ground level for economy of shield costs. The path taken by the intermediate svstemn can be either a high-level path, immediately up from the primary svstem, or a low-level path, immediately down from the primary system, and then horizontally to an area outside the primary system area. The intermediate system in this arrangement follows the high-level route to the steam plant. Sodium lines are brought straight up to an annular area around the reactor maintenance chamber. Since access to 890 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24 ! Turbo-Generator Building Sodium Cold Dump Component __Tanks 25200 Sterage ; N, - } l bhl,Jp!flq i . s | i . Foacs \ \.) \ | and : -M G,Q o achine | | |Receiving] 3 e LN \ :\T"__“_'" : Pool \ ) '/Fi ‘ © 7 b i r‘\m\ ! I—_ Platform apsule: I\ L o L. Decontamination % ) J‘.LU___.__ Area \T:,Zj,: - jP|utform ‘@ Blowers @ and /‘ Venhlclhng Stack system Pyroprocessing and Aqueous Capsule Removal and Equipment Storage First Floor Plan Fra. 24-16. LMFR-6: Capsulate loop conceptual plant layout. this chamber will not be permitted during reactor operation, a heavy shield wall i1s not required around the chamber. Within this annulus are the sodium pumps and the steam generators. Final layout of this equipment will require considerable ingenuity, but it is feasible. Steam lines will cross the roof of the reactor building to the turbine building, Because the primary loop hot maintenance shop for this concept serves such specialized functions, its usefulness for maintenance of chemical processing equipment is doubtful. Accordingly, the chemical processing facilities for this two-fluid six-loop plant, together with its supporting hot and conventional laboratories, fuel addition and other systems, are located in a separate building. The turbine building 1s of conventional construction and will be in all essential respects wdentical for both plants. Startup heating switch gear, gas heating and cooling systems for the reactor and dump tanks, inert gas storage systems, control rooms, and other auxiliaries are located relative to the above systems as logically as possible in the light of their functional requirements. With respect to contamination control the basic philosophy 1s (1) con- trolled access to areas having different order of magnitude activity levels and (2) controlled circulation of ventilating air to assure flow from low- 24-3] SYSTEMS DESIGN 891 Reactor Service 5 Ton Crane Capsule Area Sodium Acc:ss - Sodium rea —k Systems S Systems Shielding Windows Mechanical Arm Rails Electr'y'c Core Reduct'n Liquid Return F1a. 24-18. Fluoride volatility processing of core and blanket. 24-3.15 Maintenance. The maintenance of the reactor and primary system components will be completely remote, because of the high levels of radioactivity of the circulating fuel. The entire plant and reactor system are arranged for remote maintenance [5}]. 24-3.16 Chemical processing. The pyro process chosen for this economie study 1s the fluoride volatility method applied to a two-region reactor. Work of adapting this process to bismuth fuel processing is presently under way at Argonne National Laboratory. Figure 24-18 presents the main steps 1n this process. As shown, the process may be used for either blanket or core liquid. When the plant i1s processing core liquid the basic steps in this process are (1) hydrofluorination to oxidize uranium and some fission products, (2) transfer of the oxidized material to a fused salt phase, (3) Auorination of the salt carrying the uranium and fission products for sepa- ration of uranium as volatile UFg, (4) reduction of the Uly to UF4 by H. in o fused salt phase, and (5) reduction of UF4 to uranium metal and transfer into the metal phase (bismuth). The volatility method can be conveniently used to process a thorium bizmuthide blanket. The process must be preceded by a phase separation ~tep which separates the thorium bismuthide solids from the liquid carrier hi=muth (Fig. 24-19). The modification of the core liquid process flowsheet v ax follows: (1) salt effluent from the hydrofluorination step must be stored in order to achieve Pa decay to uranium, and (2) the bismuth liquad 15 returned to the blanket head end process without the addition of uranium. 