.;m‘\ | ORNL —eBEORET- MASTER CCFPY FOR INTERNAL USE ONLY :f ORNL & i - ] Central Files Number ; ! 56-8-208 OAK RIDGE SCHOOL OF REACTOR TECHNOLOGY REACTOR DESIGN AND FEASIBILITY STUDY “> R This document has been reviewed and is determined to be ,0 MW FUSE D SA LT H OMOGE NE OUS APPROVED FOR PUBLIC RELEASE. Name/Title: Leesa Laymance/ORNL TIO REACTOR POWER PLANT Date: _10/04/2018 g‘;‘a‘fi;“* gfiiig‘“ fi By magctoriyg 1. ; -'*’i 5 ;g‘#-;? /' -~ o~ 5/ -2 S /3 R G e S8k inld L[ RILUTZE T e Fu: B ¥ Vo, Ve Lederzimy Motords Haskaw GREL NOTICE This document contains information of a preliminary nature and was prepared primarily for internal use at the Oak Ridge National Laboratory. |t is subject to revision or correction and therefore does not represent a final report. OAK RIDGE NATIONAL LABORATORY OPERATED BY UNION CARBIDE NUCLEAR COMPANY A Division of Union Carbide and Carbon Corporation 4 POST OFFICE BOX X + OAK RIDGE, TENNESSEE This document contains. trigi_g%,,n_q " VEY=REY 67 1754, Its transmittal or the disc i the Atomic osure of its contents .—M prghibited. DISCLAIMER This report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency Thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. 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AEC Technical Information Service Extension Oak Ridge, Tennessee ~ THIS PAGE WAS INTENTIONALLY LEFT BLANK THIS PAGE WAS INTENTIONALLY LEFT BLANK o THISPAGE | WAS INTFNTIONALLY | | LEFTBLANK PREFACE In September, 1955, a group of men experienced in various scientific and engineering fields embarked on the twelve months of study which culminated in this report. For nine of those months, formal classroom and student laboratory work occupied their time. At the end of that period, these seven students were presented with a problem in reactor design. They studied it for ten weeks, the final perlod of the school term. - This is a summary report of their effort. It must be reelized that, in so short a time, a study of this scope can not be guaranteed complete or free of error. This '"thesis" is not offered as a polished engineering report, but rather as a record of the work done by the group under the leadership of the group leader. It is issued for use by those persons competent to assess the uncertainties inherent in the results obtained in terms of the preciseness of the technical data and analytical methods employed in the study. In the opinion of the students and faculty of ORSORT, the problem has served the pedagogical purpose for which it was intended. The faculty joins the authors in an expression of appreciation for the generous assistance which various members of the Oak Ridge National Laboratory gave. In particular, the guidance of the group consultants, E. S. Bettis and D. A. Carrison, 1is gratefully acknowledged. Lewis Nelson for The Faculty of ORSORT ACKNOWLEDGMENT The group wishes to express its appreciation to the many peéple of the Oak Ridge National Labora£ory wvho gave us the benefit of their experience and knowledge by supplying advice and information toward the completion of this desigp study. Our special thanks go to E. S. Bettis; our group advisor, for his guid- ance and continuing interest in tfie project. Finally, we want to thank the ORSORT faculty and staff for providing us with the education and knowledge which made it possible for us to undertake this design study. ABSTRACT 2 The reactor is a fused salt, homogeneous, intermediate energy reactog - that operates at a power density of 55 watts/cc with an enriched U-235 in;V vehtory of 240 kilograms. The fuel bearing fused fluoride salt funpfiions as neutron moderator and heat transfer medium. It 1s cfreulated through the multi-pass reactor veésel and transfers the fission heat to a sodium loop within the reactof vessel by an annular, U-tube heat exchanger which surrounds the central core. This heat is then transferred from this radiocactive sodium loop to an intermediate, non-radioactive sodium loop, which is used to generate 1000°F superheated steam at 1800 psig to drive a steam turbine-generatpr unit. The net thermal plant efficiency is 38.2 percent at a maximum net electrical output of 229 MW. TABLE OF CONTENTS Page Number CHAPTER 1. SUMMARY DESCRIPTION AND CONCLUSIONS ek 15 1.0.0 Introduction ' 15 1.1.0 Fuel 15 1.2.0 Materials 16 1.3.0 Reactor 17 1.4.0 ‘Sodium Coolant Systems 18 h.5 1. 50 Steam Pofier System l 20 h.é 1.6.0 Conclusions n .1 st il s 20 \ 4.7 CHAPTER 2. FUEL : 23 4.8 2.0;0 Ifitroduction ¥ 2.1.0 Composition 2.2.0 Physical and Thermal Properties 2.3.0 Nuclear Properties 2.4.0 Availebility and Cost 2.5.0 Additiz;Hgf Uranium Fuel 2.6.0. ‘Fuel Reprocessing CHAPTER 3. STRUCTURAL MATERIALS 3.0.0 Introduction ’ L0 Reaétor Structursal Materialé 3:1.I Inconel . . 3.1.2 Nickel-Molybdenum Alloys 3.1.3 Nickel Clad Stainless Steel 3.2.0 Sodium System Materials 3.3.0 Steam System Materials '-hp REACTOR ANALYSIS Introduction Critical Size and Fuel Concentration sutron Flux BAiy of Reactivity [ L ID"' AT EXCHANGER DESIGN ary Heat Exchanger "y .1.1 Internal vs. External Arrangement .1l.2 Internal Arrangement ~.l 3 Basic Design Criteria .1.4 Parametric Study 5.1;5 Desigh Considerations Reactor Vessel > 1 Shell Design 2 Internal Arrangement 3 Structural Arrangement .4t Fuel Circulating Pumps : 5 Pressurizer, Gas Removal and Expansion System 6 Effect of Volume Heat Sources SODIUM AND STEAM POWER SYSTEMS - Introduction Primary Sodium Coolant Loop 6.1.1 Schemes for Removing Heat from Reactor 6.1.2 Choice of Coolant 6.1 3 Sodium Coolant Activation * 6.4.0 6.5.0 6.6.0 6.7.0 6.8.0 | .' . (cgn;tnued) "Intermediate Sodium Coolant Loop l'lfJuatification of Intermediate Loop 2 Choice of Coolant Fluid 3 Sodium-Water Isolation Problem - 4 Proposed Design for Sodium Water Isolation Intermediate Heat Exchanger 6.3.1 Isolation of Radioactive from Non- Radioactive Sodium 6.3.2 Calculations for Intermediate Heat Exchanger Steam Gengratqrs 6.4.1 Once-Through Steam Generator 6.4.2 Natural Circulation Boiler-Separate Superheater Steam Generator Steam Generator Calculation Procedures 6.5.1 Heat Transfer 6.5.2 Pressure Drop Summary of Heat Exchanger Data Sodium Piping Systems W2 g, 6.T.1 Pipe Length and Size 6.7.2 Pipe Slope 6.7.3 Expansion Tanks 6.7.% Cold Traps 6.7.5 Valves 6.7.6 Pressure and Instrument Taps 6.7.7 Drain and Charge Tanks 6.7.8 Cover Gas 6.7.9 Sodium Pumps 6.7.10 Heating the Salt and Sodium Loops for Startup Steam Power System 6.8.1 Steam Turbine and Steam Cycle Heat Balance 2 Calculation Procedure .8.3 Auxiliary Power Requirements, Net Plant Qutput and Efficiency 125 130 T-5.0 9. Page Number PTER 7. POWER PLANT LAYOUT, OPERATION AND MAINTENANCE 133 7.0.0 Introduction , 133 .0 General Piping Arrangement 133 T-1.1 Primary Loop Piping 133 7.1.2 Intermediate Loop Piping 134 7.1.3 Steam Generator Connections 134 T-1.4 Steam and Water Piping 134 2,0 General Arrangement of Plant Components 13k 3.0 Shielding % i g ik 135 4.0 Maintenance b 139 T.4.1 Fuel Circulating Pumps 140 T.4.2 Primary Heat Exehanger Tube Bundles 140 T-4.3 Intermediate Heat Exchangers and Primary 141 Sodium Pumps .0 Partial Load Operation 141 7.5.1 Mathematical Approach 142 7-5.2 Reactor Simulator Analysis 143 F0 figgnqmi; g 145 T7.6.1 Factors Reguiring Additional Investigation 151 T.6.2 Approximate Economic Analysis 152 hysical and Thermal Properties 15k esign Curves 171 ieactor Analysis Calculations 176 rimary Heat Exchanger Calculations 189 [ntermediate Heat Exchanger Calculations 201 Steam Generator Calculations 207 falculations of Steam Plant Heat Balance 227 odium Piping Pressure Drop Calculations 234 'flfiyésflip (continued) S I. Partial Load Operation (Mathematical Approach) ) Reactor Simulator Analysis (Supplied by Dr. E. R. Mann) .l e, *Test of Double Tube Sheet Design for Water-Sodium Isolation .< bols Used in Engineering Calculations 'Tejéia Used in Nuclear Physies Calculations o'y I Bi b 11 Qgraphy ¥ e 10. Page Number 237 o bm'f 3 2 " 2lgl 251 255 257 258 LIST OF ILLUSTRATIONS Heat Tranéfer Diagramé - Fuséd Salt_Powér fieactor S&stems Phase Diagram of the'Three;CofiponentfiNaF-ZrF4¥UFh'System Phase Diagram of the Two-Component:NgFfszh-System. ’ Par£1a1 Pressureé of ZrF), . | | Changes in Attack with Increasing Operating Time in Forced-Circulation Inconel Loops Corrosion of 85-15 Nickel-Molybdenum Alloy by Fused Fluoride Salt No. 30 in a Thermal Convection Loop Corrosion of Nickel.in Fused Fluoride Salt No. 30 in a Thermal Convection Loop ‘ Variation of Mui%ifilication Constant K with Concentration of U-235 and Central Core Radius Total U-235 Invéntory Variation with Central Core Radius Neutron Flux Distribution Reactor Arrangement with Straight Heat Exchanger Tubes Isometric View of Beactor Prihgry Heat Exchanger Paramétric Study, 1/2-inch Tubes . Primary Heat Exchanger Parametric Study, 5/8-inch Tubes Primery Heat ExchangerlParametric Study, 3/4-inch Tubes . Primary Heat Exchanger Parametric Study Summary, Variation of Number of Tubes with Tube Ligament Ratio and Size of Tubes Primary Heat Exchanger Parsmetric Study Summary, Variation of Pump Power, Fuel Holdup Volume, and Heat Exchanger Surface with Tube Ligament Ratio and Size of Tubes Primary Heat Exchanger Tube Layout Section Through Reactor Vessel Plan View of Reactor Vessel 11. Page Number 19 2> 26 29 39 Lo b1 50 52 62 63 69 70 T1 T2 T3 15 78 9 ¢ * Figure Number 5.11 6.1 6.2 7.1 2 T-3 T.l 7.5 7.6 Pressurizer, Fuel Expansion and Gas Removal System Suggested Test of Double Tube Sheet Design for Isolation of Water and Sodium Heat Balance Diagram for 600 MW Reactor-Steam Power Plant Sodium and Steam Power System Elevation of Power Plant Plan View of Power Plant Time Constants and Heat Capacity of Components in Once- Through Steam Generator System Partial Load System Temperatures without Auxiliary Heat Exchanger Temperature Control 7 + Partial Load System Temperatures with Auxiliary Heat Exchanger Temperature Control T 0 R Qemperature and Reactor Power Transients Due to Rapid 7.8 A.l A.2 A.3 Ak A.5 A.6 A.T A.8 A.9 Increase in Load Demand without Auxiliary Heat Exchanger " Temperature Control Temperature and Reactor Power Transients Due to Rapid Increase in Load Demand with Auxiliary Heat Exchanger Temperature Control Al 3 Thermal Conductivity of Selected Steels Specific Heat of Water Viscosity and Thermal Conductivity of Saturated Water Density of Sodium, Potassium and Sodium Potassium Alloys Viscosity of Sodium, Potassium and Sodium Potassium Alloys Thermal Conductivity of Sodium, Potassium and Sodium Potassium Alloys opecific Heat of Steam at Constant Pressure Thermal Conductivity of Steam Viscoselty of Steam 150 155 156 157 Figure Number A.10 The B.l Hea Tur B.2 Heal B.3 Heat B.h Tube C.1l Neut G.1l Pres J.l Simu J.2 Simu] J.3 Simul Heat K-1 Under K-2 Defec | Thermal Conductivity of Nickel - Heat Transfer of Steam or Subsaturated Water in Turbulent Flow Heat Transfer from Flat Plates Heat Transfer from Bare or Insulated Steel Pipe Tube Count for Tubes Spaced on Equilateral Centers Neutron Flux Distribution Pressure-Enthalpy Relation of Expansion Line Simulator Circuit fgf Fuel and Primary Sodium Loops Simulator Circuit for Intermediate Sodium Loop - Simulator Circuit for Driving Function in Intermediate Heat Exchanger and Regulator for Steam Temperature Control 1 - Underside of Brazed Tube Sheet i Defective Brazing of Tube 1k, LIST OF TABLES Table Page Number Number h.1 Summary of Reactor Analysis Data 57 5.1 Summary of Primary Heat Exchanger Parametric Study s 5.2 Specifications for Fuel Circulating Pumps 88 6.1 Pripisry #eat Exchanger Specifications 11k combines th 6.2 Intermediate Heat Exchanger Specifications 115 | reactor des 6.3 Once-Through Steam Generator Specifications 116 reactor and 6.4 Convection Boiler Specifications 117 ‘nology wher 6.5 Single Wall Superheater Specifications 118 | : ' S could opera 6.6 Double Wall Superheater Specifications 119 minimum of 6.7 Summary of Full Load Operating Data 132 coolants, h T-1 Estimated Power Generation Costs 153 been made wi A.l Thermodynamic Properties of Liquid Sodium _ 165 | to fulfill 1 A.2 Selected Properties of Stainless Steels 169 ment work wi A.3 Selected Physical Properties of "L" Nickel 170 into operati C.1l Basic Microscqpic Cross Section Data Used to Obtain _ 178 3-Group, 34Region Code Constants intended to c.2 Case Data and Computed Multiplication Constants cC.3 Macroscopic Group Constants for 3-Group, 3-Region Code H.l Sodium Piping Data TABLE OF CONTENTS SUMMARY DESCRIPTION AND CONCLUSIONS Introduction Fuel Materials Reactor Sodium Coolant Systems Steam Power System Conclusions FUEL Introduction Composition Physical and Thermal Properties Nuélear Properties Aiailability and Cost Addition of Uranium Fuel Fuel Reprocessing STRUCTURAL MATERTALS "Introduction Reactor Structural Materials Sodium System Materials Steam System Materials I} 5.2.0 REACTOR ANALYSIS Introduction Critical Size and Fuel Concentration Neutron Flux Fuel Inventory Fuel Burn-up Uranium-233 Fuel Fission Product Poisoning Temperature Coefficient of Reactivifif Decay Heating | REACTOR AND PRIMARY HFAT EXCHANGER DESIGN Introduction | Primary Heat Exchanger 1l Internal vs. External Arrangement 2 Internal Arrangement .3 Basic Design Criteria 4 Parametric Study 5 Design Considerations Reactor Vessglv Shell Design Internal Arrangement Structural Arrangement Fuel Circulating Pumps - N AV AN O PPPPOND v FWw o Effect of Volume Heat Sources SODIUM AND STEAM POWER SYSTEMS Introduction Priméfy Sodium Coolant Loop 6.1.1 6.1.2 Choice of Coolant 6.1.3 Sodium Coolant Activatior Pressurizer, Gas Removal and Expansion System Schemes for Removing Heat from Reactor Page Number CHAPTER 6. 6.2.0 6.3.0 6.4.0 6.5.0 6.8.0 O\ O\ O\ O\ O\ O\ O\ O\ O\ O© (continued) Intermediate Sodium Coolant Loop 1 Justification of Intermediate Loop 2 Choice of Coolant Fluid 3 Sodium-Water Isolation Problem - 4 Proposed Design for Sodium Water Isolation Intermediate Heat Exchanger 6.3.1 Isolation of Radioactive from Non- Radioactive Sodium 6.3.2 Calculations for Intermediate Heat Exchanger Steam Generators 6.4.1 Once-Through Steam Generator 6.4.2 Natural Circulation Boiler-Separate Superheater Steam Generator Steam Generator Calculation Procedures 6.5.1 Heat Transfer 6.5.2 Pressure Drop Summary of Heat Exchanger Data Sodium Piping Systems | Pipe Length and Size Pipe Slope Expansion Tanks Cold Traps " Valves Pressure and Instrument Taps Drain and Charge Tanks Cover Gas ' Sodium Pumps Heating the Salt and Sodium Loops for Startup L] o I, O, RN, IR QN T, T, UK AR, (-] =0 - OV W N O (-3 o -] © [ L] -] -] Steam Power System 6.8.1 Steam Turbine and Steam Cycle Heat Balance 6.8.2 Calculation Procedure - 6.8.3 Auxiliary Pover Requirements, Net Plant Output and’ Efficiency 99 100 100 102 102 102 10k 107 107 110 113 113 113 120 120 121 121 121 121 122 123 123 125 125 127 130 CHAPTER 7. POWER PLANT LAYOUT, OPERATION AND MAINTENANCE 7.0.0 Introduction T.1.0 General Piping Arrangement T-1.1 Primary Loop Piping T.1.2 Intermediate Loop Piping T.-1.3 Steam Generator Connections T.1.4 Steam and Water Piping T.2.0 General Arrangement of Plant Components T.3.0 Shieldirg T-4.0 Maintenance - 7.4.1 Fuel Circulating Pumps T-4.2 Primary Heat Exchanger Tube Bundles T.4.3 Intermediate Heat Exchangers and Primary Sodium Pumps 7.5.0 Partial Load Operation T.5.1 Mathematical Approach T.5.2 Reactor Simulator Analysis 7.6.0 Economics 7.6.1 Factors Requiring Additional Investigation 7.6.2 Approximate Economic Analysis APPENDIX A Physical and Thermal Properties B. Design Curves C. Reactor Analysis Calculations — D. Primary Heat Exchanger Calculations E. Intermediate Heat Exchanger Calculations F. Steam Generator Calculations G. Calculations of Steam Plant Heat Balance H. Sodium Piping Pressure Drop Calculations 13. Page Number 133 1133 133 13 134 134 134 134 135 139 140 140 141 1 142 143 145 151 152 15k 171 176 189 201 207 227 234 APPENDIX (continued) I. Partial Load Operation (Mathematical Approach) J. Reactor Simulator Analysis (Supplied by Dr. E. R. Mann) K. Test of Double Tube Sheet Design for Water-Sodium Isolation Symbols Used in Engineering Calculations Symbols Used in Nuclear Physics Calculations Bibliography 1k, Page Number 237 2l 251 255 257 258 k4 15, CHAPTER 1. SUMMARY DESCRIPTION AND CONCLUSIONS 1.0.0 INTRODUCTION This report is a study of the feasibility of using a fused salt fuel reactor in a central station electric generating plant. The propoBed reactor .combines the desirable features of various homogeneous and high temperature reactor designs. The basic philosophy of the study has been to design a reactor and pbwer plant system which could be built using'present day tech- nology wherever possible; whicfi would be reliablé, safe, and efficient; which could operate for long periods of timé; and which could be maintained with a minimum of difficulty. The selection of the fuel, reactor materials, reactor coolants, heat transfer system, steam turbine and associated equipment has been made with the above objectives in mind, and the resulting design appears to fulfili these conditions. This gtudy indicates, however, that some develop- ment work will be necessary beforefifi plant of this type can be built and put into operation. The basic plant arrangement presented in this report is not | intended to represent a finished or optimum design, but it does appear to be: reasonable and possible and warrants further study and consideration. 1.1.0 FUEL The plant has been designed to utilize current technology in the deyelop- ment of fused salt fuels. A reactor experiment has been conducted successfully using a solution of sOdifim, zirconium and uranium fluorides for the fuel system. This experiment indicated that a fused salt fuel of this type has the char- acteristics required for the homogeneous reactor under consideration. The moderating material in the reactor is the fluorine in the fused salt. With the 16. composition contemplated in this design, the number of fluorine atoms is in- sufficient to thermalize the majority of neutrons before fission takes place. As a result, approximately 34 percent of the fissions occur in the intermediate and fast energy ranges. | A fused fluoride composition of 43 mol percent Zth and 57 mol percent NaF was selected as the fuel base to obtain a compromise between a reasonable fielting point (9329F), corrosion, and the ZrF), "snow problem". The concen- tration of the uranium is low and has little effect on the over-all characteristics of the salt. Uranium is added to the salt in the form of (NaF),UF). Pellets, powder or & dissolved solution of the compound can be added during operation to replace uranium burn-up and to overcome the build-up of fission product and corrosion poisoné° Economically, it probably would not be desirable to repro- cess fuel to remove the solid fission products in order to minimize the amount of uraniwm which must be added to override the effect of these poisons. 1.2.0 MATERIALS Fused salts are extremely corrosive to conventional engineering struc- tural materials. It has been shown that alloys containing large percentages of nickel are the most corrosion-resistant metals to fused fluoride salts. Inconel, nickel-molybdenum alloys and pure nickel were considered as possible structural materials for the reactof. It has been shown that the chromium in Inconel diffuses from the metal! at a fairly substantial rate and, therefore, Inconel does not appear $0 be a satisfactory structural material for a long life reactor. Nickel-molybdenum alloys have been shown to have excellent corrosion ‘resistance to fused fluorides. A series of these alloys is being developed e 17. in an attempt to obtaifi a strong, fabricable material for use in high tem- perature fused salt reactors. It appears that this development program will be successful, but at the time this report was prepared, the exact properties and fabricability of these new alloys were not sufficiently established to presuppose a satisfactory reactor design. The amount of available data on the corrosion of pure nickecl in fused fluorides is limited; however, these data indicate little or no corrosion or. mass transfer. On this basis, it appears that nickel has the corrosion resis- tant properties for long reactor life. Unfortunately, at the operating temperatures contemplated, the strength characteristics of nickel are inadequate, and the nickel must be strengthened by suitable.structural materiglsa There- fore, to present a design using present day technology, nickel clad stainless steel was selected as the most suitable structural material for the reactor in this study. Because of the limited experimental data available, there is still some question as to the suitability of nickel to resist corrosion to the degree desired. The use of nickel cald stainless steel also poses fabricating and inspection problems. 1.3.0 REACTOR The reactor is a homogeneous circulating fuel reactor with the primary heat exchangers an integral part of the reactor vessel. The heat exchangers: are located in a baffled annulus surrounding the seven-foot diameter cylin- drical core. Fuel flows upward through the core, through the U-tube hgat exchafigers, énd through an annular downcomer at the periphery of the vessel. Eight vertical axial flow pumps located symmetrically around the top of the vessel maintain the fuel circulation in the reactor. 18. . The reactor is designed to produce 600 MW of heat with fuel inlet and outlet temperatures of 1050°F and 12009F, respectively. The pumps and the entire reactor vessel, with the exception of the top, are subjected to the ' minimum fuel temperature. The reactor vessel is a sixteen-foot diameter vertical cylinder with a ¢ dished bottom and protrusions near the top where the fuel pumps are located. The only penetrations in the vessel are a series of drain connections in the bottom; all other fuel and coolant 1inés enter the top of the vessel. The heat exchanger baffles are supported internally, and the heat exchanger bundles are supported from a circular cantilever arrangement extending radially inward at the top. The heat exchangers and pumps are arranged to permit removal.vertically from the reactor. The layout of the equipment was arranged to provide for replacement of components with a minimum of difficulty, since outages for maintenunce should be nf short duration. Provisions are made for the continuous removal of xenon, and space is provided, external to the reactor, for the expansion of fuel. The reactor will have a negative temperature coefficient of reactivity which will furnish the only means of reactor load control. The design average temperature will be controlled by the proper addition of uranium to the fuel solution. "1.4.0 SODIUM COOLANT SYSTEMS The sodium system is made up of two independent coolant loops. A heat transfer diagram which illustrates the temperature drops through these loops is shown in Figure 1.1. The sodium in the primary loop flows from twenty; four heat exchanger bundles located in the reactor vessel to a common ring . header above the reactor. Six independent circuits of the primary loop are ORNL=LK=Dwg,=18169 19. Figure % | ; HEAT TRANSFER DIAGRAMS : FUSED SALT POWER REACTOR SYSTEM 1200 Fuel Sodium o = S 3 Primary , SRS 7 {Heat Exchanger 1100 Sodlum ~ N~ . o e N SRR £ 48.0 x 106 1b/hr. - i h-NEENEE NN Water- . 