898 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24 Stripped Bi l From Volatility Plant Th Slurry Mixer Feed Sluery 1 Recycle L___>. Heat Pulser Crystallizer b Phcse Separator Slurry Return Y - To Blanket U Rich Bi Liquid F1g. 24-19. Head end processing, bismuthide slurry. Certain of the fission products are not removed by volatility processing. These may be removed by zine precipitation (Fig. 24-20). This process requires that the bismuth feed be free of uranium, and the volatility plant provides such a bismuth feed stream. The head end process transfers bred uranium, protactinium, and fission products out of the solid phase portion of the slurry and into the liquid phase. After this step the two phases are partially separated. A liquid portion transferred to the volatility plant carries bred uranium, protac- tinium, and fission products with it for stripping with HF. The stripped liquid bismuth is returned to the head end plant for mixing with fresh slurry feed. The head end process is not 1009% efficient; i.e., the uranium and protactinium are not completely removed from the slurry before reconsti- tution and return to the blanket region. This problem has been examined in some detail and was taken into account in determining economics. 24-3.17 Turbine generator plant. A flow of 3,330,000 1b/hr of super- heated steam at 2100 psi and 1000°F is delivered to the turbine. The generator has a gross output of 333,000 electrical kw, and the condenser removes 1.677 X 10° Btu/hr at 1.5 in. of mercury absolute, thus giving a gross heat rate of 8450 Btu/kwh. About 18,000 kw of electrical power is used for the various pumps and auxiliary systems in the plant, making the net output 315,000 kw. Therefore, the net heat .ate is 8940 Btu/kwh, which corresponds to an efliciency of 38.2%. 24-3] SYSTEMS DESIGN 899 From Zn Rich, NFPN Volatility ‘ TPlus Bi Plant Zinc Concen— . _Bi Rich | trator Zinc Plus Zn Waste Zn In Zinc Zn Crystal - lizer | Still Bi Return To Volatility Plant Fia. 24-20., NFPN fission-product removal. At full load there are 1,825,500 Ib/hr of steam leaving the turbine and being condensed in the condenser. Also, 113,800 Ib/hr of water from various parts of the cyele are being cooled by the condenser. The total head load on the eondenser is 1.677 X 10 Btu/hr. The condenser cooling water enters one water box at 70°F and leaves the other at 80°F. 24-3.18 Off-gas system. The actual design and efficiency of any con- ceptual degasser are as yet unknown quantities, and a knowledge of these important details will have to wait until in-pile loops have provided suffi- cient data. The off-gas system will consist of a cooler followed by a series of storage hottles. Gaseous fission products that have been separated from the Liquid bizmuth in the degasser are first sent through a cooler which offers a resi- denee time of about a day, or enough for most of the short-lived isotopes to decay. From the cooler, the gasses are compressed into storage bottles, each capable of holding 30 days’ accumulation. The storage bottles will cach be .25 {t3 in volume, and at 212°F and 60 psia at the time of dis- cotnection from the compressor. ~ome sweep gas may be included in the above gas stream, but the present dexign philosophy indicates that no extra sweep gas should be required; however, if some sweep gas 1s required for efficient degasser operation, this gas could be obtained by a recycle of previously removed gas. This recycle ~weep stream would most probably be taken from the storage bottles after ~ufficient cooling. 900 LIQUID METAL FUEL REACTOR DESIGN STUDY [cuap. 24 The gas in the storage bottles may be vented to the atmosphere after 90 days of storage, since then only the Kr35 activity is still present in ap- preciable amounts, and this can be released provided there is sufficient di- lution. However, the most probable course of action will be to process the stored off-gas through a gas separation system, where the valuable Kr8? will be recovered. 24-4. SineLE-Fruip ReacTor DESIGN 24—4.1 General description. The single-fluid LMFR concept has been investigated to determine the characteristies and economie attractiveness of this design. In general, the core consists of a large array of solid modera- tor blocks stacked to provide the desirable geometry of a cylinder. Vertical cylindrical channels are drilled through the moderator to allow circulation of the liquid metal slurry containing both the fuel and fertile material. The fission heat generated in the fuel-coolant stream is transported by forced convection to heat exchangers external to the reactor vessel. The unique feature of this concept is that only one coolant, the slurry, is used for removing heat from all parts of the reactor. The desired slurry-to- moderator ratio is achieved by selecting the appropriate combination of channel size and spacing. 24—4.2 General specifications. The general specifications for the system affecting reactor design are tabulated below: Power 825 Mw (thermal) 315 Mw (electrical) Slurry temperature: T; 750°F Tout 1050°F Maximum slurry velocity 10 fps Fuel J235 op {7233 Fertile material Thorium Moderator material (iraphite or BeO Slurry carrier Bismuth or lead Slurry-to-moderator ratio Variable Fertile material content in slurry Variable Core geometry Cylinder Core size Variable 24-4] SINGLE-FLUID REACTOR DESIGN 901 24-4.3 Parametric study. A parametric study was performed to deter- mine the optimum nuclear parameters for a single-fluid concept. The variable parameters investigated and their range of values are: Slurry-to-graphite ratio, V,/V. = 0.05 to 1.0, Fertile material content, g/kg of Bi= 0 to 80, Equivalent bare reactor diameter, D, ft = 10 to 20. The choice of fuel for the first full-scale LMFR will depend upon the availability of U238, which is much more attractive than U2*% because of hetter neutron economy, and a sufficient quantity for fueling an LMFR may be available in 10 to 15 yr. In the carly stages of this study, how- over, U2% was arbitrarily chosen as the fuel for the parametric study. The =election of the reference design should be valid for either fuel. In each case the critical concentration and conversion ratio were de- termined by multigroup diffusion theory, using 37 neutron energy groups. To handle the large number of caleulations, a digital computer was used once the range of values for the parameters was established by a series of criticality calculations by hand. The use of BeO as a moderator has the advantage of reducing the core A 8 S 20— — Q o | i | | | | | I 0 10 20 30 40 50 60 70 80 90 Reflector Thickness, cm Fig. 24-26. Reflector savings vs. reflector thickness for a single-fluid LMFR. These data are obtained from case 11435, where V,V,.==0.3, Wz =50 g/kg, and Dp=17. TasLe 24-7 EstiMaTED SiNGLE-FLuip Reacror Cost i : : Total 201 Reactor Graphite | Misc. | Irection| Total c(())si $/yr 10 | 160,000 | 350,000 | 167,000 | 24,000 | 701,000 | 1,000,440 | 151,400 14 | 380,000 | 970,000 | 167,000 | 30,000 | 1,547,000 | 2,227,680 | 334,152 171 570,000 | 1,700,000 | 167,000 | 35,000 | 2,472,000 | 3,559,680 | 533,052 20| 900,000 | 2,800,000 | 167,000 | 40,000 | 3,907,000 | 5,627,080 | 843,912 Bismulh inventory charges. The bismuth inventory is determined by the primary system volume external to the reactor vessel, the volume of bis- muth in the core, the volume of bismuth external to the core but inside the reactor vessel, and the holdup external to the reactor system. The primary svstem external to the reactor vessel is made up of three heat-exchanger loops containing a total volume of 1640 ft3. The volume of bismuth in the core 1s V/Ve , ey wh 7= ol VT, where § core volume Va=V The volume of bismuth external to the core and inside the reactor vessel i~ tiuhulated in Table 24-8. No additional holdup is included to account for temperature expansion during startup, fuel feed system, and other sources of bismuth inventory. The assumption used throughout this study that the volume of bismuth is cqual to the volume of slurry accounts for an additional 3 to 10% excess bismuth due to the ThO2 content of the slurry. 