45.7 x 100 1b/hr 1000 Steam A\ N 14 ‘ \. - o e = e ¢ \ ‘\\( Intermedliate g "Heat Exchanger N \\ 900 N\ / N HF s / - Fx :: i ; / N ol ] A 1.2 T \ + 5 T 800 T \ ! _ . W [T Steam A — \ Generator ] SR L\ 15.9 x 10 1b/ar ] - \\ o AT = 700 \‘ \ 6003 \ \ \| 1.96 x 10 1b/nr 500 400 * Percent of Heat Transfer 20, connected to the ring header. This division is used so that piping and pump sizes in the primary loop are reasonable and practical. The six intermediate heat exchangers are located in shielded compartments surrounding the reactor shield. There are two intermediate loop circuits, each connected to three inter- mediate heat exchangers and one steam generator. Superheated steam from the two steam generators is used to drive a steam turbine. 1.5.0 STEAM POWER SYSTEM Fused salt reactors are especially attractive for central station appli- cation because of their high temperature characteristics. In this study, the steam turbine and associated equipment have been selected for operation with 1800 psig and 1000°F steam. A 250 MW tandem compound triple flow turbine will be connected to a 3600 rpm conductor cooled generator. Five closed feedwater heaters and a deaerating heater are included in the feed water heating cycle. At raled load, the gross generator output is 245,400 kw, the plant auxiliary power requirement is 16,400 kw, and the net thermal efficiency is 38.2 percent. 1.6.0 CONCLUSIONS This design study of a fused fluoride salt reactor central station power plant has shown that the proposed system is technically feasible. It is con- sidered that such a plant can be built with a minimum of development work, and that the developments would primarily be associated with the reactor. The concept of using a fused fluoride salt as a homogeneous self-moderating fuel solution has been established as practical. It has been shown that a large fraction of fissions are caused by epithermal neutrons, which classifieg this design as an intermediate reactor. 21. No completely satisfactory materials are currently‘available for the con- struction of the reactor vessel. Nickel clad stainless steel has many of the necessary requirements; but fabrication and inspection techniques must be improved. A series of nickelamolybdenufi alloys are in the process of develop- ment which appear to have excellent corrosion and high temperature strength properties. The success of the fused fluoride salt reactor may depend upon satisfactory completion of this metallurgical development program. The high temperature characteristics of the fused fluoride salt reactor- make it possible to use a high efficiency steam turbine cycle for the con- version of fission heat intoc electrical energy. The net thermal efficiency of the proposed design is 38.2 percent, and it is possible to design for even a higher efficiency if the over-all plant economics show that this is desirable. The operation of the reactor is relatively simple. The reactor power iS'controlled by its negative temperature coefficient of reactivity. Uranium will be added to the fused fluoride salt to replace fuel burn-up and tc over- ride the build-up of fission product poisons. Continuous fuel reprocessing should not be necessary for extended periods of operation. | The reactor has been deésigned to burn uranium-235. However, sincg thig is an intermediate reactor, a number of advantages would result if uranium-233 would be bred during operation. The production of new fissionable material vould reduce the fuel addition rate. The fuel bred ip the reactor could be utilized without the neceséity of employing chemical separation techniques. .. The mechanical design features of the reactor are novel. The reactor vessel is & multi-pass cylindrical container. The primary heat exchanger tube bundles, which are located inside the reactor vessel, and the fuel cir- culating pumps can be removed for replacement by lifting them vertically 22, from their sufiportan The critical components of the reactor and primary sodium coolant circuits can be maintained with a minimum of difficulty. All components in the system, with the exception of those inside the reactor - primary shield, can be maintained directly. It is the considered opinion of the authors that the fused salt reactor - not only shows great promise as the heat source for a large central station ‘power plant, but that it is one of the most desirable reactor types for this application. 23. CHAPTER 2. FUEL 2,0.0 INTRODUCTION The wmost.outstanding advantage of fising a fused salt as the fuel solvent and heatv exchanging medium is that it enables the attainment of high tempera- tures without high pressures. 1In addition; the advantaggs inherent in homo- geneous systems are possible; i.e., no neutron absorbing structural material in the core proper; possibility of fission product removal, with the important gaseous fission products {xenon) being especially easy to remove; nc fuel element fabrication; extended periods of operation possible by the periodic addition of a fuel conéentrate; and high permissible fuel burn-up. The particular fused salt used should have a reasonable melting point and must be stable at high temperatures and under high radiation. The elemental constituents of the salt should have small neutron capture cross sections. As in the cagse of the reactor in this study, it may be required that the fused salt solvent also serve as the neutron moderator. In addition to these specialized requirements, the salt must meet the more general con- siderations of availability, reasonable cost, permissible toxicity, and tolerable corrosiveness. A good deal of thepretical and experimental work has been, and is.still being done, by other groups in this field to select the constituents and the composition of a fused salt that best meets the above or very qimilar require~ ments. The result of this work indicates that a fused salt composed of NaF, ZrF), and UF) is satisfactory for use as a liquid fuel for reactors. A reactor - has been successfully operated using this fuel system. After considerable 2L, literature‘research and discussion with Oak Ridge National fiafioratory per- sonnel; it was considered that this particular salt system did fulfill the design requirements for the proposed reactor and also that this was currently 8till the most desirable system. By utilizing this salt system, there was immediately available a considerable amount of experimental information, physical properties data, and practical experience necessary for the intelli- gent design of a plant using such a salt as the reactor fuel solvent. 2.1.0 COMPOSITION Within the NaF -ZrF) -UF) system; there remained to select the optimum composition of these constituents for this design study. Figure 2.1 shows an isometric view of the temperature-composition-phase plot for the three component NaF -ZrF) -UF), system. Nuclear calculaticns indicated that the frac- tion of UF) required is small (<0.2 mol %), so that for many purpéses the small salt system may be considered as a two-component system of NaF -ZrF) . Figure 2.2 shows a plot of the temperature—composifiion-phase diagram for the two-component NaF-ZrF) system. From both the two-component and three=component plots; it can be observed tha£ there exist two relatively low melting eutectics in the low UF), content region of interest. These occur at about 42 and 59 mol % Zr'Fh° In order to obtain the minimum melting point, it would be desirable to specify the salt to be one of these eutectic compositions; however, there are other factors that need to be considered in the composition selection. Operational experience with NaF-ZrF) fused salt syétema revealed the existence of a "snow problem" resulting from the low, but finite, vapor pres- sure of ZrF) at operating temperatures. A simplified explanation of this pPhenomenon is that ZrFu-vapor will escape from the fused salt to any existing gas volume within the fused salt system. The subsequent sublimation of ZrF) -25- ORNL-LR-DWG 3398A NaF COMP. NO, 34 43 mol Y% ZrF4 Phase Diagram of the Three-Component NaF -ZrF,-UF, System. Figure 2.1 TEMPERATURE (°C) _gz_ 1000 ORNL-LR -Dwg. -12}J1A N\ .‘ 7 o \V//’ ////// s / | &% / = - O \ %E | / 600 \\IJIQ // Y N S N No3ZrF7—// i NaZrFg (METASTABLE) 400 - | il Na,ZrFs. +— | I O {O 20 30 40 50 60 70 80 90 {00 ZrF,4 (mole %) rigure 2.2 Phase Diagram of the Two-Component NaF —ZrF, System. I I 27. crystals on lower temperature surfaces results in the formation of snow-like Zth crystals which can build up and accumulate to such an extent that plugging of tubing or even filling the entire gas volume may result. This most undesir- able phenomenon can be significantly imfroved by reducing the fraction of ZrFu in the fused salt composition. Other factors are in opposition to shifting the fused salt composition in this direnfion" For nucleur considerations, it is desirable to have the highest fraction of fluorine atoms possible in the fused salt solvent. .The Zth_salt constituent has four times as many fluorine atoms for each metal atom as the NaF constituent. Further, the corrosiveness of the fused salt is decreased by increasing the fraction of ZrF) and decreasing the NaF. These - factors both suggest having as high a mole percentage of Zch a8 possible. For this design study, the relatively slight effect on corrosion and nuclear properties by the small changes in composition that are possible were ‘considered of lesser consequence than the more significunt effect these small changes in composition have on the "snow problem". For these reasons, then, it is desirable to utilize a fused salt of composition close to the lower ZrF) content eutectic (41 mole percent). An examination of the temperature-composition-phase diagrams shows that Just below the 41 mole % Zth eutectic, a large increase in melting point - temperature results with only a slight_decrease in ZrF), composition. Such a decrease would quite probably occur during operation, since the uranium fuel concentrate is added as the complex salt (NaF),UF,. To safeguard against the poesibility of the fused salt composition falling into this region after a period of reactor operation, it was considered advisable to accept a very slightly higher fraction of ZrFu than the exact eutectic. For these considerations, 28. an already standardized composition, Composition No. 34 (Ref. 31) was there- fore selected. Comp. number Component Mole % Weight Liquidius Temp. 34 NaF 57.0 2l ,98 500°C, 932°F ZrF), 43.0 75.02 - 2.2.0 PHYSICAL AND THERMAL PROPERTIES The physical and thermal properties of this salt composition are as follows (Ref. 17): | fiensity: 3.86 gm/cc at room-temperature € (gm/cc) = 3.65 - 0.00088 x (°c) * 5% € (1v/£t3) = 228.8 - 0.0305 x (°F) * 5% Heat Capacity: 0.27 cal/gm - °C at 700°C Thermal Conductivity: 1.3 BTU/hr-£t-°F Viscosity: 600°C =-mcnemcan 7.5 cp (Ol — 4.6 cp 8009C --c-cemmeaa 3.2 cp A good indication of the pfobable degree of the "snow problem" discussed briefly above 1s given:by the vapor pressure of ZrF) that exists at the tem- perature of operation. Figure 2.3 shows the variation of the partial pressure - - of ZrF) with temperature for various compositions of the NaP -2xF), éystem (Ref. 26). An interpolation of the plotted data to the 43% ZrF), compésition and extrapola- tion down to the design operating temperature (1050°F, 565°C) of the fused salt - 1000 T - S l h 10C N 748 N CNIN o O ‘* 100 891 T 13 — £Y7 W . 419 oI\ T SOO\\\ ‘:\‘ '&,\ < \\ N *\1‘ N \ N\ T \:\ "~ . 10— ~ ~C ~ T~ ~ s N I BN q N R CNERN \\i\\.\ N - N NN E i \\ M ’\\ ~ DN £ S ~ - - X ‘w - N S AN N a A i e N j‘\\ o 283 \ ~ C% Pa ' <3 N O'S'-Z‘é . N [ 207 Joy X EONSENS ¢ ‘\|95 4 2 ™~ 0.01}—x : : % - - \:\ ~ °-tfius 00— =8 80 82 ¢ B¢ 88 9C 97 97 9% 38 00 102 104 106 108 110 112 1€ 16 N8 ’ : Reciprocal Tempaerature, -T'-?K—] | : 1 e | b 1050 1000 900 800 ) TO00 600 " Temperature, C 565 C 1050 F - FIGURE 2.3 PARTIAL PRESSURES OF Zrf, BASED ON ASSUMPTION THAT ONLY Nof AND Zrf, EXIST IN VAPOR PHASE Figures on plot denote mole per cent ZrF,. Source: BMI -~ 1064; "Vapor Pressures of the Sodium Fluoride - Zirconium Fluoride System ard Derived Information", by Karl A, Sense, et, al. ' 0LTQT= *3MI=HT~-TNIO ‘62 30. — at the location of the gas surface shows the partial pressure of Zth to be less than 0.01 mm mercury. It was the considered opinion of personnel experienced in this field that a "snow problem” would be virtually non-existent for Z1F), vapor pressures less than 0.0l mm mercury, even for extended periods of operation of the order of years (Ref. 18). In addition, mechanical appliasnces called "snow traps" have been developed which prevent plugging of the critical areas, and these, or similar innovations; could be adapted to this design if required. 2.3.0 NUCLEAR PROPERTIES The nuclear characteristics of the NaF-ZrF) salt are not outstanding, how- ever, they are tolerable. The absorption croes sections, particularly of the sodium atom for thermal neutrons; are quite large. Neutron moderation is primarily performed by the fluorine atom which, with an atomic weight of 19, is rather poor in this respect. Alternate salts could be suggested (i.e., BeFp) with superior nuclear properties to those of the NaF -ZrF) salt, but other con- siderations such as corrosion, viscosity, stability; melting point, etc.; make the selected Na¥F-ZiF) sult the most demirable at this time from an over-all consideration. 2.4.0 . AVAILABILITY AND COST The availability and cost of this salt were investigated (Ref. 19). The NaF constituent is commercially available at $0.12 per pound. Hafnium separated - ZrF), has been produced by Oak Ridge National Laboratory from Hf separated ZrCl) at a total cost of $3.15 per pound. The NaF and ZrF) salt constituents are blended as powders and then fused in a batch process at 1500°F to form the com- position desired. This fused salt must then be processed to remove trace quantitiés of sulfur, iron, nickel,‘EQO, oxides, chlor;des, and other impurities for corrosion pufposes° This treatment is performed at 1500°F with hydrogen and hydrogen fluo- ride'gaa° At the present time, ORNL can produce a hafnium-free, processed NaF-ZfiFu fused salt ready for use as a reactor fuel solvent for approximately 31 $7-.50 per pound. It is the considered opinion of the responsible Oak Ridge National Laboratory personnel that increased capacity and technical develop- ments would very probably reduce this cost to $5.00 per pound, or less. - The nuclear calculations for this project were based on the assumption that the elemental zirconium constituent of the salt is pure, the hafnium impurity having been removed. This has the effect of minimizing the uranium inventory by decreasing parasitic absorption at the expense of an increase in the cost of the salt. 2.5.0 ADDITION OF URANIUM FUEL - -The uranium fuel burn-up is compensated by the periodic addition of a uranium concentrate—(NaF)EUth From the three constituent phase diagram (Figure 2.1), it is observed that as the uranium concentrate compound is dis- gsolved in the composition number 3l salt solvent, only compositions of con- sistently lower melting points are formed in the dissolution process. The dissolution should, therefure, proceed without difficulty. The physical addition of the uranium concentrate can be accomplishéd in several ways. The concentrated uranium salt could be added directly to the fused salt in the form of pellets. It could be added more or less continuously as a powder, or it could be dissolved in a small, isolated quantif& of the fused salt solvent in-a continuous or batch process and then injected into‘the g reactor. The final selection oflthe preferred methodlwould need to be deter- mined by a consideration of the hazardé involved and the ease of operation. The quantity or rate of fuel concentrate addition is controlled by the average temperature of the fused salt in the reactor core (1125°F design). As fuel is burned up, this average temperature would be gradually reduced by the reactor's negative temperature coefficient, which keeps the reactor "just critical”. When this temperature falls below 1125°F by some predetermined ) 32. temperature increment, a carefully controlled quantity of fuel concentrate would be added or, if this is a continuous process, the rate of addition would be increased. The average temperature would thereby be raised a regu- lated increment above the average design temperature and the cycle would begin again. 2.6.0 FUEL REPROCESSING Fission products and corrosion products will accumulate in the fuel solu- tion during operation. Additional uranium will have to be added to override the poisoning effect of these nuclear poisons if they are allowed to remain in the reactor. For this reason, it may be necessary to provide a means of fuel reprocessing to remove some of these poisons. 8 Xenon has a very low solubility (n{lO moles/cc @ 1 atmos. of xenon) in thg fused salt fuel solution and will therefore be removed by the off-gas system. The gaseous iodine fission product which is the precursor of xenon will also be removed by the off-gas system. If this is accomplished before the iodine;, which has a 6.7-hour half-life, decays to xenon, the reuctor core will operate with little or no xenon present. Some of the metallic fission products (i.e., Ru, Mo) do not form fluoride compounds and will not dissolve in the fused salt solvent. They will probably plate out on the surfaces within the reactor. This has the advantage of placing these nuclear poisons in regions of low flux. The effect of these plated metals on corrosion and heat transfer will require further investigation. The leasgt. soluble of the fluoride forming fission products is believed to be that of the rare earth, CeF3, vhich is soluble up to 3 wt. %.At the tempera- ture of this design study.(Ref. 32). Approximately one-third of the fission products resulting frqm,fission are rare earths.. Even by making the pessimistic 33. assumption that all the rare earths must be less than 3 wt. % in order tpgt they will not precipitate, the solubiiity limit of the rare earth fluori@es would not be reached for many years of operation (>20 years). - The conclusion that may be derived from this is that the fused salt would need to be reprocégsed for other reasons long before there would be 'any possibility of precipitating these fission products. The nuclear poisoning of the dissolved fission products would eventually cause the uranium fuel inventory required to attain criticality to become excessive. A discussion of this effect is given in the chapter on Reactor Analysis, Section 4.5, and Appendix C.6 and C.7. Although it is expected that improved reprocessing techniques will be developed in the near future, present technology would restrict fuel repro—. cessing to the following method. Contaminated fused salt would be drawn from the reactor system and the uranium recovered by the UF6 volatility process. The uranium-depleted, fission product-contaminated fused salt solvent would | then be discarded to a hot-waste storage system. The salvaged uranium wquld‘ - be dissolved in fresh, pure salt and returned to the reactor system. This same salt might also provide a convenient solvent for the uranium which must be added because of burn-up. A possible alternative which might be developed would be to treat the uranium-depleted, contaminated salt to remove some of the fission prohucts,, This would permit re-using the fused salt, thereby eliminating the consider- - - able expense of replaging the fused salt, and the problem and expense of storing the radioactive, contaminated salt. This treatment process might possibly be a solvent extraction type pro- cess, using fused salts, or it might be accomplished by cooling the salt to Just above its freezing temperature and precipitating some of the fiesion products. 3k, The economically optimum extent of fuel processing is determined by the relative costs of increased uranium inventory and the costs of processing. Based on the currently available processihg method outlined above; an economic analysis to determinelthe optimum length of processing cycle time was made and is outlined in Appendix C.6 and C.7. Based on the assumptions made in “ this analysis, the reactor should be operated at a uranium inventory-fission product equilibrium that would replace all the salt in the core system by a continuous process over a 19-year period. This means that unless a superior method of fuel reprocessing can be developed, it would be economically desir- able to perform no fuel reprocessing with the lifetime of the plant (20 years). g CHAPTER 3. STRUCTURAL MATERIALS 3.0.0 INTRODUCTION The major metallurgical consideration of a fused fluoride reactor is thg choice of a structural material for the reactor vessel and primary heat exchanger vhich will resist corrosion, be fabricable, and have satisfactory strength pro- perties at elevated temperatures. The nuclear characteristics of the structural materials in this reactor are not critical. Materials problems associated -with the sodium cooclant system and the steam power plant are important, but previous experience is available; therefore, the design_of these systems introduces no serious problems. | 3.1.0 REACTOR STRUCTURAL MATERITAILS The power plant design presented in this study is proposed for central station electric generation, and it would be desirable to have a reactor which would have an operating life eqfialAto the 20 or 30-year life expected from the steam plant equipment. Present technology is not sufficiently advanced to in- dicate that such an extended operating time can be obtained from fused.fluoride reactors, but it is mandatory that a material be obtained which will resist fluo- ride salt corrosion at elevated temperatures to the extent that these reactors can be considered for this type of operation. A homogeneous reactor of the contemplated design does not have as many materials problems as other types of reactors. The fuel is in solution and is self-moderating, and there are no fuel fabricating, cladding or moderating material problems. The reactor vessel is a low pressure container and does not have to be designed for high pressure operation. Since there are no structural materials in the high neutron flux region of the reactor, the effect of the nuclear characteristics of these materials on neutron economy is of seconda interest. Fused salts, however, are considerably more corrosive on materials that many other reactor fluids. Experiments have shown that nickel base alloys | the best corrosion resistance to fused fluorides. Inconel, nickel-molybdes alloys and pure nickel have been considered as possible structural material this reactor design. 3 1.1 Inconel There: afe considerable data on the corrosion of Inconel in fused fluorl at temperatures in excess of those contemplated in this design. An exmqflr Inconel corrosion is shown in the photomicrographs in Figure 3.1 which ilh trates the time dependence of the attack° Inconel corrodes in fluoride sal by a leaching of chromium frofi the surface of the metal. New chromium the| diffuses from the interior of the metal from which it in turn is leached. voids remain and concentrate at the grain boundaries; these voids are not | nected to each other-or to the surface. Corroqiofi data indicate that after the first 50 hours, the(depth of appearo to be linear, with a rate of between 3 and 4 mils per 1000 hours operation. The photomicrographs and corrosion data were obtained from fo circulation loops with 1500°F maximum fluid bulk fluoride temperature, a 1600-1620°F maximum wall temperature, 200°F temperature gradient, and 10, Reynolds number. The fluid circulated was Fuel No. 30 (Ref. 48, p. 97). There is also an initial impurity attack of approximately 3 to 4 mil which continues until the UF) -Cr reaction reaches eqpilibriuo and the chr content in the fluoride mixture remains constant (Ref. 48, p. 96, 97). T 37 tion will take place with each additional new charge of fluoride fore, the long.range corrosion rate will depend upon the number it the fluoride salt is changed. psion resistance of Inconel to fused fluoride salts is not satis- } reactor of this design. kel -Molybdenum Alloys nown Mickel-molybdenum alloy, Hastelloy B, has shown excellent fstance to fused fluoride salts (Ref. 50). This alloy, however, actory for reactor use because of its low ductility at elevated and because of its undesirable fabricating characteristicso el -molybdenum alloys are currently being developed which are in- tain the negligible corrosion rates, but which will have satisfactory ure and fabricating properties. ogion resistance of an 85-15 nickel-molybdenum alloy in a fluoride jermal convection loop is illustrated in Figure 3.2. These photo- jhow that there is essentially no corrosion or mass transfer in L a temperature of 15009E. Certain of these new alloys appear to .*ary'physicair;:zperties for reactor use; however, it was felt opment program is not sufficiently advanced to completely sub- expected excellent properties. These new alloys were not d this reactor study because of the philosophy originally estab- bit a reactor design using present day technology wherever possible. fckel Clad Stainless Steel opears to have excellent corrosion resistance to fused fluoride iic tests of nickel exposed i@ fltuoride:fuel.No. 30 Toryl00.Hours at own no measurable attack (Ref. 49). A small amount of mass transfer was * ."Ied-:.m the f;l.'rs-t 'r..lickv.al -thefmal"convection loop in which Fuel No. 30 u“flabjcirculatéd (Ref. 46, p. 63, 64). To confirm this single test, another :niékéi'loop was cleaned with Nafosfiénd ogerated with high pfi}ity fuel. No evidence of subsurfaceavoids,‘infergranular attack, dr mass fransfer’wés observed (Ref. 47, p. 73;;7h)., Thé“phétomicrOQraphs'in figure 3.3 show the absence of these effects in a nickel thermal convection loop operated for 500 hours at 1500°F with Fuel No. 30. No convection or forced circulation loops have beern constructed which . i1lustrate serious mass transfer properties of nickel, although the opinion has been expressed that nickel theoretically should mass transfer. Nickel does not have adequate strength to be used as a structural material; howéver, nickelqca# be clad satisfactorily on stainless steel. This duplex material will have the required cor?osion resistance and physical pro- perties for reactor cofistruction. Nickel clad stainless steel can be obtained at the present time as plate for the reactor vessel and as extruded tubing for the primary heat exchangers (Ref. 51). The inspection of éhe cladding is, how- ever, a problem. A metallurgicaliibfifi is obtained between the nickél cladding and the stain less steel backing material. Due to the intimate contact between tfie materials and the adherence of the bond, there 1s essentially no thermal resistagce° The efficiency in heat transfer propefties bf_this clad material is almost equal t& that of solid stainless steel (Ref. 52). The welding of this clad material will also present difficulties, but with the proper weld area preparation and use of special welding rods, this problem can be solved (Ref. 53). T-8227 1000 HRS. Figure 3.1 Changes in Attack with Increasing Operatlng Time in Forced- Circulation Inconel LOODPS. Fluoride mlxture circulated, fuel No. 30; maximun fluoride-mixture temperature, 1500 F; Temp- erature differential, 200 F:; fluoride-mixture Reynolds number, 5000. 6¢ - ———— 7-8798 HOT LEG 250 X T-8799 FIGURE 3.2 corrosion of 85-15 Nickel-Molybdenum Alloy by Fused Fluoride Salt No. 30 in & Thermal Convection Loop. Meximm fluoride mixture temperature, 1500 F Duration of test, 1000 hours. COLD LEG O T-4233 COTTOBLUM Va Wy ==y === Tused Fluoride aalt No. 30 in a Thermal Convectlon Loop. Meximm fluoride mixture temperature, 1500 F Duration of test, 1000 hours. 250 X p-L23l BOT LEG FIGURE 3.3 jckel in Fused Fluoride Salt No. 30 Corrosion of N in a Theruml.ConNection 1.00D . Maximum.fluoriae mixture temperature, 1500 F; puration of test, 500 houx's. COLD LEG “Th L2 The choice of nickel clad stainless steel as the structural and primar heat exchanger material for this reactor has made it possible to prepare a * preliminary design which could be constructed with a minimum of development work. -;t 1; recognized, however, that this duplex material has shortcoming and that a sipgle material with all the desired propertieewwould be superio Before a reactor of this basic deeign is constructed, it would be advisable to examine the nickel-molybdenum system Very carefully to determine if those alloys have the required properties to warrant substitution of the selectio of nickel clad stainless steel with a nickel-molybdenum alloy. 3.2.0 SODIUM SYSTEM MATERIALS Selection for the materials for.the'sodium system was based on the foll ing considerationsi 1. Ability to resist corrosion and mass transfer in high temperature sodium. _ 2. Adequate mechanical properties at elevated temperatures for extended period. 3. Stability under high radiation fields at elevated temperatures for extended period. : P L. Fabricability into the required shapes. 5. Weldability 6. Reasonable neutron'ebsorptionneross section. T. Availability and cost At the present time, the austenitic stainless steel alloys more nearly satisfy these requirements than other materials, except these alloys have a relatively high neutron absorption cross section. However, since in this design the heat exehangers carrying the sodium are outside the central core and in a low flux region, this is of little consequence. 43, There are three types of stainless steels, ndhely, Types 30k, 316 and 347, vhich have been previously used in sodium systems with satisfactory results. - There remains the problem of selecting the type that not only has the best mechanical properties at the elevatéd temperature under consideration, but also the type most readily weldable with no cracking of the weld and no embrittle- ment after long-life high temperature service. The material chosen fo contain the sodium in this reactor system was Type 304-L stainless steel (0.03% maximum carbon content). Although the three types of austenitic stainless steels show acceptable resistance to corrosion in sodium up to 1000°F, columbium stablized Type 3h7 has been generally favored. However, investigation at the Knolls Atomic Power Laboratory showed tha£ Type 304 stainlesfi steel is equally resistant to corrosion from sodium as Type 347 and that it has good strength properties and contains less strategic materials. No mass transport difficulties have been encduntgred up to 1200°F, which is higher than the sodium operating temperatures contemplated in this désignoi Since the austenitic stainless éteels‘appear reasonably corrosion resis- tant, namely, Types 304, 316 and 347, the strength versus ductility at elevated tempefatures should be given considerable weight in selecting the material. Type 304 stainless steel causes only slight loss in room temperature ductility due to carbide precipitation (Ref° 56) , whereas Type 316 stainless steel, under the same conditions and treatment, indicates rather severe decrease in room temperature ductility. ‘ N Table A.2 lists some of the pertinent properties of Types 304, 316 and \ 347 stainless steels at elevated temperatures. Therefore, from mechanical properties consideration, Type 30h'appears to be the superior. ”~ .. et . Lk, Since there will be many welds in the sodium piping and components, the corrosion effect of sodium on welded joints must be considered. The corrosion resistance of welded 18-8 stainless steels is affected because of chromium carbide precipitation at the grain boundaries in the heat affected zone. The formation of chromium depletes the surrounding.area of chromium and thus the area loses its corrosion resistance. The precipitation'of carbides can be pre- vented by: (1) reducing carbon content below 0.03 percent; (2) using columbium or other elements which form stable carbides with the carbon present; (3) cool- ing the heat affected zone quickly through the critical range (1 to 1-1/2 seconds from 1500°%F to 900°%F); or (4) heating to above 1750°F following welding 80 as to redissolve chromium carbide, followed by rapid cooling to retain them; Until recently the use of (2) above to prevent the precipitation of carbides had been considered the superior method; however, in recent years, the use of Type 347 weldments in tfiick sections for high temperature-long-time service has resulted in cracking during welding and embrittlement during service. It has been found that columbium free alloys do not become embritfled) even when Lhey contain normal amounts of ferrite (Ref. 55). Also, the tendency for cracking of the weld metal during welding is reduced with the columbium free alloys. To prevent carbide precipitation; the use of a low carbon content austenitic stainless steel appears now to be the better method. Rapid coqling and re- heating to above 1750°F are not practical methods for projects of this scope. The material which best meets the above requirements is Type 304-L. Therefore, Type 304-L stainless steel was selected as the structural material for the sodium systems. 3.3.0 STEAM SYSTEM MATERIALS Since the temperatures and pressures of the steam system for this design study are consistent with those of large conventional steam power plants today, 45, the selection of materials %or the steam and water sysfems presents no parti- cular problems. The main steam line piping would be fabricated from standard - Cr-Mo alloy steel pipe, ASTM Spec. A-335, Grade P-1]1 or P-22. The higher pressure boiler feed water, condensate, and extraction lines would be fabri- cated from seamless carbon steel pipe, ASTM Spec. A-53, Grade B. Other miscellaneous lines would be fabricated from standard seamless carhon stecl pipe. All valves would be fabricated from the same material as that of the piping in which they are located. h6. CHAPTER k. REACTOR ANALYSIS 4,0.0 INTRODUCTION The nuclear characteristics of this fused salt "burner" reactor are relatively flexible and are not restricted by the stringent requirements of breeding. In addition, fhe high solubility of UF) in the NaF-Zth fused s8alt system allows considerable variation in uranium concentration which provides a wide latitude for compromise between nuclear and engineering considerations. The basic reactor design consists of é vertical cylindrical core sur- rounded by primary heat exchangér tube bundles. The nuclear requirements of the reactor determine an optimum core diameter for minimum fuel inventory. The core dimensions selected represent a compromise between these considerations. In order to proceed with thé nuclear calculations, a preliminary arrange- ment of the reactor vessel was selected. This section of the repurt describes the reéégor analysis calculations that were performed to determine the optimum core size and fuel concentration and other nuclear rgquirements, A summary of the results is included in Table k4.1. 4.1.0 CRITICAL SIZE AND FUEL CONCENTRATION A preliminary criticality calculatioh was made by W. K. Ergen, OaknRidge : National Laboratory (Ref. 29), for a reactor similar to that of this project using a bare sphere, Fermi Age approximation. The optimum size and fuel con- centration for minimum fuel investment calculated by this method was 136 cm (4-1/2 feet) in radius with a fuei concentration of 0.0208 gm/cc of uranium-235. The fused salt used in these calcfilations was a 50-50 fiole percent composition of NaF and ZfiFh with the uranium dissolved as UFh“ b7, The reactor contains no moderating material other than the fused salt itself. The atom having the smallest mass and existing in greatest abund- ancé in the fused salt to be used in this reactor is fluérine, Fluorine; then, is largely responsible for the moderating characteristics of thé fused salt. With an atomic weight of 19, fluorine is a relatively poor moderator compared with the lighter elements usuaslly used as neutron moderators. It was suspected, and later calculations have shown, that this reactof was not " a thermal reactor, but that a high percentage of fissions would occur from epithermal neutrons. .Hand calculations for a bare sphere with th?ee energy groups indicated that approximately 50 percent of the fissions would involve epithermal nefitrons} ~With these preliminary hand calculations as background, a series of three energy group, three geometry region (3G3R5 calculations were performed on the. Oak Ridge National Laboratory digital computer (ORACLE) .. The energy group limits were_éelected for reasons given in Appendix C.l. The intermediate group extends from 0.5 ev to 100'ev,_.The fast group extends from 100 ev to 2 mev. The central core of the reactor is region 1; the lrinch.‘ " thick baffle surrounding the core is region 2; and the heat exchanger and downcqmgr volumes gre homogenizedlfor region 3. The phygical properties and nuclear parameters were based on a salt of composition 43 mole percent ZrF) and 57 mole percent NaF (Section 2.1.0). ‘The fuel is highly enriched uranium-235 (over 90% U-235). The fast and in- termediate energy group nuclear cross sections were obtained from the "eyewash code” (Ref. 27). Thermal cross sections were obtained from BNL-325 (Ref. 30). The detailed,détg and calculations for the code input are given in Appendix C. In order to minimize the number of computer calculations, 1t_was_de§ided to vary only the radius of the central core and the fuel concentration. Other 48. parameters such as the height of the reactor and the thickness of the heat exchanger were held constant. Since previous estimates indicated a h-1/2 ft radius core to be optimum, a 1C-ft core height was chosen to provide com- patible arrangement with this core radius and the heat exchangers. The heat exchangers contain a considerable amount of fuel, acting somewhat as a reflector to the central core, which makes 1t desirable to have a central core height larger than the central core diameter. Calculations were per- formed with central core radii of 3, 3-1/2, 4, k<1/2, and 5 feet. For each radius, calculations were made for fuel concentrations of 0.005, 0.010, 0.020 and 0,030 gm/cc of enriched uranium. | The above information and the data in Table C.2 were coded (Ref. 28) for the electronic computer (ORACLE). The electronic computer calculated the multiplication constant (k) for each radius and fuel concentration (Figure 4.1). *From these data, the optimum size and fuel concentration were determined (dis- cussed in Section 4.3.0) as a 3-1/2 foot central core radius and 0.0088 gm/cc of enriched uranium-235. A re-ecvaluation of the height of the central core is probably desirable in light of the smaller central core radius. In the final design, the height is shown to be 8 feet, which is probably a more com- - patible figure. A concave head and bottom are used to obtain this reduction in height and to direct the fused salt flow (See Figure 5.9). However, the nuclear, heat transfer, and other cfilculations were not revised to reflect this change. | In order to check the inaccuracy introduced by using the l-inch thick stainless stee; and nickel baffle plate as a diffusing medium (region 2), two check calculations were made. One calculation considered the Baffle as an absorbing "shell".between the central core and the heat exchanger regionms, and the other eliminated the baffle plate e?tirely. The multiplication 49, 1,5 —t T - ISI‘I-JOI 1 - = 16" 1.,.'-. T ]'l.LOI“I- R O O 1 3]_6" _ _ ! 1.3. I' o~ Y /'4, > ’A3 '_O" / ~ A |/ A4 y o 1,2 4 A /7 =S y ™ T :—' ] __101? 4 {1 mmb- mel: / EEb= v i - J WREf By / ¢ — st 1.0 : T & 0,005 0,015 0,020 0,025 0,030 | - / HEEL !.!!2!_3J' HHHHH ;:gr 7 Concentration U - gm, /cc. E b 7 Elb= " e 9.9 7 Mg i g AV Y 1L 104N (T . 0.8 / ] ] - 41 0.7 I/ I T O S T () [ s ) 5 D g 1 = ) S I L) VS i . Al = PR Gl 1§ L Figure lj,1 Variation of Multiplication Constant k, E with Concentration of 121235 and Central ' Core Radius for the Hot, Clean Condition [ ] o 7 [ (MA“T =1 T T QLIML. m m a - ;4 3 6 B B WA M o | S Fiieinenreneet el MU o - It Y i 0 LI 5 L P9 = CESEmSiEEiEEmEaatiic o 1 3 A (i =1 i I an 0 S B 2 TR ] < g &8 _.._n G AN R S ||TL|,11 1 | 1 (] A AEERE B 1 T = ° N WS , 1=y T 1 1 h- T T T i \ i \ ' 1 5 _ T SHNK ! w WA W i T | | * b ! = .y;th. T T _ | | N | o P ] L 1 I | S | u 1] ? O T =, Lol I i 1l las ‘4- | I s sunnsuun s wd ol R 1 I T 5 um i 3] T ; | i)} L | ,Ai L] s ] ) T L, ! T I _ L% [ m SEREENNERSINES S BEGESERaRE iy Rh [ i N eSS o b = AR ENREN SN TR L SO REEE W BE ! i _1 N ,,A_fi A IR S el B 5 S e S ! L 6 (M ARG NS S i L i B SR EERE 350! B % M BF BN N T SR 3 SRS EER R S REEEan Ran aU SR Euses rannans & SONER e S0 88 Ses ases nnses p spyeeses o 1si R R T A ] 1 i3 I 1 0 2 (A O SR SIS T T 0 AR T = L.IJ L@T. 680 5 SR T R , , o : o T : I | T 1 * i I A i T | T [ I + 7 +—T T T o T g L | L |1 | i T i 1 T I § i A B 1R ] W T D PN S Y T BsoamsmEs=s eShianEnso = meass - ; — -t - 3 r _ + P LY : .LM\M H\ i 1 _ _q ; A7 s i = ||. _ i ‘ i 3 EwEEmsannss A : ; i SIS N 1 it | 1 M P A0 B ,.. i BEE £ | e | i M2 TP S | 5 == G & RSN NSE R SEEN 2 B B ; i i ] i W e R T M N 08 I i G I R \mN 0 oo 50 2o o Y O L — SR a1 T R z 11 e v i i 2o | AR R % 1 H 0 (B 5 1 S 1 O e Y R e N 028 (S B S BRI s g 1 L T T T Y T __ ! Zeee fO g .,LJLOFII\\ ,\fie S ~ O et N B & 1 h ! I ) | BT N | I A L.mmmp .wn ¥ -.i. d.oaa«uo 5.8 ‘308 TeI0L| ,,_ flu.m w | 1 I 51. constants (k) and flux plots obtained by the three methods of treating the baffle plate were almost identical (Table C.3). It was therefore inferred that using the baffle plate as. a region did not in itself introduce any appreciable error. 4,2.0 NEUTRON FLUX The electronic computer (ORACLE) calculated the neutron flux distribution for each encrgy group as a function of radius (Figure 4.3). At 600 MW total power, the fast neutron flux is approximately 1.2 x lO15 neutrons/cm2 sec.; the intermediate neutron flux is approximately 6 x lOl,+ neutrons/cm2 sec., and the thermal neutron flux is approximately 1.6 x lOlu neutrons/cm2 sec. (Appendix C.4). These values, coupled with the fission cross section for each group, show the fissions to be distributed among the three energy groups as follows: 9 percent in the fast group; 27 percent in the intermediate group; and 64 percent in the thermal group (Appendix C.5). This means that this reactor is what is often termed an "intermediate" reactor. A nuclear calculation was prepared to utilize the 30 energy group "Eye- wash" code on the UNIVAC digital computer. Unfortunately, the results of this calculation were not available to be included in this report; but should be shortly thereafter (Ref. 40). The results from this more detailed calcu- lation will serve as a check on the 3G3R code calculation results. Also, and the original reason this alternate calculation was prepared, the 30 energy groups will provide a much better indication of the neutron flux energy spec- trum in this reactor. This is of particular interest for the high neutron energies, to better establish the possibility of utilizing this reactor as a large flux source of high energy neutrons for purposes of radiation damage research. TTT T m FF M o _ P g 8 - - - g 5/ T . — 59 3 EEEEN N DM o] ‘HII_JJU | 1 I Wa EE . 3 ) @ L I a m 2a) ur ° ¥, @ % . I o p’ g @ 3 NN m H o .ur o - H 1 - o mm CIAM =4 ] I A A B 9% H o Z 4+ ..-... Q.J Sk 1] o~ I 1 & IT d _ ] = Lo B S 1 T g E g siEaise EeERaas | e & 13 HHHH ey S 1 @ ] EEEE ] | " o £ i [ =Th=] [} e ECTTRE g8 £ ERExau 2o © T ! A 4 W - 0 N0 L4 n Oflu!t D i m 3 “ s 1F8 £+ S - P> " = m — T ® [ 1 (- 7} © =5 1 ) = = .m - : h e mis g™ 0 =1 (3} el © ! ] smmmm | amaaas man e mav-ke 8 3 4 H 2] | ,I_ll, .-u I | - & | B 74D 1 & i == 2= N s Be ,, fim =1 b T i L [ 4 . i =T : A g 1 | wE J\m ~ 2 o 5 3 2. o} [} p. W (T X .Q 4 ERNENEE P t ax M\OO ; ! L mv. “wh‘ & ~ “ N H gL &8 H = : u_foao %%& 5 7 R - aap 4 T o T g v TS e B ] L aw m g 1 EE 0 v 7 B mw T = " & T‘w 8 By — L1 o - . m H R T & fa R \ am B 1= - - 55 A i Sl ] 3 _ N SRR A e T | ! 0 0 0 20 L - e 6 6 ) (R : ! Erte \‘J_vj%m nw EEEE MT t o o Mlm _Tm ] 0..71 & ] M EE ~ 1 i T e e T = B ,.wwfi 9883 Jo \._.Qoe.uomfi ow Woua.n.flruo ..H.flflm + . | ] I | (I R N o ] T O d | | N I ,_ L ) 1 6 I [ 1] 4.3.0 FUEL INVENTORY The critical uranium concentr;tion for each central core radius was determined by the results of the three gfoup, three region calculations. ‘The critical uranium concentration decreases withran increase in the central core radius. The volume obviously ;ncreases with an increase in the central core radius. The combined effect fésmlts in s minimum uranium inventory required for criticality at some value of central core radius. This vériation ‘of critical mass (uranium inventory) with central core radius is shown ifi Figure 4.2, | For 600 MW rating, the heat exchanger volume, and ihus the fuel volume in the heat exchanger and downcomer; will remain nearly constant. The fused salt solvent inventory is decreased by a decrease in central core radius. Therefore, the central core'radius was chosen slightly smaller (3-1/2 feefi) than that for the minimum fuel inventory indicated (Figure 4.2), in order to allow some effect of the decreased salt inventory. The total uranium-235 fuel inventory in the entire fuel system for the hot, clean critical condition 18 approximately 240 kg. 4L.4,0 FUEL BURN-UP The fuel burn-up at 600 MW will be about 780 grams of uranium-235 a day. The fuel would be added on & temperature regulation basis to maintain an average operating temperature of 1125°F. 4L.5.0 URANIUM-233 FUEL Highly enriched uranium-235 is the fissionable fuel utilized in this -design study. The fact that a large fraction of fissions is caused by epi- thermal neutrons suggests that it may be desirable to use uranium-233 as the fissionable fuel instead of uranium-235s Uranium-233 has a uniformly low L value for the ratio of capture-to-fission cross section (oC ) of 0.1 over all neutron energles. This,méans that a considerably smaller portion of non-fissioning neutron absorptions occurs in uranium-233 than in uranium-235, particularly for neutrons in the intermediate energy range. The advantages that would be. gained by using uranium-233 the fuel, if_availableg are that - the uranium inventory required would be le:!!\hid the energy release for the uranium burned-up would be greater. 4.6.0 FISSION PRODUCT POISONING Fission products and corrosion prodficts will accumulate in the fuel solu- tion and will add considerable nuclear poison to the core. This will require an increase in the uranium fuel inventory and could become a significant economic consideration. An evaluation of the probable magnitude of this effect was made, and is outlined in Appendix c;.6° Based on the reference source utilized in this analysis (Ref. 33 and 34), and the assumptions and approximations made to adapt the data of this refer- ence to this design study, the following effects of fission product build-up are predicted. The nuclear poisoning of the non-gaseous fission pfoducts only (no corrosion products) will asymptotically approach a negative reactivity of about 0.1 during the lifetime of the reactor. If corrosion should be appre- ciable; the relatively high absorption cross section of the nickel atom could appreciably change_the rate and magnitude of this poison build-up. From Figuré L,1, it can be seen that a negative reactivity of 0.1l would require an increase in uranium concentration from 0.0088 gm/cc to about 0.0115 gm/cc to override this poison and maintain criticality. This is a{ 31 percent increase in uranium inventory from 240 kg to 314 kg. 03 Some form of fuel reprocessing could reduce the uranium required to over- come the poisons. This possibility was investigated (Section 2.6.0., Appendix € C.6 and C.7), and it was determined that it would not be economical to provide - for fuel reprocessing as limited by current technology. o The increase in uranium inventoys. to override fission product-poisoning would then progress gradually ov# life of the reactor, being added as . required along with the uranium added for burn-out. L,7.0 TEMPERATURE COEFFICIENT OF REACTIVITY The temperature coefficient of reactivity of the reactor was not calcu- lated, but is believed to besufficieqtly negative for the»reactor to be stable. This is based on the fact that a reactor using a similar fused salt - fuel has been operated, and this reactor had a negative temperature coEfficiefit of reactivity of approximately 5 x 1072 Ak/k per F (Ref. 39), p. 1009). The fact that this fuel has a desirable temperature coefficient of reactivity is also indicated by the slope of the curves in Figure L.1. 