908 LIQUID METAL FUEL REACTOR DESIGN STUDY fcaap. 24 TaBLE 24-8 BismurH INVvENTORY IN REACTOR VESSEL ExTERNAL TO CORE Core diameter, ft Bi inventory, ft3 10 550 14 600 17 650 20 700 The density of bismuth is taken as 9.83 g/ce, and the price is assumed to be $2.25/lb. Bismuth is a nondepreciating capital investment with a 12% annual amortization rate. The annual bismuth inventory charges may be represented by the equation . -~ V/Ve Cl(fb/yr) = 012(220) V., 'H__V/I/T + Vp PRBi, where V» = total primary system volume except core, ft3, pri = density of bismuth, Ib/ft3. Fuel inventory charges. The annual lease charges on the UZ23% are as- sumed to be 4%. Treating Pa23? as fuel, the annual fuel inventory charges can be expressed as C2($/yr) = 0.04VosM o5+ VasMaz + VisMys, where Vaz = Viz = value of U??? as fuel, Vo5 = value of U%3% as fuel, $17,760/kg, M ; = average mass of element j in entire reactor system during life of plant. To simplify the work in the absence of information concerning average values of fuel mass, the total mass of fuel was considered to be the hot, clean critical loading at startup. The value of M2s is taken as the initial value with M3 and M3 taken as zero. 24-4] SINGLE-FLUID REACTOR DESIGN 909 Fuel burnup costs. Using U235 as fuel, the yearly burnup costs are C3($/yr) = 17.76(202)P3(1 — CR), where P = power, 825 My, B = grams of fuel burned per MwD), 1.25, CR = average conversion ratio. The initial value of the conversion ratio is used, since only relative costs are needed. Thortwm burnup costs. Thorium is periodically replenished in the reactor to maintain the desired concentration in the slurry. The thorium burnup costz may be expressed as C4($/yr) = V2 PBCR(292), where Vo2 = value of thorium, $12/kg. These costs are very small, ap- proximately $10,000/yr, and are neglected. (‘hemical processing costs. The chemical processing is assumed to use solvent extraction aqueous chemistry in a central processing plant. The irradiated fuel is removed from the reactor on a batch processing cycle. The processing costs are represented by Cp = 292 [95.875 “5“— 14705 MQ;}ET) 4 w 4 596] : where My2 = total thorium inventory kg, M5(T) = Moz + Mz at time T after loading of fuel charge, kg, T = chemical processing cycle time, days. Results of economic optimization. Since chemical processing costs are very sensitive to the chemical processing cycle time and the optimum cycle time may vary with reactor design, the relative energy cost of each reactor design was determined neglecting the chemical processing costs. The results of this study are tabulated in Table 24-9 and are shown graphically in Figs, 24-27 and 21-28. The pure burner, Wo2 = (0, shows costs more than twice as high as several of the more attractive concepts (Fig. 24-28). In general, the minimum 910 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24 28 ' ‘ ! T ' Vs/Ve=0.5 5 ¢—=\U. 26F T D=20 - ———D=17 / 24, ——— b=l4 Vs/Ve=0.5 \ e o VsIVe=0.5 2.2 p @ : viof 7 2 \ Vs/Ve=0.3 18| : - \\ Vs/Ve=0.2 et NN ; Vs/Ve=03 ' h : -Vs/Ve=0.3 5 ST~ <\\_,, T - -'_/_:// Vs/Ve=0.2 1.4 T ;;:"_2*}/‘5’—1?* o Ys/Ve=0.2 12 F 1.0 1 ] 1 . | ! | | 0 10 20 30 40 50 60 70 80 90 Wg2., Grams Thorium/KG Bismuth Fic. 24-27. Relative cost vs. thorium concentration for a single-fluid LMFR. 3.0 T | T r T I T ! T 1’ T [ l ! 55 |- _—— — — Wp2=0| 26~ -—- D=20 / ] D =17 s 24 - D =14 /W2 =80 — % D=1 s Wg2 =80 822 ’ RO / ] o . /«/./ @ // S > 20 +— . s ] 3 \ e ’ 18 e /"4/ W 0 W =30 ™ \ o =157 02 =38~ "027 16 - el /\‘xoz“ 0] \ = - ,,,:/,\_4';'_"'7 Ry 02 =50 14 TR _ <