4L.8.0 DECAY HEATING The decay heating of a reactor during shutdown is dependent upon the amount of negative reactivity injected into the reactor_and the delayed neu- trons and gamma decaf rate.. In a reactor with control rods, a considerable amount of poison can be inserted very quickly. In the fused salt reactor, however,‘no control rods are present, and the ‘amount of negative reactivity achieved depends upon the magnitude of the temperature rise above the average L operating temperature. In the event of complete loss of pumping power, the reactor must be shut down and the decay heat removed. 1f it is assumed that no heat is removed and the temperature is allowed to rise, the power generation would decrease rapidly, 1 56. and the reactor period would level off to an 80-second period in a fraction of a minute. At this time, the power generation of the reactor would be re- duced to a few percent of operating power. The temperature rise under these i conditions may be several hundred degrees within the first minute. This condition of rapid temperature rise is undesirable, if not intolerablg_° Imme- ~ diately following the loss of pumping power, some heat would be radiated from the reactor vessel; some heat would be removed by the reactor vessel surface cooling; and some would be removed by the momentum of the sodium in the primary loop. A detailed calculation of the flow characteristics of the primary and intermediate sodium loops after l&ss of pumping power was not made; however, a simplified calculation showed that the inertia of the fused salt in the reactor, and the sodium coolant in the primary and intermediate loops is sufficient to maintain the flows in the turbulent range for approximately one minute. This would be adequate to carry away a considerable amount of - decay heat. and limit the fused salt temperature to a safe value. o7 TABLE 4.1 SUMMARY OF REACTOR ANALYSIS DATA 1. Dimensions of reactor central core: Radius - 3 1/2 feet Height - calculations based on 10 feet 2. Critical fuel concentration: 0.0088 grams/cc (for hol clean critical) 3. Power density: 55.2 watts/cc at 600 MW 4. Flux and fissions for each energy group: _Flux Fissibns:l Fast {100 ev to 2 mev) 1.2k x 1017 neut/cc-sec 9% Intermediate (0.5 ev to 100 ev) 5.96 x 10M* neut/cc-sec 27% Thermal (up to 0.5) 1.62 x 101% neut/cc-séc 64% 5. Fuel Inventory: 240 kg (for hot, clean critical condition) 314 kg (for hot, fission product contaminated condition) 6. Fuel Burn-up: 780 grams/day at 600 MW T. Temperature coefficient of reactivity: Not calculated; but reasonable assurance that it is satisfactory. 58. - CHAPTER 5. REACTOR AND PRIMARY HEAT EXCHANGER DESIGN 5.0.0 INTRODUCTION - The basic design of the reactor consists of a vertical cylindrical pressure tight vessel to contain the homogeneous circulating fuel. The size of the reactor depends on the nuclear characteristics of the fuel and the method of removing the heat from the circulating fuel system. Critical size and fuel concentration calculations have been discussed in Chapter &, based on the design concept of using primary heat exchangers internal-tp the reactor vessel. Several arrangements were considered prior to the | selection of the internal arrangement and reasons for this decision are dis- cusesed 1n the subsequent sections. 5.1.0 PRIMARY HEAT EXCHANGER The function of the primary heat exchanger is to transfer the heat from the circulating fuel to the primary sodium cooclant loop. The design of the exchanger is based on a heat removal rate of 2°Oh8 X lO9 BTU/br for both &ides of the exchanger with mixed mean fuel temperatures of 1200°F entering and IOSOQF leaving, and mixed mean sodium coolant temperatures of 1000°F entering and 1150°F leaving the exchanger. The minimum fuel temperature has been arbi- trarily set at lOSOQF, approximately 100°F above the fuel melting point to allow a temperature drop of 50°F‘for fuel processing and an additional 50°F for cold spots in the system before reaching the solidification temperature. ~ - The upper fuel temperature was set at 1200°F, which was considered the maximum allowable for tolerable corrosion and mass transfer rates for a long life power reactor. The sodium coolant temperatures were chosen to obtain a practical 09 e, log mean temperature difference across the heat exchanger which assures reason- able flows and heat exchanger surfaces. These temperatures provide sufficient margin above the fuel melting temperature to prevent operational difficulties ° from possible cold spots caused by poor.distribution in the flow pattern through the heat exchanger and reactor. 5.1.1 Internal vs. External Arrangement In determining a practical arrangement for the reactor and primary heat exchanger, the following two basic design concepts were considered: (1) Use of internal primary heat exchargers with associated piping and pumping equip- ment integral with the reactor. In this design, the fuel remains in the reactor vessel and the heat is transferréd from the fuel circulating inter- nally to the sodium coolant piped inside the reactor vessel. (2) Use of external primary heat exchangers through which the fuel is continuously cir- culated outside the reactor with piping and pumping equipment adjacent to, but external to, the core vessel. Each arrangement offers certain advantages and disadvantages, and the final selection depends on an extensive evaluation of the economic operation and design factors involved, and their effect on the design chgracteristics of the reactor and system. - The important advantages of the interfiél primary heat exchanger arrange- ment over the external arrangemeflt are:; 1. Lower fuel holdup because external fuel piping eliminated. 2. More compact arrangement résulting-in reduction of requirements for shielding, piping, containment, building and services (i.e., crane handling equipment,; etc.) 3. Less pumping power because of reduced piping friction losses. 4., Provisions for electrically heating the fuel containment vessel and fuel piping simplified. These 5&vantages must, in a complete economic and feasibility study, overbalance the following apparent disedvantages: 1. Difficulties in maintaining and replacing heat exchanger tube bundles. 2. Structural problems in supporting integral tube bundles. | 3. Higher radioactive sodium loop caused by the higher flux level inside the reactor increasing the desirability of an extra sodium lcop. k. Problems in fabrication of the larger reactor vessel and complexity of heat exchanger tube layout. 5. Difficulties in locating leaks and isolating the defective tube bundle. Time did not permit a complete anslysis of this problem. However, based on preliminary studies; it was concluded that a design employing internal heat exchangers and pumps could be developed wvhich would result in an economical and practical arrangement for a high power reactor. Therefore, the design of an internal arrangement was pursued from the outset of this study. 5.1.2 Internal Arrangement Several schemes for the arrangement of the internal primary heat exchangers were‘considered and are schematically shown on Figures 5.1 and 5.2. In all the 'schemes proposed, the requirement of counter-flow heat exchamge was followed in order to minimize tube surface and to obtain the best heat exchanger character- istics possible. A. Separate Upper and Lower Header and Single Pass Arrangement (Filg° 5,.1) The tube bundles are located in an annular space around the core section with baffles to direct the fuel flow down through the exchanger on thg outside of the tubes. The sodium cooclant is through the tubes. The tube bfindles ter- minate 1n concentric toroidal headers; with the coolant outlet header located at the top of the reactor vessel and the coolant inlet header locgted at the bottom. This arrangement imposes several apparent disadvantages. 61. 1. To obtain adequate heat transfer surface using practical tube sizes and a reasonable number of straight tubes would require tube lengths considerably longer than the optimum length dictated by the nuclear requirements for the reactor core size. 2. Use of a lower header would necessitate many large penetrations in the bottom section of the reactor vessel, which is undesirable and should be avoided if possible. 3. This arrangement requires pumping the high temperature fuel, which tends to decrease pump life because of increased corrosion rate at the higher temperatures. Other design difficulties inherent in this arrangement are thermal stress and expansion problems between the two headers of the exchangers, and the large number of tube-to-header welds required to permit a practical length. To overcome these disadvantages, Scheme II was devised and is the basis for the final design. B. Combined Header and Two-Pass Arrangement (Fig. 5.2) The tube bundles consist of U-tubes with the coolant inlet and outlet headers located annularly and concentrically at the top of the reactor vessel. The tube sheets have the shape of truncated circular sectors. The fuel makes two passes on the outside of the tubes, with the flow directed by concerntric baffles into the suction of the fuel circulating pumps which are located on the top outer periphery of the reactor vessel. 5.1.3 Basic Design Criteria In establishing the design, certain criteria were followed to take advan- tage of the present state of development of pumps, materials and heat exchangers. 1. The pumps would be positioned vertically and located symmetrically at the top of the reactor and would pump cold fuel. This would minimize seal problems and provide convenience in arrangement. Pumps would be sized practically. 2. Annular vertical U-tube heat exchanger bundles would be considered with duplex tubing, with 304 stainless steel as the structural material and clad with type "L" nickel on the outside to resist the corrosive fuel. Type 304 SS provides compatibility with the rest of the Na coolant circuit. -62- ORNL-LR-DWG-15339 ENRICHER LINE-— SODIUM AXIAL FLOW PUMPS T —~ SODIUM HEADER SODIUM SODIUM TO HEAT EXCHANGER TO HEAT EXCHANGER FIG 5.1- REACTOR ARRANGEMENT WITH STRAIGHT- TUBE HEAT EXCHANGER PRIMARY MEAT EXCHANGER TUBE BUNDLE BAFFL FIG. 5.2- REACTOR 63. ORNL-LR-Dwg. -18172 GRAPHITE REFLECTOR ARRANGEMENT WITH U-TUBE HEAT EXCHANGER PRIMARY SODIUM COOLANT OUTLET PRIMARY SCDIUM COOLANT INLET FUEL CIRCULATING PUMP 6. 3. The flow pattern through the reactor and heat exchanger would tend to maintain the circulation of the fuel in the event of complete loss of pumping power. k. The heat exchanger tube bundles would be sized to minimize fuel holdup and the number of tubes while maintaining practical fuel and sodium coolant velocities for reasonable friction pressure drops and heat transfer coefficients. This requirement necessi- tated a parametric study for the primary heat exchanger. 5.1.4 Parametric Study A parametric study was performed for the primary heat exchanger to deter- mine the optimum tube size and number of tubes for the required duty. The study was based on varying tube outside diameters with different ligaments (minimnm distance between outside diameter of adjacent tubes) for a number of cases, utilizing the following design conditions: Heat exchanged 2.048 x 107 Btu/br Fuel inlet temperature 1200°F Fuel outlet temperature 1050°F Sodium coolant inlet temperature 1000°F Sodium coolant outlet temperature 11509F Tube material: duplex tube, Type ALSL, 304 stainless steel (.06 € max) clad with Type "L" nickel Tube thickness: total 0.065, SS 0.042 in., Ni 0.023 in. Fuel properties based on: 50 mol % NaF - 50 mol % ZrF) Justification for the selection of these design conditions has been dis- cussed previously. Tube sizes of 1/2 in., 5/8 in., 3/4 in. OD each with ligaments of 1/4 in., 3/16 in., and 1/8 in. were studied, using a staggered arrangement with triangular pitches of tubes in bundles. For the purpose of this investigation, average values were used for the physical and thermal pro- perties of the fuel, sodium and structural materials. This simplification was considered valid over the temperature ranges involved. The complete derivation 65. iof the formulae for the study is outlined in Appendix D; however, in the development of the expressions, the foilowing basic_eqpétions-were used for establishing the applicable heat transfer coefficients. Tp simplify the study, the internal heat generation in the eirculating - fuel and sodium caused by delayed neutrons and gamma heating was neglected. Therefore, if a uniform wall heat flux is assumed, the Martinelli and Lyon equation (Ret. 15, p. T3) is applicable for determination of the sodium coolant tube side film coefficient, hdaT7%0.025 (Pe)°°8 (5.1) , k- . _ To establish the shell side film coefficient for the fuel, the Dittus-Boelter relationship (Ref. 6, p. 219) was used, 4, = (0.023) (Ry)0-8 (p.)0-% | (5.2) . | Evaluafion of fieynolds numbers for the various parameters studied indicated values‘in the fange of 1000 to 4200 on the shell side. For a-fube side calcu- lation, Reyndldé numbers of thip order 6f magnitude.would indicate that the flow was in the transition range between laminar and turbulén£ flow. Keén (Ref. 20, p. 137) indicates that for baffled heat exchangers, there is no dis- continuity in flow at Reynolds numbers near 2100, such as occurs with:tu§e side flow, and that the character of the flow on the shell side cannot be evaluated on the basis of the Reynolds number only. Further, this,referenée indicates fhat at Reynolds numbers‘as low as 100, fully turbulent flow can be developed on the shell side of a baffled heat exchanger. The equation given , by Kern (Ref. 20, p. 137) for shell side film coefficients gives values cfin- siderably larger than obtained from the Dittus-Boelter relation. Therefore, 66. use of the Dittus-Boelter equation results in conservative values for design, even though the bvaffling geometry is unknown. Calculation of the tube wall resistance was based on averaging thermal conductivity of the tube at the hot and cold mixed mean sodium coolant temperatures. The over-all heat transfér-caefficient was obtained by combining the indivi- dual resistances for the duplex tubes and film coefficients based on external tube surface, 1 =1 +Rys +Rsg+ 1 {5.3) Uo Bbs bya The required heat. transfer surface was determined by using the equation for ~constant over-all heat transfer coefficient (Uy) and counter flow adiabatic heat exchange between fuel and sodium (Ref. 6, p. 190), q = Uy S (At,;) {(5.4) For this study, the fuel and sodium flowse, the logarithmic mean temperature difference, and the yuantity of heat exchanged are fixed by the design comditions. It is possible by proper arrangement of these equations in conjunction'with the one dimensional steady state continuity equation; ws: AV (5.5) to develop simultanecus equations for the dependent variable, the over-all heat transfer coefficient (U,) with appropriate constants (C) as fumctions of the number of tubes (N) and length of tubes (L). The two equations developed in Appendix D are then, 1 .SR+ 1 4__1 (5.6) U, - - c/N'® . c¥cenB | 67, and U NL=zC (5.7) where the C's are independent constants number of tubes N = L = length of tubes R = sum of the individual duplex tube resistances For an assumed length (L), these equations are then two simultaneous equations in two unknowns which cen be solved graphically by plotting the over-all heat transfer coefficient (Uo) against assumed values of N for each case of tube diameter and ligament under consideration. For the equations to B L be satisfied; there exists only one value of N which is determined by the in- tersection of the curves of the two equations. The required number of tubes. (N) to give the necessary heat transfer for each tube size and various ligamnets is shown on Figures 5.3, 5.4 and 5.5. The pressure drops for both shell and tube side of the exéhanger can be . h BT ot R L E LN N ARy IR determined from the Fanning friction pressure drop formula, once the requiréd number of tubes for a particular tube size and ligament 1a-established° ap (pst) = (2.0+f L)V P (5.8) de 2g 1M The friction factor f was determined from Moody's chart (Ref. 16, Fig. 15) for smooth Pipe. The constant 2.0 in the above expression allows for inlet, outlet, and U-bend losses in the respective circuits (Ref. 16, p. 21 et sequi). The additional fluid pressure loss across the tube bundle supports on the shell side was not included since its method of determination is highly empirical and depends on particular tube layout and design, and, in additionm, it was considered that its contribution would not effect the selection of tube size and spacing. 68. A summary of the results of the parametric study is shown in Figures 5.6 and 5.7 and Table 5.1. These graphs ého?;cross plots for the number of tubes, heat transfer surface} fuel holdup'volfimé and total fluid phmp powef a8 a function of tube ligament ratio (ligament to tube OD) for the values of N, as determined from figures 5.3, 5.4 and 5.5. Examination of the results of this st;dy indicates that for minimum fuel holdup, reasonable pumping requirements and a practical tube arrangement, the selection of approximately 14,600 5/8-inch OD U-tubes with 1/8-inch ligament (0.20 tube ligament ratio), and an effective length of 20 feet results in the most serviceable and practical primary heat exchanger design. In determining the optimmm tube size; it is desirable to minimize the number of tubes consis- tent with reasonable pressure drops and fuel holdup volume. The fewer the number of tubes; the less the probability of tube or tube-to-tube sheet failures, 201.5 Design Considerations To determine the practicability of such a U-tube design in annular con- centric sections with a 3-foot, T-inch inner core radius; a tube bundle layout study was made using approzimatély 15,000 5/8-inch 0D tubes arranged on a 3/4finch staggered equiléfiergl triangular pitch. The result of the study indicated that such an arrangement was feasible and is shown in Figure 5.8. It was finally concluded that to provide additional wargin in the design for contingencies, 24 tube bundles, each with 650 tubes to carry, or a total of 15,600 tubes with an effective tube lengfih_of 2i‘feet would be specified for the primary hegt exchanger., | Due to the difference in areas caused by the concentric sections,; some crossing over of the tubes is necessary to retain the same tube density on both sides of the tube bundle. Although this increases the fabrication and 69. \ \ 1100 \ L UL = G \ [ B T \ -y_‘-—lOOO : - AT ET = g gy 3 &~ \ L Figure 5,3 Primary Heat Exchanger||- - | LI TN HE (e (5 0 v e 5 1) A S T [ oL | \ Parametric Study [ \ LTI TT T T T Ll SR T At 1/2 Inch Tubes & 900 \ fi S ‘ Q Basis: External Tube Surface 43 \ - o [ \\ ; TAY .. ¥ [ 4 \ o e \VEA Ep 5 o 800 N EEEA \ —~ 1/2" OD - 1/8" Lig, e \ : Gy \ \ ® o N ) N 5 | \ \ [ 700 Gy : X \ \ N : N @ \ £ \ EH N A5 A\ \ N - 600 \ N Rl o -~ . N \\ — o & o . b: 2 N 1 SOO : DO \\ \\\\“ — 1/2" OD = 3/16" Lig.‘: . i _ _ 4 ) 1 s \\ 100! 1/2" OD - 1/L" Lig, £ N \\ \\\ - R 300 N S 10,0007 120,000 30,000 1,0, 000150, 00055100, 000 N'- i‘Iumber ‘o:f' Tubes | ||||||||| | | | 1 R R l ORNL=LR=Dwg,=1817k e | i = UONL = Cé | | : | - : CEE ::_N::*;:’ \ ey \ Figure 5,li Primary Heat Exchanger S Parametric Study E— = Y : Rl ‘ +H 5/8 Inch Tubes [T i o Basis: External Tube Surface A [ty e [l et \ VA | o 1 o \ B \VHEA s Q_,' . \ Bk b \ [ \ \ L O 5 \ EL O el c \ L @ T & EEELEE }45/8" 0D - 1/8" Lig, Bkl é: C \\ [, \ \\ \\ e \\HEEA N L1 " ® \ \\‘ mago N, = N A EIRENE, \ SEp S 3 \ N =z a P ‘\‘ R I o] A A N el ; ‘\‘\ T \ \ 8" 0D - 3/16" Lig,- :: Do: 1fl+ll Lig. = N T 5/ 3/ g 7] i : N\ AN : N N ™ :F . = i - 20,00 30,000 1,0,000 50,000 60,000 e N - Number of Tubes u 10 2 10 0 O O ) 1 O o e O 1 [ ORNL=LR=Dwg,=18175 T - 1000 b oo = 0y ! = \ & - » \ 2:-900 Y \ & i he \ ] g i e T £ 1800 4 . it vt \ Figure 5,5 Primary Heat Exchanger il ‘ = ) (T IS (| (U 8 Rl W T (O ) ) (| 1] '%4: UREE Parametric Study T g \ 3/lt Inch Tubes: L el : i : H— & —700. : S Basis: External Tube Surface R \ \HHA - 11 3 \ ||| f;," \ \ \ L} ) Q_'- . sERfes: EERLE N A0 o0 - 1/6" Lg. R 2afd e VEEEA \ : e \ yé 0 TS \ @ N 1 - 13 \ ) e oL A\ N e ga N = 1 ' \ muil- L 5 B S' i - A A N N um Doilfiéo \ \ - N3/l oD - 3/16" Lig. 3 i OD - 1 _;.T ng. e NS 300 h . AN = N - SN 200 L % 0 10,000 111-20,000 - 30,000 |+ 1,0,000 |1 50,000 N - Number of Tubes T (e e o e P o ORNL~LR=Dwg,=18176 T2, F £ 0 1y 2g° S (4} [0) o G4 N 4 O + 0N T Al rMm =8 P o & A o OHir o ot o 16 P W.u eO-m &) NS g6 i) 298 g - @ m.flm 3Humm i @ > s == I () Q et m.w.m : 9_wuHm 0 =t & o HHH P i hs wm m_ lnw ; n AL W = K & B = 3 R & = 1 | Figure 5,7 Primary Heat Exchanger Parametric Study Summary - Variation of Pump Power, Fuel Holdup Volume and Heat Exchanger Surface EgfifiEIZOOHHWWrTIMIuITHTH with Tube Ligament Ratio and Size of Tubes s Total Fluid Pump Power HiEIN{HHIIHsE T I ssccl URNL g =181 . 13- B vs, Tube Ligament Ratio : ¥ @ 11800 | it . Ef = D3h00 i - 3 THEEHS /g , g 3 o8 op HE © SYetios e 0 2500 e 1t S, gt ~ 1 Bum Fuel Holdup Volume vs,! 5 &y 2000+ s ™ > ® FF L ent Ratio H ; H HHE D R z’} Y 1500 2:;: 0 e S i 10004 & : g HH o = O it 500 - it S 0 & i T © ! P R ! = SH180 o Heat Exchanger Surface / N H160 f e vs, Tube Ligament Ratlo i O - i e 'Ha)o 84 i < FHS HHH ér__ §L - _ BH HE120 & ) ‘ @ / ¢ > Raf- t \Y isess &::1010 ~y gfn ’ FEH80 T E-'I”'"‘HH saaes +H EEicogs satcisa - 9F 60 e : H 1 1,00 0.1 fiflo, 2HH#HHO, 3 0l 0.5 Tube Ligament Ratio (Ligament/Tube OD) TABLE 5.1 :7‘5%{ SUMMARY OF PRIMARY HZAT EXCHANGER PARAMETRIC STUDY Physical properties used for study: FUEL Absolute viscosity «¢ 190 lb/hr-ft Density ¢ 209 1b/cu ft Conductivity k 1.5 BTU/hr-ft-OF Specific Heat Cp 0.285 BTU/1b The study covered six arrangements using 1/2 type 304 stainless steel with ligaments of l/fi" SODIUM Type ALSL - 304 Stainless Steel Type "L" Nickel (.06 C Max) A 0.53 1b/hr-ft ¢ 51 1b/cu ft k 36.2 BTU/hr-ft-°F kK 12.9 BTU/hr-£t-°6° k 33.5 BTU/hr-ft-°F p 0.3C BTU/1b 5/8", and 3/4" duplex tubes with 0.023" clad Ni on outside of 0.042" wall s 3/16", ard 1/8". C 6 Reactor heat load - 600 Mw Temperature - Fuel 1200°F inlet - 1050°F outlet Flow rate - fuel 48.0 x 10> 1b/hr Sodium 1000F inlet - 1150°F outlet sodium 45.7 x 10 1b/hr Tube OD (in.) 1/2" 5/8" 3/4" Ligament (in.) 1/4" 3/16" 1/8" 1/4" 3/16" 1/8" 1/4" 3/16" 1/8" Tube Ligament Ratio 0.5 0.375 0.25 0.40 0.30 0.20 0.33 0.25 .17 No. of tubes required with 20 ft 60,000 32,500 18,000 50,000 26,000 14,600 46,200 22,200 13,200 effective length Velocity - Fuel outside tubes, ft/sec 0.50 1.3 3.6 0.50 1.3 3.4 0.46 X3 3.2 Sodium in tubes, ft/sec 5.6 10.3 18.6 27 7.2 12.5 2.6 5.4 9.0 Reynolds No. - Fuel 1310 2310 4150 1200 L 2350 4000 1070 L 2250 4050 Sodium coolant 5.95x10% 1.10x105 1.98x10° 5.34x10* 1.025x10° 1.78x10° 4.6x10 9.6x10" 1.61x107 Over-all heat transfer coefficient- 260 485 885 252 492 865 228 470 820 BTU/hr-sq ft Heat transfer surface 157,000 84,700 46,500 164,000 84,000 47,500 181,000 87,200 51,000 Pressure drop - Fuel, psi X1 .89 8.0 .13 .89 7.8 .13 .84 Gl Sodium, psi 2.5 1.2 22.0 .91 3.0 8.5 S 1.4 3.6 Pumping horsepower - Fuel, hp 1.8 14.9 134 2.2 14.9 130 2.2 14.0 102 Sodium, hp 163 468 1430 59.0 195 552 2k .0 91.0 234 Total, hp 165 483 1564 61.2 210 682 26.2 105 336 Fuel holdup in tube bundle, cu ft 2420 967 355 2470 950 365 2730 980 405 .flL ] /DIVISION LINE BETWEEN BUNDLES \ " e" """ ‘"‘ = 3" CLEARANCE "A ‘v‘v‘v.v‘u - S0 — \K 4 RN T | Q) % 6‘ e )‘ ’ "’ N QERERRON. 55 7B e % -‘ NN “""""" VeV VeV . 'A vvvvvvvvvvvv ’ ‘:I. ,// f/\\\/ \ &v’ / y \ .AV AV‘VAV VA"V‘V"‘~" "Aa ' (‘) Q XXX XAHKAXXAKAN 0‘ RQL) ‘°' e A4 NS i IR IRANX XA / ] 3 /)3/\&)&{\{‘\ % \/ / % ’ / g';x AA‘A VA .Ag" SE5K § N\ \ v’vo‘ "' c:, X «.v -.A" %) 9 / QQM. 0‘0(') “?‘“" £ 'o ‘ “ """M.‘.' "“ QL )‘""o"wo'w"o‘." A ‘:""" S __030. 0‘ N (,'g& Ao .-.0. 0.’ b = = ke , 0""‘% ‘ ‘ ‘ ‘.(veA% %AaAV‘A'A' é /i// |/ "’ “""A A’A"VA’ ' = ' T / 9L R / e NOTES - | TUBES IN AREAS A, AND A, ARE A CONVENTIONAL U-TUBE BUNDLE. FIG. 5.8 PRIMARY HEAT EXCHANGER 2 TURES INHRDA 8. SE OE 125 ‘ OF A U-TUBE BUNDLE WHICH DIVIDES TUBE LAYOUT AND CROSSES OVER BELOW THE TUBE BUNDLE CONTAINS 650 TUBES. BUNDLE A,-A, INTO AREAS B,. 24 BUNDLES REQUIRED T6. assembly problems;, the heat transfer characteristics are improved. Since the physical crossing of the tubes occurs in the lower portion of the bundle; the fuel flow pattern through this section will be irregular; improving the mixing 1 of the fuel as it passes through the exchanger. In order to obtain adequate cross sectional area for welding the tubes to the tube sheets, it was initially planned to increase the pitch from 3/4" to 7/8" and extend the tube sheet radially outward over the tube bundle to obtain the additional area; however, an actual layout of this arrangement in- dicated that the outer radius of the tube sheet became excessive and therefore this plan was abandoned. To solve this difficulty, the following various alter- nates were proposed: (1) place the tube sheets at angles from 45° to 90° to obtain the required additional area. This method of arrangement offers the disadvantage of imposing severe stresses on the tubes at the bend where the tube enters normal to the tube sheet. (2) Allow the fuel to flow inside the exchanger tubes and the sodium coolant ocutside with the tube sheet positioned in the vertical. This arrangement offers some advantages such a&s reductlon in the required heat transfer surface because of the larger over-all heat transfer coefficient brought about by the fuel film coefficient; the main resistance to heat flow in the system, being increased with the fuel inside the tubes. The difficulties involved in draining the tube bundle completely and problems in successfully cladding and inspecting the inside of the tubes ruled this arrange- ment out. (3) Finally, it was concluded that the most expedient solution to the problem would be to swage the tubes to 1/2" OD size Jjust before the tubes - . enter the sheet; maintaining the 3/4" pitch the full length of the tube. This allows a spacing of 1/4" between the tubes on the tube sheet with 5/8" OD tubes, which is considered enough for a successful welding procedure. No difficulty is anticipated in the swaging of the duplex tubes. e The final exchanger design comsists of 24 U-tube bundles with 650 tubes positioned vertically in the reactor vessel; each with an effective length of 21 ft. The size of the duplex tubes suggested is 5/8" OD with .O42" type 304 stainless steel wall clad with .023 type "L" nickel. Approximately a total of 20 tube spacers and support plates will be required in order to in- sure proper tube spacing and support. An arrangement with the tube spacers staggered on 12" centers will result in a practical configuration. The tubes will be swaged to 1/2" OD as they enter the tube sheet. The method for supporting the individual heat exchanger assemblies is discussed in Section 5.2.3. 5.2.0 REACTOR VESSEL The reactor; shown schematically on Figures 5.9 and 5.10, is basically a vertical cylindrical vessel with an over-all height of approximately 13 ft and a minimum diameter of 16 ft for most of its height. The diameter in- creases to 20 ft at the top of the vessel to allow for the installation of eight equally spaced fuel circulating pumps. In order to reduce fuel holdup, the minimum diameter is maintained for the height of the vessel; except at the location where the fuel circulating pumps are positioned. At these loca- tions, the vessel flares into protuberances to provide the necessary flow area for the pump discharge transitional sections. An annular plenum chamber to serve as a suction header is located at the top of the vessel directly below the pump base plate. This distributes and permits flow to all pump inlets in the event of possible failure of any single pump. Both the bottom and top of the reactor vessel are concave hemispheroidial sections to minimize fuel holdup, to allow for more effective length of the heat exchanger, and to improve the flow characteristics of the fuel into and 2% t FUEL CIRCULATING PUMP 8 REQUIRED | PRIMARY { * ’ L CCIOLA'NT TO PRESSURIZER, GAS REMOVAL, | PRIMARY * Y ? AND EXPANSION TANK //—SF)DIUM COOLANT . a: Q| V7777 b o B Z Bf 1B LR T I Sy H | '1 Q.. 1. X L\'\OPENING FOR T AN A O fq GAS ESCAPE RETURN FROM C c H o e : PRESSURIZER AND \ EXPANSION TANK [ \ by N N \ REACTOR CORE SWING CHECK VALVE N Bl 24 U-TUBE BUNDLES o 650- 5/8"0.D. TUBES EACH [ | { ‘ 4" beer— —— | 8- 27" - 7 2222722222 22272277 ELIELLLLL : -~ ‘O - GRAPHITE REFLECTORd' EN : £ o TN e : Pt e OPENINGS FOR . FUEL DRAINAGE FUEL DRAIN AND FILL LINE| F1G-5.9-SECTION (A-A) THROUGH REACTOR VESSEL —BL- + QMI-YI~INY0 V79241 ORNL-LR-DWG: 15228 A PUMP DISCHARGE COLUMN . QZ% SECTION C-C /'/. 24 U-TUBE BUNDLES 650-5/8"0.D. TUBES EACH RETURN FROM PRESSURIZER AND EXPANSION TANK TO PRESSURIZER,GAS REMOVAL AND EXPANSION TANK FUEL CIRCULATING PUMP A SECTION B-B STRUCTURAL MEMBER PUMP SUCTION PLENUM CHAMBER FIG. 5.10-PLAN VIEW OF REACTOR VESSEL —GL_ PRIMARY SODIUM COOLANT INLET PRIMARY SODIUM COOLANT OUTLET -80- i =S ORNL-IR-DUG: 15348 PRESSURIZER, GAS REMOVAL, AND EXPANSION TANK / \_ TO GAS OFF-< TAKE SYSTEM ~__ TO HELIUM == S 24 FUEL CIRCULATING PLIMP MOTOR PRIMARY SODIUM COOLANT LINES LIQUID LEVEL @ 1300F @ 1125F FIG: 5.11- PRESSURIZER, GAS REMOVAL , AND EXPANSION SYSTEM 81. out of the centrai core section. Use of the concave sections reduces the need for internal baffling and perforated plates to distribute the fluid flow,“ . To minimize neutron leakage at the top and bottom of the reactor, graphite blocks are packed into the concaved sections to serve as a neutron reflector. A side stream of the primary sodium coolant is used to remove the hesat caused by the gamma and neutron interactions in the graphite. Similar provisions could be uwmde fér cooling the feactor shell if further investigation indicates that the gamma heating in the reactor shell is excessive. ) 5.2,1 Shell Design -Since the fused salt in the reactor has arvéryAlow vapor pressuré at the operating température (,Ol.mm Hg), the system is pressurized with helium at a, pressure of only 10 psi (See Section 5.2.5). Because of this low pressure, - the shell thickness reqnired for the Qessel-is relatively thin as compared to most other types of high temperature reactors. By using a design pressure of 50 psi in the vessel, a maximum temperature of 1250°F of the fuel, Type 304 stainless steel as the structural material, and a joint efficiency of 80%,-a - . shell thickness of 1.8 inch is required for the reactor vessel. This was obtained byvapplying the ASME formula for cylindrical shells listed in Unfired: Pressure Vessel Code (Ref. 35, p. 113). To allow an additional margin of safety, the shell thickness 6f the vessel 18 designed for 2 inches and is fabricated from Type 304 L stainless steel with 0;03“¢ maximum. - The choice of Type 304 L stainless rather than Type 347 was made since it has been demonstrated that the possibility of incipient cracking:at the weld Jjoint is less brobable with Type 304 L. The weldability.of Type 304 L compares favorably with Type 347 as long as the carbon content is limited to 0.03 maximum. 82, The vessel is clad with approximately 0.20 inch of type "L" nickel at all points which are in contact with the fused fluoride salt to prevent corro- sion. It should be noted that in establishing the thickness of the vessel, no increase in strength was allowed for the nickel cladding. This is in accordance with the ASME boiler code. The sides and bottom cf the vessel will be fabricated from 2-inch plate into sections that will be welded to- gether to make the complete veséelo The assembly of the top of the vessel is more complex since the heat exchangers are seal welded onto structural members and the top ecover section. 5.2.2 Internal Arrangement The internal arrangement of the reactor vessel consists of three prin- cipal components: the central core section where mdst fissioning takes place; the'primary heat exchangers in which the heat of fission is removed; and the fuel- circulating pumps which provi@e the necessary head to maimtfiin continuous circulation of the fuel. By referring to Figure 5.9, the fuel circulation path can. be followed. Fuel flows upward through the T-foot diameter cylimdrical core;, outward radially at the top, and down through the first pass of the U-tube heat exchangers which are located in fhe baffled apmulus surrounding the core. The fuel reverses its direction at the bottom of the heat exchangers and flows up- ward thfough the second pass of the heat exchangers. At the top of the second pass, the fuel flows radially outward again to eight axial flow pumps located symmetrically around the upper periphery of the reactor vessel. These pumps force the fuel downward through an annular downcomer which surrounds the entire vessel. At the bottom of the downcomer, the fuel flows radially inward and ‘back into the central reactor core to repéat the eircuit. 4) <) 83. The internal baffling is 7/8-inch Type 304 stainless steel base plate clad on both sides with 1/16-inch minimum thickness Type "L" nickel. The inner and outer baffles of the heat exchanger are connected to the reactor vessel through internal structural members; not shown on the diagrams. The center baffle .is welded to each primary heat exchanger tube sheet channel and is supported at the bottom by the radial-support plafies of the heat exchanger bundleu- Leakuge and streaming of the fuel between the hot and ¢old legs of the heat exchanger bundles is minimized by using a tongue and groove arrange-- ment at the points where the center baffling ends between adjacent tube tundlies. In addition; special internal baffling is required to prevent fuel streaming between bundles in the voids caused by the radial structural members at the topr of the vessel which support the heat exchangers. Perforatiofis are provided at the bottom of the U-bend of the baffling to permit complete drain- age of the fuel, and at the void spaces around the fuel pump discharge transi- . tional sections to avoid overheating in this stagnunt volume. Eight fuel drain and fill lines are provided at the bottom of the vessel and are connected to. the fuelldump system through let-down valves and safetly devices. 5.2.3 Structural Arrangement . The reactor vessel is supported by eight vertical members embedded in - the concrete floor of the containment vessel. These members are welded to the outer structural ring beam which acts as the main support member for the reactor vessel. Twenty-four equaliy_spaced beams extend radiaily inward from the outer ring beam and are in turn welded to an inner concentric.ring beam. The resultifig design is a ¢ircular cantilever functional arrangement similar in appearance to a large wheel with 24 spokes between the hub and rim. 8l. The primary heat exchangers are located between these radial members. The heat exchangers are seal welded individually by radial and concentric seal strips. The radial sea; strips are located on each side of the heat exchanger channels and are welded to the radial beams running between the heat exchangers. The concentric seal strips aré located at the iwmer and outer periphery of the heat exchanger channels and are welded.to the concentric ring beams to complete the closure. These seal strips are also designed to restrain the heat exchangers from the upward force caused by the buoyant effect of the fused salt on the tube bundles. This design is unique in that it utilizes the inlet and outlet ch;nnels of the primary heat exchanger as an integral part of the reactor veaéel° - Ffirthermore, the primary heat exchangers can be removed by remotely cutting out the seal rings of any particular bundle and lifting the entire bundle out vertically after the primary coolant piping has been cut away. 5.2.4 Fuel Circulating Pumps The fuel circulating pumps have several special reguiremcnte that must be satisfied for a practical design. These requirements are as follows: (1) absolute leak tightness, since any leakage of the highly radioactive fuel would conteminate the area and would eventually cause a shutdown; (2) opera- tional reliability with minimum maintenance for uninterrupted; continuous service, since all maintenance must be accomplished remotely once the reactor has been operated; (3) be easily replaceable as a unit since plant shutdown is required for access; (4) vertical installation fof complete removal of- pump internals without necessitating complete draifiage of the fuel; and | {5) reasonable cost. | ‘ In selecting a design for the fuel circulating pump that would reason- ably fit the above requirements, two types were considered sufficiently 85, developed to warrant investigation for thié application: .(l) the totally enclosed or camnned rotor pump; and (2) conventional motor drive coupled to . the pump by shafting that has special provision for frozen seals to pre- vent leakage. The canned rotor pump met many of these requirements; however? such a pump has not been fully developed to operate with fused salt at temperatures in the required range of 1000°F to 1300°F, Thg bearings of a canned rotor pump are generally lubricated by the fluid being pumped and cooled by a secondary coolant. Application of a canned rotor pump to a fused salt system would necessitate that the rotor; sxnd consequently the electrical igsulation, operate at temperatures considerably ifi excess of those now considered prac- tical. Coolant passages would require judicious design to avoid freeze-ups in the coolant circuit. Further, the cost of development and fabricatiom would be considerable for materials suitable to resist the corrosiveness of the fluo-= ride salt. For these reasons, this type of pump was rejected. A veriical pump cbupled to a conventional motor drive with special adaptations of freeze seals and hydraulic, piston-type, self-lubricating bearings was selected as more suitable for this application. Calculations indicated that the total fuel flow is 32,000 gpm. This requires a total dynamic.pump head of 11 ftor 15 psi. This load is distri- buted among eight pumps; on the basis that this number resulted in a éymmetrical arrangement in conjunction with the 24 primary heat exchaqgers; and also that failure of one of eight pumps would not seriously reduce the capacity of the plant. Further, the impeller‘aize for a LOOO gpm pump is reasonable and fits into the chosen pump location. The required'drive‘motor. rating for each pump unit of 60 hp is practical. A value of 60 percent was used for the phmplefficiency in the design vhigh-is considered consgrvative for axial flow pumps. 86. To determine the pump classification; the specific speed was obtained from the following expression (Ref. 21, p. 296), Ns = I::) GPM - | | (5.9) where Ns 1s the specific speed; N is the revolutions per minute of the pump shaft, GPM 1is the capacity in gallons per minute, and H is the total dynamic head of the pfimp in feet. | Application of the above expression when considering a pump speed of 1750 rpm results in a specific speed of approximately 18,000, which is in the axial flow pump range. Use of an axial £low pump simplifies the arrange- ment of the pump discharge nozzle and transitional pieces since volute diffuser sections are not required and,; therefore, both suction and discharge can be axial. | In order to reduce the possibility of cavitation for the top suction pump design, it was decided to provide a pressurizer forrthe fuel in the reactor vessel. The'pressurizer will maintain the fluid level several feet above the pump suction. The location of the pump in- the reactor vessel is critical since the maximum shaft overhang tolerable is of the order of 18 inches. For this reason,; the pumping element is located as near to the top of the vessel as flafi permits. A new development in pump bearings, called the hydraulic, piston-type bearing, permits the bearing to operate in the fluid and is lub- - ricated by the fluid being pumped. This design would have application in the pump installation proposed, as it is specially designed for high tempera- ture corrosive fluids. To prevent leakage, a freeze seal is proposed. Essentially, this consists of a-section of the pump shaft where cooling coils 87. freeze the fuel around the shaft, thereby_preventing fluid leakage. These -seals have been developed and are considered practical. In order to prevent backflow through a pump which has been shut down; a check valve is required at each pump discharge. The check valve is located in the pump discharge downcomer and arranged so that it can be removed when the pump internals are withdrawn. The check value must offer a minimum resis- tance to flow sou tfiat natural circulation will not be restricted on complete eleectrical failure of the pump motor drives. The choice of suitable materials for the fiump internals is 1imited at the present time to Inconel. Experiments (Ref. S57) have been performed with an Inconel pump operating in a forced cireulation loop with fluoride fuel Noo 30 at a cold cperating temperature of lQBOOF and with a maximum loop tem-~. perature of lSSOOF° The pump was successfully operated for over 8000 hours, - when an electrical power failure forced a shutdown and restart was prevented by a pipe break. It was considered that the pump was cepable of operation o for many more hours without mechaniéal failure. The pump was equipped with a conventional seal of §ilver impfegnated graphitar against tool steel and with commercial ball bearings. This seal and bearing arrangement is different from that proposed above. On inspection of the pump impeller, it was con- cluded that the effect of corrosion and mass transfer on the pump materials vas negligible. On the basis of this experience, the use of Inconel for pump internals is considered satisfactory for the design and operating conditions planned for the fuel circulating pumps for this design study. A summary of the specifications for the fuel circulating pumps is given in Table 5.2. TABLE 5.2 SPECIFICATIONS FOR FUEL CIRCULATING PUMPS, (8) Type-Vertical-Axial flow B Capacity 4000 gpm . Temperature of salt entering pump 1050°F ‘ Total dynamic pressure head 15 psi Speed - Constant | 1800 rpm Pump horsepower | 56 hp Motor horsepower 60 hp Impeller outside diameter 10 in, Hub diameter - 5 in. Materials - Trim ' ‘ . Inconel Seal Freeze Type Bearing Hydraulic Piston Type éfigo 5.2.5 Pressurizer (Gas Removal and Expgnsian System The reactor vessel is pressurizéd to a maximum pfessure of 10 psi. The pressure is limited to this value to keep the design pressure below 50 psi in - the reactor vessel. Helium is used to pressurize the system because of its- negligible solubility and inertnesé in the fuel sclutiom. Pure helium does not become radiocactive and therefore can be vented di%ectly to the atmosphere. (Ref. 15, p. 115). The pressurizef syéfiem is required for several important reasons: 1) To maintain a positive flooded suction head on the fuel-circulating pumps to prevent pump cavitation and vapor binding. 2) To provide gas pressure for the fission product gas removal systemok 3) To minimize the possibility of the "snow problem" described in Chapter 2. 4) To reduce the tendency for formation of fission product gas voids in the circulating fuel. 5) To provide a free ‘surface for release of the fission product geses which will accumulate in the helium volume at the top of the pres- surizer vessel. 6). To serve as a surge tank for changes in the fuel volume. The pressurizers shown on Figure 5.11 consist of eight connected chambers, adequately sized to accomodate the expamsion and contraction of the fuel fluid | caused by the variation in dengity between the solidification temperature snd | 13OO°F° .Above the free surface in the pressurizer, édditional volume is avail- able for collection of the fission product gases which are scavenged to the | off-gas system. A pipe connection is provided between each fuel pump at the hiéhest elevation of the reactor vessel for connecting to the top of the pressurizer vessel for gas removal. To provide continuous circulation of the low temperature fuel through the pressurizer, the return line is connected from the bottom of the pressurizer to a cconnection at each fuel pumfi suction. 90, 5.2,6 Effect of Volume Heat Source Baffles - and flow dividers are located within the reactor vessel to di;ect the flow of the fused salt fuel through the heat excha.nger*s,;n.mzps_9 ddwncomer, and back to the central core. These baffles are positioned sc that they are in direct contact with a volume heat source on’both sides. Initially, this created some concern because of the possibility of devel§p= ing excessive temperatures and thermal stresses in metal that is in contact with a volume heat source. References 42, 43 and 44 analyze this phenomenon f;n congiderable detail, and indicate that the magnitude of the effect is a ‘ffinqtion of the relative values of neutron flux and coolanfi velocity and their distribution over fhe flow area. An authoritative source was consulted on this problem, as it appiies to this design study (Ref. hS)o- It was con- clfided that the neutron flux and velocity profiles in this reactor are such that the temperature rise in the baffles due to the volume heat source effect would be negligible. 9l. CHAPTER 6. SODIUM AND STEAM POWER SYSTEMS * ) 6.0.0 INTRODUCTION The ultimate use of the reactor hest is the generation of useable power for distribution to the consumer. In order to accomplish this safely, economi-. cally, and vith & maximum flexibility of control, it was necessary to investigate this portion of the plant in detail. This chapter includes the primary loop; intermediate loop, the steam generators,; the sodium piping, the steam cycle and auxiliary power requirements. Other items of interest included are the intermediate heat exchangers, the steam cycle, a plant heat balance and.tfie plant efficiency. | 6.1.0 ~PRIMARY SODIUM COOLANT LOOP The heat of fission in the circulating fused fluoride salt fuel can be transferred to the steam system for the turbine unit by seversl alternate schemes. The acceptance of a scheme depends on the limitation that under all conditions of operation; the heat exchange media minimum temperature will always maintain the minimum fuel temperature in eficess of.the melting point to prevent local solidification. The minimum temperature qf the reactor has been set at 1050°F, for reasons discussed in Chapter 5, whereas the boliler feed water temperature is set at 450°F by the steam turbine cycle design; resulting in a temperature difference of 600°F. This large temperaiure dif- %: - ference must be held nearly constant over the entire load range in order to maintain the system design temperatures which, in 50 doing, creates difficult problems in controlling the heat transfer rates. 92. 6.1.1 Schemes for Removing Heat from Reactor The possible schemes for transferring the heat from the fuel are dis- cussed below. A. Direct Heat Transfer by Boiling Water - It is possible to circulate the fuel external to the reactor directly through a once-through superheater and steam generator without interposing additional loops in the system. The fuel would undoubtedly flow on the out- side of the tubes and the steam inside in the heat exchanges. The following are several apparent disadvantages which make direct heat transfer unacceptable. l. Large Fuel Holdup - The over=-all heat transfer coefficient‘would be lower than if transferred to another fluid. This results because the tubes in the heat exchanger have to be thick, with probably double walls for added safety, to withstand the high pressure steam and also because the film heat transfer coefficient for the superheated steam is low. A low over-all heat transfer requires more heat tranafer area and therefore a bigger heat exchanger, which results in larger fuel holdup requirementis. 2. Reduced Safety - The larger the heat exchanger and the higher the préssure, the greater probability of a leak. Any leak would be from the‘high pressure wvater system into the low pressure, but high temperature fused salt. - Such a leak would cause a violent reaction, increasing the hazards of plant operation and containment. 3. Radiocactive Steam - The fuel will activate the steam, which increases the problems in maintenance, shielding, and designing the turbine and feed - systems completely leaktight. 4. Decomposition of the Steam - Due to the high radiation field, de- composition of the steam results, which requires an extensive O,-Hp recombination system. 93. 5. High Thermal Stresses - Because of the large over-all temperature drops, excessive tube wall stresses will result. 6. Heat Exchanger Design Problems - Because of the large temperature difference between the cold legs of the heat exchanger, the effect of film boiling must be considered in the design with the resulting problems of instability. B. Heat Transfer to Liquid Metal In order to avoid the disadvantages of direct heat transfer, the use of a liquid metal for a primary coolant was considered. The liquid metal coolant will flow inside the tubes and the fuel on the outside of the tubes, and the heat exchanger will be internal to the reactor vessel. Because of the high thermal conductivity of the liquid metal and low vapor pressure, the heat exchangers will be much smaller, therefore decreasing holdup. Safety features will be improved; decomposition and thermal stress problems will be reduced. Furthermore, existing deeign information can be used with confidence in deter- - mining the required heat transfer surface and the possibility of operational instability no longer exists. 6.1.2 Choice of Coolant In making the selection for the primary coolant, the properties of several metals such as mercury, sodium, bismuth and NaK were considered. The general requirements for a coolant, exclusive of nuclear requirements, are that the density, thermal conductivity and heat capacity should be high, whereas the viscosity, melting point and vapor pressure shofild be as low as possible. 1In addition, the coolant should have a high degree of wetting ability, negligible toxicity and, most important, should not expand upon freezing. 9l Combining these properties for an over-all comparison, sodium proved to be the most desirable coolant for the primary loop for the following, several reasons: . : , . , o 1) Lowest pumping power required for a given rate of heat removal. 2) Superior heat transfer properties, resulting in the smallest equipment sizes. 3) Contraction upon solidification. L) Superior neutron economy. Since the fuel primary héat exchanger is internal to the reactor; it is ilmportant to keep neutron absorption losses to a minimum. The coolant also acts as a reflector to reduce neutron leakage from the reactor core. 5) Complete stability under high irradiation. field. A discussion. of these requirements is extended in Section 6.2.2 of this report, in the presentation of the arguments for and against an immediate heat exchanger. 6.1.3 Sodium Coolant Activation As a result of the pfimary sodium coolant flowing through the primary heat exchangefs in the reactor vessel; the sodium will become very radio- active. When the sodium coolant captures a neutron, radioactive sodium-2k is formed and will build up in the coolant system to an equilibrium value which depends upon the particular design and flux level of the reactor. The radioisotope sodium-24 has a half-life of approximately 15 hours, and emits beta particles and two hard gamma ray photons of an energy of 1.38 and 2.75 mev (Ref. 8, p. 525). -The sodium activation imposes many limitations on the freedom of the design of the plant since it affects shielding, plant ‘ - layout, maintenanée, servicing, and personnel agcessibility to the radio- active areas. 95. Impurities in the sodium cause long-lived radioactive contaminants and their effect must be considered in the over-all design, in particular when draining the system in preparation for access to the shielded compartments. Care must be taken to assure all such containmepts are completely flushed. Some of the plant layout and design problems considered becausé of this acki- vation are discussed in Sections 6.7.7 and T.3.0. An approximate value for the level ot activity in £he primary coolant sodium can be determined by applying the following expression when the exposure time is much longer than the mean sodium-24 half-life (Ref. 15, p,_398). Naah activity in primary coolant = ¢ SlaV curies qb = Neutron thermal fiux per cm® * sec S a - Sodium thermal macroscopic cross section cm™! V - Total volume of sodium in primary exchanger Considering only the activities from the thermal flux and uslng an average flux in the sodium in the primary heat exchanger of 0.8 x’lO12 per cm2~sec, obtained from the ORACLE calculations, gives an activation of 1.46 x 106 curies of the radioactive sodium. When this activation is diluted by the approximately 2000 £t3 of sodium in the primary system, the resuitant activity is 0.0258 curies per cc. | Since this is considerable activity, it is necessary that all heat exchangers, pumps and piping containing the primary sodiufi-coolant be ade~ quately shielded to protecthoperating personnel. 6.2.0 INTERMEDIATE SODIUM COOLANT LOOP It is possible to transfer the heat directly from the primary loop to’ generate steam or to interpose an intermediafe loop between the primary and (—— 96. steam portion of the system. Several excellent discussions of the advantages and disadvantages of using an intermediate circuit are given in the literature (Ref. 7, 15, 25). The primary basis for any such selection must be in terms of the economics, safety, reliability and operation of the systen. 6.2.1 Justification of Intermediate Loop In order to mske the decision to include such a loop in the system; the following advantages and disadvantages of using an intermediate loop were considered: Advantages 1. Decreased thermal stress problems in the steam generators. 2. Increased flexibility of the system for partial load operation. 3. Sepafation of the primary heat transfer loop from the steam loop in case of a leak. L, Simpler maintenance of the steam loop outside of the shielding. 5. Avoldance of the possibility of film boiling by decreased At in the boiler. . 6. Reduced shielding problems. T. Reduced possibility of steam decomposition as a result of gamma radiation from the primary sodium loop. As is noted in Section 6.4.1, the presence of oxygen and hydrogen in the steam loop would be very detrimental from a corrosion standpoint. 8. Lighter reactor construction, since the reactor will not be called upon to withstand the full steam pressure in case of a leak in the steam generator portion of the system. Disadvantages 1. Increased piping, pumps and valves for the intermediate loop. 2. More hesat egchangers, | 3. Increased system éize° 4, Larger holdup of heat transfer fluid. 5. More equipment which may be potential socurces of leaks. 97, From the previousiy mentioned discussions (Ref. 7, 15, 25) and the list- ing of the advantages and disadvantages, it is apparent that the main advantages lie in the areas of safety, reliability and ofierafiility, while the principal disadvantages lie in the increased cost of having such a loop. Safety and reliability must of necessity govern the design of any reactor system or its components;, and cost, vhile important, must be sécondary in the design philo- sophy. On this basis, the choice was made to includé an intermediate loop in the system. 6.2.2 Choice of Coolant Fluid The choice of heat transfer fluid to use in the intermediate loop was readily narrowed down to one of the liquid metals or a fused salt. Organics or aqueous media were eliminated becauée of the gamma radiation and high tem- peratures in this loop. Gamma radiation and high temperatures cause these compounds to deccmpose or dissociate. Some of the variables that were con- sidered in choosing the fluid were: 1. High heat transfer properties at reasonable velocities 2. Cost 3. Toxicity k., Relative pumping power. For example, mercury requires from 5 to 10 times the pumping power of the alksli metals. 5. Corrosiveness of the fluid in contact with common materials at the temperatures of interest.: 6. Tendency to dissociate, combine with oxygen or, in genéral, be chemically degraded at operating temperatures and in the gamma flux from the primary loop. T- In case of a leak, the fluid must be compatible with the sodium in the primary circuit and also with the fused salt fuel. 8. The materials of construction necessary to contain the fluid must be compatible with those of the primary loop to prevent the possi- bility of corrosion or mass transfer. . 98, 9, The melting point, heat capacity and thermal conductivity must all exhibit reasonable values for a practical system. 10. Handling characteristics must be practical. For example, the expansion of bismuth on solidification creates many extremely difficult problems. 11. The boiling point and vapor pressure must be reasonable for this high temperature application. Mercury, which has good heat transfer characteristics, has a boiling point of 675°F at atmospheric pres- sure. In order to take advantage of the high heat transfer characteristics of this material, the system would have to be pres- surized to maintain it in the liquid state. Of the wany materials possible, the alkali metals appeared to be best suited for this system. As is noted in the Reactor Handbook (Ref. 7, p. TT3- 777), the heat transfer, where the liquid metal is the sole fluid, does not vary appreciably regardless of the fluid. Sodium was selected as the liquid metal which fulfills more of the requirements listed above than any other fluid. However, this material presents one major problem, namely, the exothermic sodium - water reac;tion° 6.2.3 Sodium-Water Isolation Problem In,order to preclude the possibility of sodium and water contact, systems have been designed (Ref. 15) using such design details as double wall tubes or a third fluid to separate the sodium and water. At present, " one system (Ref. 5) contemplates the use of single wall tubes and éingle tube sheets to separate these fluids. A relatively small mock-up of this system will be tested in the near future and this test shduld provide valuable information regarding the ability of such a unit to run for a period of time, comparable to central station steam pracfice, without leaks. As is noted in Section 6.4.2, it is possible to adequately treat the feed water and toAproperly design a heap exchanger to preclude almost entirely the péssibilitj of stress corrosion failure of the tubes or tube sheets. On the basis that the previous statement represents the present state of boiler _22, and heat exchangér practice, then the tube-to-tube sheet Joint must be care- fully considered. By exercising extreme care and quality control, units.havé been built for nuclefir applications which, prior to service,'have been made- - leakfree. fioweverj in service, the tube-to-tube sheet jJoint is subjected =~ - to a complex stress patterh'and thermal cycling. Under these conditions, there is always the distinct possibility of having some tube-to-tube sheet welds fail. The other type of posoible fallure lies in the realm of faulty tube material being inadvertently used. To prevent this, it 15 necessary to. make some, or all, tests such as radiography, dye penetrant methods, physical tests, ultrasonic inspection, helium leak tests and mass spectrometer, both ' prior to and after the tubing is installed in the unit. 6.2.4 ° Proposed Design for Sodium-Water Isolation of the potent1a1 types of failure mentioned above, the tube-to-tube sheet weld failure presents an area with the greatest possibilit&_of.occur- rence in service. As has been mentioned previously in this section, safety must be a prime consideration in this system. For this reason, the design detail shown in Figure 6.1 is proposed to minimize the effect of a leak in a tube-to-tube sheet joint. In essence, it consists of using single wallj tubes with double tube sheets. A thin walled chamber separates the tube - sheets and helium would be used to purge this chamber to preclude the possi- bility of'é sodium-to-air reaction, and also to act ae a monitor to detect a leak. Should a leak occur at either tube sheet, the heat transfer fluidg would not mix and the magnitude of fihe leak would be monitored fiy the helium. If leaks occurred simultaneously at both tube sheets, only small volumes of materials would be involved and corrective action, such as isolation of the ~ component, could be initiated. Lt el e TS e Tt ST Preliminary work has been initiated* to determine the feasibility of this type of construction. This work consists of a test section of approx- imately 400 1/2-inch OD tubes on 3/4-inch pitch braze welded to a 1/2-inch thick tube sheet on both sides of the tube sheet. Metallurgical and physical tests will be utilized to check the integrity of the braze welds. A summary of this test work up to the date of publication of this report is given in Appendix J. 6.3.0 INTERMEDIATE HEAT EXCHANGER The reasons for placing the intermediate sodium loop and heat exchanger in the cycle are given in Section 6.2.1. In order to maintain reasonable equipment size, six intermediate heat exchangers were used and therefore each heater transfers 100 Mw of heat from the primary to the secondary sodium loop. The heaters are of thé U-tube type to minimize thermal stresses. 6.3.1 Isolation of Radiocactive from Non-Radioactive Sodium Cince the intermediate sodium loop extends beyond the secondary radiation shield, the leakage of rédioactive sodium from the primary loop intc the non-~ radiocactive sodium of the intermediate loop through the heat exchanger is intolerable. It was decided that the intermediate heat exchanger would be designed as a two-fluid type rather than the more costly and difficult to manufacture three-fluld type. The non-radicactive intermediate sodium will be pressurized above the radioactive sodium in the primary loop. Any leakage would then be into the radioactive sodium rather than out of it. The isolation design, as discussed and proposed in Sections 6.2.3 and 6.2.4, would be appli- cable and desirable for the intermediate heat exchanger. * Work Order iésued by W. D. Manly to P. Patriarca, dated July 16, 1956 ORNL=LR=Dwg,=18178 101. STEAM @ 1000°F ?QPV | 7z 7\l o8 IMPOSSIBLE TO BACK BRAZE. PLATE TOO HEAVY. z - HELIUM [ ; : : HOLES TO BE 0.006" LARGER THAN TUBES iFO. LEAVE 0.003" CLEARANCE. BRAZE r'_"—' '—"'-——-’— i Rl et __._.‘| " ' 7777 | | SODIUM | J| — i = j__ i S b /M SUGGESTEDTEST ; FABRICATE TEST SPECIMEN FROM 172" PLATE AND 172" 0.D. TUBE SECTIONS MAKING LARGEST ASSEMBLY WHICH WILL FIT IN FURNACE AVAILABLE IN BRAZING LABORATORY. ( APPROXIMATELY 400 TUBES.) FIG.6.l -SUGGESTED TEST OF DOUBLE TUBE SHEET DESIGN FOR ISOLATION OF WATER AND SODIUM 102. 6.3.2 Calculations for Intermediate Heat Exchanger The calculations for the intermediate heat exchanger design are given in Appendix £ and the specifications are listed in Table 6.2, . » The equations and methods used in determining film coefficients, over-all heat transfer coefficient, heat exchanger area and pressure drop are given in Sections 5.1.4, 6.5.1 and 6.5.2. 6.4.0 STEAM GENERATORS The design of a steam generator for a nuclear system presents several unusual aspects when compared to conventional central stationm practice. The choice between a once-through unit or natural circulation boiler with sepa- rate superheater is not clearly defined for this application. Therefore; in order to determine the proper selection for this system; both types of steam generators were investigated. 6.4.1 Once-Through Steam Generator In principle, the once-through steam generator represents the simplest and most direct method of transferring the rcactor heat to the steam loop. It consists essentially of a number of tubes in parallel with feed water entering the tubes at one end and superheated steam leaving at the other. The heat transfer fluid, which in this case is sodium from the intermediate loop, is on the shell side of the unit flowing parallel to the tube axis. The feed water flow to the unit may be adjusted so that either wet, dry or superheated steam is obtained at the exit, dependent on the design condi- tions. In the event that wet steam is produced, it can be dried by use of - » a relatively small in-line steam separator, eliminating the need for a steam drum. The principal advantages and disadvantages for this type of unit as compared to the natural circulation boiler with a separate super- heater are as follows: 103. Advantages 1. qggactness - For a given set of design parameters, the unit is 20 more compact.. Cdst.- Since the requirement for a separate steam drum is eliminated and the amount of interconnecting piping, tube sheets and shells is reduced, the unit is inherently less costly. ' Control - By the relatively simple method of controlling feed water flow to the tubes, wide ranges of load can be achieved. Arrangement - Since-this type of steam generator does pot require a steam drum and interconnecting piping to a separate superheater, it is adaptable and ideal for a nuclear plant. layout. In particular, the problems of shielding or isolation for reasons of safety are minimized, resulting in simplified arrangements and substantial savings. Internal Arrangement - The use of low pressure sodium on the shell side and high pressure water cn the tube side results in the best over-all heat transfer and a low pressure shell design which is ideally suited to this reactor system. Disadvantages 1. Lack of Water Storage - In the event of complete loss of feed water.. . flow caused by electrical or mechenical failure of the feed pump, the once-through unit has essentislly no water storage and provision must be made to insure the immediate aveilability of an independent water supply to the unit to provide mesns of decay heat removal from the reactor. Water Treatment - Experience with high pressure once-through boilers in Eurcpe has demonstrated that the major difficulty encounteres.in . these units has been the inability o maintain extremely high purity - wvater required for successful operation. According to Shannon (Ref. 1), investigations have indicated that the oncemthrough unit can be operated successfully only if the following conditicns are mets:: a. Solids content of feed less than 0.5 ppm - b. Oxygen content of feed not to exceed approximately O 006 ppm c. Removal of oxygen formed by the dissociation of steam d. The absolute exclusion of chlorides from the system e. pH control by use of aqueous ammonia It is believed that these conditions can be met when operating the plant (Ref. 1). 3. Thermal Stress - The outlet of the evaporating section is particularly subjected to conditions that lead to stress reversals and the possi- bility of fatigue or stress corrosion failure resulting from the intermittent heating and cooling of the tube walls by the finely dispersed particles of water in the steam at this location. OGS 104, L. Flow Instability - In forced circulation boiling, the problem of flow instability (Ref. 2, 3, 4) in parallel circuits can endanger the per- formance of the unit. If an obstruction or scale builds up in a tube, the flow rate in the tube may actually lead to an increasing flow resistance due to the increase in the average specific volume of the + tube contents. Ultimately, the tube may become steam bound and become completely ineffective from the standpoint of constituting active heat transfer surface. A further effect under these conditions is to de- crease the total feed water flow to the unit, which also decreases . its ability to perform properly. 5. Design - The calculation of the heat transfer surface required to perform a specified duty with a complete change 6f phase is especially difficult in those sections where there are high percentages of steam by volume. At present, such calculations represent "educated guesses”, and the reliability of such calculation procedures must be validated against future tests such as are presently being contemplated by APDA (Ref. 5). 6.4.2 Natural Circulation Boiler-Separate Superheater Steam Generator The natural circulation boiler with a separate superheater requires more basic components than the oncé-through unit and; therefore; represents a more complex arrangement for the removal of the reactor heat. Inthe natural circu- lation boiler, steam is generated on the shell side and low pressure sodium is on the tube side. The steam-water flow in the boiler is directed by baffles which also serve as tube supports. The steam-water mixture flows through the risers to an external steam drum by natural circulation. As in conventional practice, feed water is introduced into the system at the steam drum to avoid ‘the possibility of temperature shock to the steam generator. The feed water is heated to saturation by condefising steam in the steam drufi and saturated water is recirculated to the boiler by the downcomer system. Saturated steam, mechanically separated in the steam drum, is Piped to the separate superheater. The principal advantages and disadvantages 6f the natural circulation boiler with a separate superheater, as compared to the once-through boiler, are: 105. Advantages lD Reliability - This type of steam generation system has been used in principle in several actual and proposed reactor installations. Among these are the Mark I and Mark II STR units, the PWR, SIR,; SAR, HRT, etc. The reliability has been established under actual and test conditions of reactor operation. Design - Design methods are available which enable the designer to predict the performance of these unite to the required degree of accuracy for engineering of the system. Water Storage - This system has an inherently large water storage built intu 1t in the steam drums. Thus, in the event of a forced outage or the failure of a feed pump, the system is still capable of removing heat from the reactor for several hours. In addition, the effect of a perturbation of the reactor part of the system is minimized on the steam portion of the system because of the damping effect of the large amount of water storage. Water Treatment - In both this unit and the once-through boiler, careful control must be exercised in treating the feed water for proper operation of the units; however, for this unit, the water treatment problem is less stringent than for the once-through - boiler since scaling and tube plugging are not as critical. The problem of stress corrosion prevails in both types of unit and needs to be carefully considered. Inherent Stability - This type of design has established its in- herent stability under various operational conditions. The type of instability that occurs in forced circulation boiling does not exist in the natural circulation system. Disadvantages ld Cost - Because of the necesslty for a steam drum, piping from the steam drum to the superheater, interconnecting piping in the steam generator, multiple tube sheets, shells and heads, this type of unit is more costly than the once-through unit. Arrangement - The engineering design requirements of the natural . circulation equipment make it less adaptable to either shielding or compartment arrangement in a nuclear system. Size - For a given set of design parameters, the natural circulation system will occupy more volume than the once-through system. The natural circulation design is predicated on the basis that the high pressure water is on the shell side, which subjects the tubes to collapsing pressure. This results in thicker tubing with an attendant increase in weight and decrease in heat transfer capa- bility per unit volume. 106, 5., Control - As described in Section 7.4.0, the control of this unit presents problems in operating the system over wide ranges of load. In order to accomplish the desired regulation; it is necessary to provide elther external regulation of -the reactor or special opera- tional control of the heat transfer circuits. - In a previous section of this report (6.2), the problem of a possible godium-water reaction has been discussed and a tentative scheme has been pro- posed to minimi?e the pogsibility of this reaction by the use of double tube sheets. In all of the systems built to date (Ref. 15), sodium and water have been separated by at leagt two metal walls, with the one exception of the APDA tests (Ref. 5). The boilers and superheateré for this present sfiudy'have been designed on the basis of fising a single metal wall to separate the sodium from the wa.t'.er.T In order to indicate the magnitude of the size change when two metal walls are used, the superheater of the natural circulation boiler with a separate superheater was calculated on two bases. The comparison design was for two metal walls having perfect thermal contact (no interface resistance), but not mechanically joined. Each wall thickness was taken as being separately capable of withstanding the full design préssure° After careful consideration of all the factors involved, it was not possible | to select either the once-through or natural.circulation boiler with separate superheater as having an over-all advantage over the other, for the proposed application. Thefefore, for the purpose of this study, it was considered that prelimingry designs for both units should be developed. The calculation proce- dures used in making fhese designs are given below. These designs were made on the basis that the units would conform to present technology wherever possible, and no attempt was made to ofitimize the performance of this equip- ment, as this optimization can be realistically done only after the system has been finalized. 107. 6.5.0 STEAM GENERATOR CALCULATION PROCEDURES 6.5.1 Heat Transfer A. Boiling In making such design studies, it was apparent that the greatest'diffi- culty was the inadequacy of the information available in the field of boiling heat transfer. Since these units must operate with large temperature differ- ences between the fluids, there is the additional problem of the possibility of unstable film boiling occurring. The literature (Ref. 6, T7) indicates that a temperature difference of 45CF between the surface causing boiling and the saturation temperature of the fluid being boiled is the maximum value pogsible to have nucleate boiling. To a first approximation,'this temperature differ- ence is independent of pressure. When the above-mentioned temperature differ- ence exceeds 459F, unstable film boiling results. At the present state of development, it is Qirtually impossible to design engineering eqfiipment with any assurance of predicting performance when in the unstable film boiling region. In Appendix F, é check calculation was made to assure that these units were operating in the nucleate boiling range. A solution to the problem of designing & heat exchanger with variable fiucleate boiling coefficients 1s available (Ref. 8, p. 699). It is exceed- ingly difficult to utilize this for design, s}nce it 1s necessary to use iterative techniques to obtain a single soiution° Therefore, an anaiysis was made in Appendix F to determine a reasonable value of an average boiling plus scale coefficient to use for fhe design of these units. It is to be noted that the operating temperature différences in the units analyzed were less than the temperaturé differences in this study. The numerical value of the boiling plus scale coefficient was 2000 BTU /hr/sq ft/F,Aand since the 108, temperature differences in the equipment will be greater than for the units analyzed, this value is conservative for design. This value of 2000 was used in all cases where there was nucleate boilifig, B. Tube Resistance The resistance of a tube wall to heat transfer based on the outside tube surface is, ro In ro : | (6°l) k Ty ' where the thermal conductivity of the tube material is taken to be the value at the average tube wall temperature, as given in Figure A.l. C. Superheated Steam Resistance Colburn (Ref. 9) has correlated the heat tranéfer data on many fluids and his work (or winor modifications) has been the basis for most of the corre- lations for fluids in turbulent flow inside tubes. The greatest difficulty | in fising any heat transfer equation lies in the corfect values of the physical properties of the.fluid to use. In order to correctly apply these équations, it is necessary to use the same properties as the correlator used. At times, this is difficult to ascertain, and in these instances,; the designer must exercise Jjudgment in the choice of physical properties. lThere has been, and still is (Ref. 10), a great deal of controversy regarding the properties of superheated steam. It is the general opinion (Ref. 11, 12) that the properties of steam given in Appendix A represent the "best" values available at this time. .These values were used in a recent investigation (Ref. 13) of the flow of super- heated steam in annuli. The recommended equation for the heat transfer coefficient when superheated steam is flowing-turbuI§ntly in pipes is based 109. on the properties in Appendix A, and ia,' hd - .023 (gfi)lk} aveP-8 | (6.2) k k A where all properties are evaluated at the avefage bulk temperature. Since this is‘based on the inside tube surface, it is necessary to multiply the | h from the above equation by the ratio of the inside tube radius to the outside tube radius to base h on the outside tube surface. The resistance is the inverse of h . Equation 6.2 has been plotted in convenient form for both subsaturated water and superheated steam in Appendix B (FigurelB.l)° This form was suggested by Dr. G. E. Tate (Ref. 14). D. Sodium Resistance a. Sodium flowing inside tubes - In the natural circulation boiler,_. sodium is on the tube side. In this case, the most accepted value of the sodium film coefficient is given by the Lyon-Martinelll equation, which is listed as Equation 5.1 in Section 5.1.4. It is noted that the sodlum coeffi- cients are generally so high as to constitute only a smail part of the overQall fesistance of the equipment. b. Sodium flowing on the shell side - Several correlations have been proposed for calculating the film coefficient when godium is floying turbu- lently on the shell side of baffled and unbaffled heat exchangers (Ref.‘ 15, 25). Kern (Ref. 20, p..136) notes the difficulty of calculéting the heat transfer coefficient on the éhéll gide of a baffled heat exchanger, even when common industrial heat exchange fluids are used. Kern gives a éemi-empir}cql equation of the Colburn type (Ref. 9) and states that "....excellent agreement is obtained if the hydraulic radius is calculated along instead of across the long axis of the tube....". For the range of Reynolds' numbers involved, Dwyer 0. (Ref. 25, p. T78) recommends the use of the Lyon-Martinelli equa.tion° This was used and all physical properties were evaluated at thé average fluid £ulk temperature. The reciprocal of the sum of all the above applicable resistances is the over-all coefficient of heat transfer for the equipment. In general; the over-all coefficient of heat transfer will vary throughout the length of the unit since it is temperature dependent. In such cases, Equations 8-17 of Reference (6) should be used to determine the surface required for a given set of design paraméters. For this study, a constant average value of over-all coefficient of heat transfer was used, since the variation of fluid‘properties is not great and the error so incurred is well within the accuracy of the heat transfer correlations used. The surface was then calcu- lated from Equation 6.k4. q =Uy x8 x Aty (6.4) 6.5.2 Pressure Drop According to Reference (15), "The use of liquid sodium and NaK as heat transfer media involves no new problems in the application of fluid mechanics”, Therefore, the calculations of the various pressure drops in the equipment resolved themselves into identifying the various resistances involved and applying known. -solutions. Pressure losses due to nozzles and interconnecting piping are not included as part of the equipment pressure drop. These losses are included in the piping pressure drops; however, the calculation procedures of this section were used to evaluate the piping pressure losses. A. Entrance (Contraction) Losses (Ref. 16) The entrance loss was taken to be independent of Reynolds number and also independent of the ratio of the'tube flow area to the header flow area. 111.° . This loss is given below (Eq. 6.5), where the:velocity is the average‘tube velocity as determined from the one-dimensional, steady state continuity - relation (Eq. 6.6), AP - 0.5 !3 (psi) | (6.5) 2g 1 ' VoW | | | (6.6) B. Exit Loss (Ref. 16) It was assumed that there would be no regain in pressure at a tube exit and the decrease in static pressure at the exit would be one veloéity head, AP - 1.0V2 @ (psi) (6.7) 2g 1LL ' _ _ C. Friction (Single Phase Incompressible) (Ref. 16) The friction factor uscd in the Fauning equation is a function of the relative roughness of the tubing and the Reynolds number. The relative roughness of the tubing was taken to correspond to fhe smooth pipe curve of the Moody diagram. The friction factor was taken directly from the Moody diagram at the proper Reynolds number, oL v @ (pst) | . (6.8) AP = f 3 5z I pe : , D. Friction and Acceleration (Two Phase) (Ref. 7, p. 69) The method of Martinelli and Nelson was used to calculate the friction 7 and acceleration pressure losses in the boiling section of the once-through boiler. 112, E. Friction and Acceleration (Single Phase Compressible) (Ref. 7, p. 69) The acceleration and pressure drop for single phase compressible flow such as occurs in the superheater can be calculated from Equation 23 of Refer- ence (7), page (69). The preliminary calculations of these pressure losses showed that they were of the order of 3 percent or less of the system pressure. Therefore, the arithmetical average density was used and the flow was treated as before (Section 6.5.2, Subsection C), neglecting the acceleration pressure drop (Ref. 21, p. 234). F. Bend Losses (Ref. 16) The bend loss pressure drop is known to be & function of at least three variables. These are relative roughness, Reynolds number, and the ratio of the radius of bend to tube diameter. The value of this loss for 180° vend (Ref. 16) was taken to be 1/2 velocity head, p.o.5s¥ £ (pst) | (6.9) Shell side preséure drops were caléulated ou the same baoio as the tuhe side pressure drops above. The equivalent diameter concept (Ref. 6) was used based on typical unit cells of the tube layout pattern. No attempt was made to estimate the effect- of tube baffles on the shell side pressure drops. The drum size, its location above the stesm generator, and the unit's abllity to maintain an‘adequate‘circulation rétio vere investigated ofily from -8 feasibility standpoint‘(Ref° 22). The results only of this pfeliminary calculgtipn are tabulafied on ihe equipment data ;heet for thexnatural circu- 1ati§n uni#. | | | In ;ll instances, a minimum distance of 1/4 inch was allowed between adjacent tubes at the tube sheets, to permit the welding of the tubes into 113. the tube sheets. Calculafiions of the péssible relative expansion of the tubes and shells of these units (Ref. 24) indicated that a straight tube exchanger with fixed tube sheets would not be suited for this application. Therefore, the "horse shoe" shaped unit was used for all of these exchangers. 6.6.0 SUMMARY OF HEAT EXCHANGER DATA The specifications for the primary heat exchanger, intermediate heat exchanger, once-through steam generatér, convection boiler, single tube wall superheater, and double wall superheater are included in Tables 6.1 through 6.6. 6.7.0 SODIUM PIPING SYSTEMS The technology of sodium piping is well established (Ref. 15), and it is considered that the piping systems proposed in this study can be designed satis- factorily. Rough piping designs and layouts have been prepared and the follofiing sections briefly describe the considerations which were used to Justify these preliminary designs. Qertain design criteria, which are not shown on the lay- out drawings or in the tabulations, are also discussed. 6.7.1 Pipe Length and Size AY In estimating pipe length, an attempt was méde to provide'a genefous ~ allowance for expansion loops. Figures 7.2 and 7.3 were used'to make,reasonmr able estimates of over-all pipe lengths. Pipe wall thickness was chosen on the basis of'commercially available pipe. Sin;e the sodium system is a low pressure system, Scheduié 40 pipe was generally selected. Pipe diameters were influenced by the-chqicé of sodium velocity. -Since 30 fps to 40 fps sodium velocity is being cénsidered for desigfi purposes (Ref. 15), all pifie slzes were selected so that the velocity_did not exceed 40 fps and, in most casés, the velocity was near the 30 fps value. Wherever excessive pressure drops were encountered in the initial estimate; the pipe size was increased. 11"' L] TABLE 6.1 EQUIPMENT FOR SPECIFICATION SHEET Name of equipment: primary heat exchanger Description: U-tube bundles Service of Unit: fused salt to sodium No. of units required: 2k Connected in: parallel Surface per unit: 2230 £42 PERFORMANCE OF ONE UNIT Shell Side Tube Side Fluid circulated fused salt sodium Total fluid entering, 1b/hr 48 x 10° 5.7 x 106 Temperature in 1200°F 1000°F Temperature out | 1050°F 1150°F Operating pressutre 10 1b/sq in. 125 1b/sq in. Velocity 3/7 ft/sec 12.8 ft/sec Pressure drop (not including nozzles) 8.5 psi 9.1 psi Heat exchanged, BTU/hr 9,09 x 107 = Log mean temperature difference SOOF Transfer rate design - - 815 BTU/hr-£t2-°F ¥ Based on a tube side surface of 2230 ft2 per bundle CONSTRUCTION Tubes: .O42 inch wall type ALSL 304 clad with .023 inch type "L" nickel; No. 650; 0.D.,5/8 inch; I.D.,4#95; Effective Length 21 ft; Pitch 3/4 inch Tube sheets: stationary Baffles and tube supports: as required Code requirements: ASME, TEMA and others Remarks: shaped for expansion | MATERTALS See Chapter 3 115. TABLE 6.2 EQUIPMEN? SéECIFICATION SHEET N Name of equipment: intermediate heat exéhanger Description: horseshoe shépe, P1xed tfibe shéét exchanger Service of unit: sodium-to-sodifim | No. of units required: 6 Connected in: parallel - Surface per unit: 1910 f£t° Shells per unit: 1 PERFORMANCE OF ONE UNIT : Shell Side Tube Side Fluid circulated. sodium sodium Total fluid entering, 1b/hr 7/62 x 106 2.67 x 105 Temperature in 1150°F 675°F Temperature out | 1000CF ~ 1100°F - Operating préssure | 50 1b/sq in. 150 l‘b/sqvin° Velocity - | . 12.29 ft/sec. 13.73 ft/sec _ Pressure drop (not including nozzles) - 1.61 psi 9.02 psi Heat exchanged, BTU/hr - | 3.41 x 10° B Log mean temperature difference ' 1&7QF ) Transfer rate design o 1216 BTU/hr-ftz-QF . o CONSTRUCTION Tubes: No.:1100 0.D.:1/2 in. I.D.:O.h}6 in. Length: 13.3 £t Pitch: 3/4 in. .~ Shell (approx. dia. & length): horseshoe shape, 30 in. dia.,, 16 ff developed length Tube sheets: stationary Baffles and tube supports: as required Code fequirements: ASME , TEMA and others Remarks: shaped for expansion MATERIALS - See Chapter 3. 116, TABLE 6.3 EQUIPMENT SPECIFICATION SHEET Name of equipment: once-through steam generator Description: horseshoe shaped, fixed tube sheet exchanger o Service of unit: sodium-to-water No. of units required: 2 Surface per unit: 10,600 £t° Shells per unit: 1 Connected in: parallel PERFORMANCE OF ONE UNIT : Shell Side Tube Side Fluid circulated ' sodium water ) Total fluid entering, 1b/hr 7.93 x 10° 9.78 x 107 - Temperature in 1100°F 450°F Temperature out 675°F 1000°F Operating pressure | 100 1b/sq in. 1950 1b/sq in. Velocity 13.25 ft/sec 5.14-67.3 ft/sec Pressure drop.(not including nozzles) | 48 psi 50 psi Heat exchanged, BTU/hr | | 1,02 x 107 - (preheat 178.5°F Log mean temperature difference (evaporating 221 .0°F - ' (superheating 192, 5°F - (preheat 690 BTU/hr-ft2 -OF Transfer rate design (evaporating 700 BTU/hr-ft=-OF (superheating 300 BTU/hr-ft€-CF CONSTRUCTION Tubes: No. 1,620 0.D. 1/2 in. I.D. 0.37 in. Length 50.ft Pitch 3/4 in. Shall (Approx. dia. & length): 34 in. I.D. 26 £t - 6 in. Tube sheete: stationary Baffles and tube supports: as required Code requirements: ASME, TEMA and others Remarks: shaped for expansion MATERIALS See Chapter 3. 117. TABLE 6.4 EQUIPMENT SPECIFICATION SHEET Name of equipment: convection boiler Déscription: horseshoe shaped, fixed tube sheet exchanger Service of unit: sodium-to-water No. of units required: 2 2 Surface per unit: 10,600 ft Shells per unit: 2 Connection in: parallel PERFORMANCE OF ONE UNIT _ Shell Side Tube Side Fluid circulated N water sodium Total fluld entering, 1b/hr 9.75 x 10° - 8 x 10° Fluid vaporized _ water — ' Temperature in L450°F 960°F 'Temperature out 635.8°F 675°F Operating pressure 2000 1b/sq in. 100 lb/sé in. Velocity ‘ | None | ' 21,9 ft/séé Pressure drop (not including nozzles) cmmea 38.6 psi Heat exchanged, BTU/hr 6.86 x 108 | Log mean temperature difference 1359F Transfer rate design ' 480 BTU/hr-fta-QF - CONSTRUCTION Tubes: No. 2000 O.D. .75 in. I.D. .42 in; Length 28.5 ft Pitch 1 in. Shell (approx. dia. & length): L47 in. I.D. | 15.8 £t | Tube sheets: sfiatiopary | | Baffles and tube supports: as required Code requirements: ASME, TEMA and others Remarksf shaped for expansion | MATERIALS See Chapter 3. TABLE 6.5 EQUIPMENT SPECIFICATION SHEET Name of equipment: superheater, single wall Description: horseshoe shaped, fixed tube sheet exchanger | - " Service of unit: sodium-to-steam No. of units required: 2 X Surface per unit: 3520 ft2 Shells per unit: 1 Connection in: parallel - PERFORMANCE OF ONE UNIT Shell Side Tube Side Fluld circulated | sodium steam ) Total fluid entering, lo/hr - ' 8 x 106 9.75 x 107 - Temperature in : . 1100°F 635089F Temperature out 960°F 1000°F Operating pressure 100 1b/sq in. 2000 1b/sq in. Velocity 17 ft/sec T4.8 ft/sec Preesure drop (not including nozzles) 1 psi Lk fisi Heat exchanged, BTU/hr | 3.34 x 108 o ST Log mean temperature difference | : 191°F ) Transfer rate design 500 BTU/hr-£t2-OF CONSTRUCTION Tubes: No. 1275 0.D. .5 I.D. .37 Length 22.25 Pitech .75 Shell (a.pproxo'dia,o & length?s 32 im. IOD; 2.5 £t | ; Tube sheets: staticnary | _ - Baffles and tube supporté: as fequired | Code requirements: ASME, TEMA and others Remarks: shaped for expansion | MATERIALS ) B z‘_ - See Chapter 3. i TABLE 6.6 EQUIPMENT SPECIFICATION éHEET ) Name of equipment: superheater, double wall - Descriptions horseshoe shaped, fixed tube sheet exchanger Service of unit: sodium-to-steam . No. of units required: 2 . Surface per unit: 6400 fta, Shells per unit: 1 Connection in: parallel PERFORMANCE OF ONE UNIT . Shell Side Tube Side — Fluid circulated ' . sodium . steam Total fluid entering, ib/hr | 8 x 10° 9.75 x 105 Temperéture in 1100°%F 635.8°F - Temperature out 960°F 1000°F Operating pressure ' | 100 1b/sq in. 1950 1b/sq in. Velocity | 10.2 ft/sec 63.5 ft/sec Pressure drop (not including nozzles) b psi 4O psi - Heat exchanged BTU/hr o 3.34 x 108 . Log mean temperature difference l9qu | Transfer rate design | 275 BTU/hr-fta-oF‘ CONSTRUCTION Tubes: No. 1500° 0.D. .666 I.D. .500 Length 25.75 £t Pitch .916 ,500 .370 Shell (approx. dia. & length): 42 in. I.D. x 14.3 ft Tube sheets: stationary - Baffles and tube supports: as required Code rgquirements: ASME; TEMA and others Remarks: .shaped for expapsion MATERIALS See Chapter 3. 120. ST The important piping data, length, size, velocity and pressure drop are given in Table H.1 6.7.2 Pipe Slope In laying out the sodium piping system, a continuous slope of 5/8 inches per foot will be used throughout. The purpose of this slope is toc facilitate charging, venting, draining and washout. Where traps or pockets are required, individual draifi lines will be provided at these locations. These drain lines are not shown in the piping layout. All of the sodium piping, with the excep- tion of the primary heat exchanger tube bundles and a portion of the feed lines, can be drained by gravity. 6.7.3 Expansion Tanks An expansion tank will be provided in each circuit of the primary and intermediate sodium loops, since their design is based on & closed loop system. The location of the expansion tanks will be at the highest point in the cir- cuits and in the vicinity of the pump suctions to take advantége of pressure drop in the system, to-minimize the cover gas pressure. Computations were made to determine the size of tanks necessary to accomodate the expansion of a full hystem of sodium from a temperature of 208% to 1200°F. This consideration indicated that tanks approximately 4 ft in diameter and 5 1/2 ft long are required for each circuit. To reduce the size of these tanks, provisions are included for draining sodium back to the drain and charge tank as the system is brought up to operating teuperature by providing an ovérflow line near the top of a reasonably sized expansion tank. An electromagnetic pump will be used in this overflofi Bystém to add sodium to the expansion ténk during operation to maintain a minimum safe fluid level in:the system. 4 121. 6.7.4 Cold Traps Cold traps are located in the cold leg of each closed sodium loop and on the suction side of the pumps when pcssible. Either static or circulating ccld traps could be used. 6.7.5 Valves A minimum number of valves are shown on the General Piping Layout, Figure 7.1, and the design presented is considered practical and operative; however, the final philosophy of the design dictates the exact number and location of valves in the Bodium piping. . A careful evaluation of the operating and maintenance procedures to be followed would determine the optimum number of valves and their location. In general, low pressure drop valves.will be used; and freeze stgm valves will be used in preference to bellows sealed valves because of the large pibé gizes required. Bellows sealed valves may be desirable in some locations for small pipe sizes. Check valves will require no stem seal. 6.7.6 Pressure and Instrument Taps No consideration was given to pressure and instrument taps since this would require going into more detail than time would permit. 6.7.7 Drain and Charge Tanks Drain and charge tanks will be provided for eech circuit of the primary and intermediate sodium loops. These tanks will be located below the lowest. point in the circuit so that draining can be accomplished by using gravity assisted by gas pressure. All components of the sodium system are drainable, except the primary heat exchangers and portions of the feed lines connecting the heat exchanger bundles and the ring headers. It is considered that the sodium level remaining 1122, in the feed lines after the system has been drained can be lowered sufficiently, for maintenance purposes, by applying gas pressure to the upper ring header to force the sodium into the ldfier ring header, from which it can be drained to the tank. Six tanks, approximately 6 feet in diameter and 15 feet long, will be installed in the priméry loop;‘ These tanks will be located in shielded com- partments directly below the intermediate heat exchangers. Valving and other provisions will be included so that all radioéctive sodium can be drained and fluéhed with non-radioactive sodium from the piping, intermediate heat exchangers,b and primary loop pumps in the intermediate compartments during outages for main- tenance work on the above-mentioned compcnents. Three tanks, approximately 8 feet in diameter and 15 feet long, will be installed in the intermediate ioop; These tanks will contain non-r;dioactive sodium and will be located below the steam gefiérators outside the reactor con- tainment vessel. It has been estimated that the total sodium volume in the primery and intermediate loops 1is approximately 2000 ft3 and 1500 ft3, fespéctively° At 208°F, the:melting point of éodium, this volume represents a total sodium mass of 202,650 pounds which will be used in these two coolant loopé° Adai- tional sodium will be kept in the drain and charge tanks to make up fortlosaes during opefation and to reduce the'thermal éhock involved in d:ainihg a hot system by allowing the cold sodium in the tank tO‘fiix with the hot_sodium. (Note: The above estimate of sodium volume in the intermediate loop was made considering the straight-through steam generator only.) L 6.7.8 Cover Gas Cover gas, probably purified heiium,'is to be provided_over the sodium‘": surfaces in the expansion tanks and drain and charge tanks. A gas equalizing 123 Iine will connect these tanks to facilitate gravity drainage. Valving will also be provided in this connecting line to allow gas pressure to be directed to the expansion or drain tanks for forced drainage or charging operatiqns° 6.7.9 Sodium Pumps Centrifugal pumps will be used for pumping the large quantities of sodium required in the coolant systems. Six pumps operating with a head of 200 feét at design flow of 18,550 gpm will be used in the primary loop system. Two pumps operating with a head of 140 feet at a design flow of 18,350 gpm are required in the intermediate loop. As shown in Figures 7.l and 7.3, a spare pump is to be installed in the intermediate loop. These pumps will have rated load motor horsepower requirements df 1335 and 950, respectively. It 1s possible that design changes can be made in the two coolant ld&ps 8o that identical pumps can be used in the primary and intermediate coolanfi lecops. Such interchangeability would fie desirable from maihtenanee considera-~ tionc and stocking of spare parts. 6.7.10 _Heating the Salt and Sodium Loops for Startup Because of the large quantities of fused salt and sodium required for the various loops, there are many profilems involved in designing methods for preheating the systems, charging the systems and maintaining approximately design operating temperatures with thé reactor subcritical during startup procedures. The design of the heatiné accessories ié bfised on the concepts that adequate heéting capacity will be inclfided to: (1) melt the sait.mixe. ture and sodium in the primary loop; (2) heat and maintain them at a minimum temperature of lOBOqF while continuofisly‘circulated in tfieir réspeétive systems;.and (3) melt tfie sodiu& fof the intermediate loop -and heat it to approximately 250°.F° 124, At the present time, induction heating is the only possible method of preheating the primary sodium and salt mixture up to 1050°F, since this tem- perature is above the temperature range of resistance heating. The use of induction heating with an austenitic stainless steel piping system and reactor vessel requires high frequencies and is quite inefficient (Ref. 15, p. 255). Utilizing this method of preheating would require development work before application. Both systems will be charged by using pressurized helium. Extreme care must be exercised to assure that all the oxygen has been purged from the system prior to charging. The following procedures were considered as possible methods for heating the systems prior to initial operation: 1) Initially, the reactor vessel is internally heated to at least 250°F to prevent the sodium from solidifying in the primary exchanger during the charging of the sodium. 2) The sodium for the primary loop is melted, charged, and then slowly brought up to 1050°F by its heating system and friction heat of the pumps in the systen. 3) The sodium is continuocusly recirculated for a long period until it is determined that the reactor vessel internals have reached a tem- perature of at least 1000°F by the sodium circulating through the primary heat exchangers. 4) During this period, the salt mixture in the fuel dump tanks is melted and the salt piping system and reactor vessel are heated to 10509F, in preparation for charging. 5) When the complete salt system has reached a temperature of 1050°F, the salt is charged into the reactor vessel. 6) The fuel circulating pumps are started and the fuel is circulated in preparation for uranium fuel addition to bring the reactor to the critical condition. During this latter operation, the salt and primary sodium systems are maintained at 1050°F by their indivi- dual heating systems. 125. 7) The intermediate sodium system is heated to 250°F and charged when criticality is reached, and means of heat removal from the reactor system is necessary. Charging the sodium in the intermediate loop last, keeps the high temperature systems isclated from the steam . system heat sink for the startup preheating. This reduces the pre- heating requirement for this intermediate loop, since it will then require only enough heating capacity to maintain the metal tempera- tures above the melting point of sodium to prevent solidification . _ during charging. 6.8.0 STEAM POWER SYSTEM A conventional regenerative steam cycle was chosen for this fused salt reactor power plant. In this type of cycle, steam is extracted from the tur- bine in various stages of expansion and is used to heat feed water (Ref. 23). This reduces the amount of heat converted to work in the turbine, but there is an even greater reduction in the heat rejected to the condenser cooling water. This results in an increase in efficiency over that of a cycle with no steam extraction for feed water heating. The over-all heat balance diagram and the conditions of the working fluids are shown in Figure 6.2. The results of the heat balance are given in Table 6.8. 6.8.1 Steam Turbine and Steam Cycle Heat Balance The stefim'conditions entering the turbine were set at 1800 psig and - 1000°F. These are standard and conservative. | The turbine exhaust pressure was set at 1.5 in. Hé, which is easily attainable with a reasonable condensef cooling water temperaturé° it is pro- béble that in most locations the exhaust pressure.could‘be lowered to 1 in. . ‘ Hg or-léss, with a resulting increase in the cycle efficiency. The temperature and quantify of cooling water available are a major factor in determining the - location of a steam plant. If a fiatural supply of cooling Qater is not a#ailm : able, it is necessary to resort to a cooling tower, with a resulting increase in capital cost and operating expense. 126, The turbine-generator losses were found by the method deééribed in Refer- ence 41. In order to apply this method, the type and size of the turbine must be known. It was known that the size of the turbine would be between 200 and 250 Mw, and from this it wae decided that a 3600 rpm single shaft turbine of the tandem compound triple flow type with 26-inch last stage blades would be used for this study. This type of turbine has a high pressure section and three low pressure sections on one shaft. All the steam passes through the single high pressure secfiion and then is divided among the three low pressure sections. This allows smaller wheel diameters than those of a single low pressure section. A detailed econcwic study would have to be made to ascertain whether the higher efficiency of a cross compounded double flow turbine would Justify the higher capital cost. This unit would have a high pressure 3600 rpm section on ore shaft and two low pressure sections at 1800 rpm on another shaft. For a steam cycle in the EOOQESO Mw range, i%t is the practice to have six or seven feed-water heaters; with the water leaving the last heater and enter- ing the boiler at 450°F (Ref. 41). 8Six feed-water heaters and a final feed water temperature of 450°F were assumed for this study. It is quite possible that a final feed water temperature in excess of'h50°F would be desirable. A detailed economic study would show if the increase in c¢ycle efficiency due to a seventh feed-water heater justifies the additional expenditure for the heater, controls, piping and maintenance. The temperature rise of the feed water was taken’to be apprdximately the same in each heater. The fioo 3 heater was made the open heater in this cycle. In an open heater, the feed watef, extracted steam, and drains from higher pressure heaters are mixed and any dissolved gases are driven out of the mixture of steam and satura- ted water. These gases are remcved from the cycle by venting of -the atmosphere. 127. All other feed-water heaters were taken to be of the closed type with no contact between the feed water and extracted steam. Two pumps were placed in the cycle. The hot well or condensate pump takes the condensed steam from the condenser hot well and discharges it into the No. 3 opefl.heater, after passing through Nos. 4, 5 and 6 heaters. The hot well pump works against the difference in pressure;, the pressuré drop due to friction, and the difference in elevatiom‘befiween the=No; 3 heater and con- denser. The boiler feed-water pump takes the feed water from the No. 3 heater and pumps it through Nos. 1 and 2 heaters into the boiler. This pump works against the préssmre difference, friction drop, and elevation difference be- tween the boiier and No. 3 heater. 6.8.2 Calculation Procedures The method used to determine the cutput of the 600 Mw reactor power plant is described below. Detailed caldulationé are given in Appendix Go_ The heat energy converted to work at the turbine wheel was found by cglcuo lating the steam flow and enthalpy change between the throttle, extraction' points and condenser (Ref. 23). The stea; flow is constant between the extrac- tion points and the wérk done between extraction points is'given by, work = steam flow ;g,x enthalpy change BTU o . (6015) hr 1b ' The steam flow at the turbine throttle is found by a‘heat.baiance betweén. the 600 Mw heat input from the reactor and the enthalpy differenée between the steam and feed wafier at the boiler outlet and inlet. throttle steam flow 1b s 600 Mw x 3413 x 103 BTU (6.11) bhr AH BTU MWH 128. In order to calculate the various extracted steam and water flows in the feed water heating portion of the cycle, it was necessary to make the following assumptions: ‘ . 1. The terminal temperature difference in the feed-water heaters was taken as shown in Figure 6.2 (Ref. 23). This is the temperature difference between the feed water leaving the heater and the con- =~ " densed extraction steam leaving the heater and drained into the next lower pressure heater. The open heater (No. 3) has no terminal temperature difference since the feed water, extracted steam, and drains are mixed. 2. The steam pressure in a heater was taken to be the saturation pressure corresponding to the heater drain temperature. 3. No. 3 heater has no drain and the feed water leaving this heater was assumed to be saturated. 4, A pressure drop of 8 percent was assumed between the turbine extrac- tion points and the feed-water heaters. 5. No enthalpy correction was made for sub cooling. 6. It was assumed that there was no heat loss between the reactor and boiler. This is closely approached in practice by proper insulation of the heat exchangers and connecting piping. T. The efficiency of pumps and motors was taken to be TO percent and 95 percent, respectively (Ref. 41). The extracted steam flows required by all heaters, except heater No. 3, were calculated from a heat balance between the heat gained by the feed water and the heat given up by the extracted steam and drains from higher presauréi o heaters. For No. 3 heater, both the extracted eteém flow and the entering feed water flow were unknown. These were found from a heat balance and a mass balance for the heater. In order to find the enthalpy corresponding to the pressure at the tur- » bine extraction points, a turbine expansion line with its end point was needed, and this was found by the method described in Reference 41. Figure G.1l shows . this pressure-enthalpy relation. 129, ) [ ] The amount of work at the turbine wheel, as found above, had to be corrected for heat put into the feed water by the condensate pump and voiler . feed pump.. This heat caused a reduction in the required extraction steam . - flow and & corresponding increase in the turbine steam flow and wbrk output, - The heat gained from the pump per pound of feed water was found by, Q BTU = work into pump BTU - work into feed water BTU - (6.12) 1b ' 1b 1b The work put into the feed water is found from, vork = (PpV, - P1Vy) BTU | (6.13) 1b where PoV, and P;V, are the respective products of the outlet and inlet pressure and specific volume. Since the change in specific volume is very slight, Equation (4) may be rewritten as, vork = (pp - p1) V; BTU ~ (6.14) 1b . The work put into the pump is found from, Ba _ work in = work out BTU (6.15) efficiency 1b : The reduction in extracted steam is found from a heat balance between heat from the pump and heat from the extracted steam, feed water flow 1lb x heat from pump BTYU W hr 1b (6.16) H of extracted steam BTU/1b 130, The turbine work gained due to the reduction in extracted steam is found from Equation (6.10). The corrected work done by the steam at the turbine wheel is; therefore, the work as calculated initially plius the work due to the correction for heat gained by the feed watef from the hot well and boiler feed pumps. The work at the turbine Qheel, as found abcve, must be corrected for exhaust losses in the steam leaving the last stage of the turbine and for mechanical losses in the turbine and generator. These losses are found by the method in Reference L, The gross electgic output of the generator is equal to the work at the . turbine wheel minus the turbine and generator losses, as described above. The method used to determine thé net plant output is described in the next section. 6.8.3 Auxiliary Power Requirements, Net Plant Output and Efficiency The method used to determine the groés electric power output of the reactor power plaht was described in the previous section. The net electric power output of the'plant is found by deducting from the gross power all auxiliary power required by the plant. Power required by the plant would 1nclude all pumps used in the cycle; i.e., fuel circulating pumps, sodium pumps, feed water pumps, and the con- denser ccoling water pump, and this power can be calculated. Additional power is required for other pumps, air compressors, ventilating fans, light- ing, etec. This power'is indeterminate for this study and an ‘arbitrary figure of 2.45 Mw (1 percent of full load) was assumed. -1€1- TURBINE GENERATOR: } 1000F . 2I5F \ 250,000 KW TCTF,3600 RPM Isalofi;?F :Is,ooopF ‘,%%%: jt8oopsig 15.6p 1800 psig, 1000F, 1.5 IN ¥g Abs~ oor 2 5/700,000 { | 5.850,000 ‘ 055000 1b/H- 1480.3H 1095 H / . \ M , GONDUGTOR COOLED, 26" BLADE ——fl N Ib /e Ib/mr AAA \ , ‘ 3 f ] FUEL| 2 3 3 245,420 KXW SAUL : ' 1050F|, * \ 48,000,000 |b/Hr J \\ | 94ca REACTOR fi00oF (O 675F N . ) 1024 H Na NG b 3.2°P 2 \T 9£,800 INTERMEDIATE HEAT zlag -~ 15/ ne EXCHANGER ol 8 [ 1.5 Hg N _mM Nl‘"’a; o padt I ~ . } { o Py A.1,600,000 REACTOR POWER-600,000 KW (HEAT) > DEAERATOR S Hwp ] Ib/Hr GROSS ELECTRIC POWER - 245,420 KW 1_- I :- HEATER T ;‘r. z| 2 PLANT AUXILIARY POWER-16,400KW tags ol 59 # 3 MY aloo NET ELECTRIC POWER- 229,020 KW a2l :gg- z|+ 2 olw Tl o e 17 - oFTo}" T~ OFTOY§ ~ SFTD SFTD § & 5FTD 450F AV 90F ANAAR BFP AAMA 2I0F| 150F| ... [SL7F W : 8Fe AT B T - 1950 p Ny 364.2h &2 fl[ 2000p ®a 8.1h a5 H7.9 h &6 §9.7h 430.1h 330F l / 380F 300.Th 270F 450F 364.2h 238.8h 275F 2IEF IS5F 430.1h 243.9h 182.1h 122.9h 600,000 X 3413 GR HEAT RATE = - = 4 TU /KWH 0SS HE E 545,920 8,3¢58TU/ 600,000 X 3413 : 2 : 8,935 BTU/XWH NET HEAT RATE ® ;45,420 - 16,400 229,020 S22 - 38.29 NET PLANT THERMAL EFFICIENCY 500,000 Yo NOTE : {1) BASED ON ZERO MEAT LOSS TO AMBIENT . {2) TWO HIGH PRESSURE MEATERS AND DEAERATOR HEATERS ZERO T.T.D. OTHER HEATERS 5F T.D, {3} ENTHALPY NOT CORRECTED FOR SUBCOOLING, , (4) STEAM EXTRACTION LINE PRESSURE DROP OF 8% ) ASSUMED FOR ALL HEATERS. FIG.6.2-HEAT BALANCE DIAGRAM FOR 600 MW REACTOR-STEAM POWER PLANT. V42251 : ONA-Y1~INHO 132. TABLE 6.7 SUMMARY OF FULL LOAD OPERATING DATA Turbine Throttle Pressure 1800 psig - Turbine Throttle Temperature 1000°F ) Turbine Exhaust Pressure 1.5 in. Hg Steam Flow 1.955 x 106 15/br 'Gross Generator Output 245 .42 Mw - - Auxiliary Power Requirements: Condensate Pumps : 0.33 Mw Boiler Feed Pumps S.34 Mw Condenser Cooling Water Pumps | 0.52 Mvw Reactor Fuel Pumps - 0.36 Mw Primary Sodium Pumps 5.98 Mw ’ Intermediate Sodium Pumps 1.42 Mw’ ! Miscellaneous 2.45 Mw Total Auxiliary Power 16.40 Mw ) Auxiliary Power as Percent of Gross Output 69'{% Net Station Output 229.02 Mw - Net Station Thermal Efficiency '38,21:_ Net Station Heat Rate 8,935 BTU/KWH 133. CHAPTER 7. POWER PLANT LAYOUT, OPERATION AND MAINTENANCE T.0,0 INTRODUCTION In order to determine the suitability of using a fused fluoride salt reactor in a central station power plant, it has been necessary to consider all components of the reactor and power plant system. A pfeliminary piping layout and drawings of a proposed arrangement of plant components have been prepared to illustrate the details of this design.. The resulting power plant ‘design is feaaible in size and arrangement. It wasralao necessary #o consider the operating characteristics; main- tenance and'econfifiics-to.determine whether such.a plant would be attractive to the power‘industry,; | 7.1.0 GENERAL PIPING ARRANGEMENT The-génefél'piping arrangement shown in Figure 7.1 has been designed to be safe, reliable, and to pro&ide flexibility in operation and maintenance. T-1.1 Primary Loop Piping The twenty-four primary heat exchanger bundles are connected to ring. \ héaders located aboie and outside the reactor vessel. Six sets of primary circuit lines connect the ring headers to the intermediate heat exchangers. The main purpose of this arrangement is to provide uniform cooling‘of the reactor in case one or.more.intermediate heat exchangers are out of service. The'pump for each primary circuit is located in the intermediate heat exchanger compartment. The piping passes through the primary shield at an éngle to minimize streaming of neutrons and gamma rays, and to retain the integrity of the shield. 13k, T-1.2 Intermediate Loop Piping The six intermediate heat exchangers are distributed uniformly around the priméry reactor shield. This symetrical arrangement minimizes shielding requirements and the cost of the containment vessel. | The six intermediate heat exchengers are divided into two groups, with each group cennected to 8 steam Qeneratoro' Thé heat exchangérs-in each group are connected in parallel to eaehvother inside the contalnment veésel, and a common line is passed tarough the containmémt vessel to the steam generatoro The connecting lines are designed to pr@#i&g equal firessure drops in order to bélance flow and heat load among the intermediaté heat exchangers. T.1.3 Steam Generator Connections The steam generators are located as near the intermediate heat gxchangera‘ 88 possible, outside the secondsry shield and containment vessel. The sodium pumps for the intermediate circuits are located between the two steam generators and are arranged so that one spare pump can serve either steam generator. 7-1.4 Steam and Water Piping- In general; the piping in the steam turbine cycle is‘represefitative of standard power plant practice. The differences are in the steam dump system, vhereby steam way bypass the turbine and be condemsed in the condenser, in the emergency cooling of the steam generators from head condensate tanks in case of a complete power loss, and in the method of attemperating or controlling the final steam temperature. 7.2.0 GENERAL ARRANCEMENT. OF PLANT COMPONENTS The general plant layout is shown on Figures 7.2 and 7.3. The major pieces of equipment are drawn to scale and gemerous allowamces have been made for pipe expansion loops and clearances. The mean diemeters of the primary and secondary . ) 135 circular shieldé are approximately 35 and 84 feetv, respectively. TheVSpace between the primary and secondary_shields is divided into six compartments by radial shields to contain radiosctive compbnefits of the primgfy and inter- mediste circuits. The containment vessel, as shown,-is TO feet in diameter and 110 feet in over-all height. The dimensions of the containment-vessel are approximate only, since fio calculations have been made to estimate the volume required to contain the gaseous fission‘products which would be reléaged in a serious nuclear incident. The reactor; primary heat exchangers and fuel dump tanks are located in- side the primary shield. The:primary loop sodium circulating pumps, expansion . tanks, drain and charge tafiks,'énd the intermediate heat exchangers are placed. in the six shielded compartments. All radioactive components and materials are . inside the sealed compartment vessel. The steam generators and intermediate sodium pumps are outside tfie con- tainment vessel and shielding, since they contain only non-radioasctive sodium from the intermediate loop and water from the steam cycle. Isclation compart- ments are provided around the steam generators and sodium pumps to prévent the ;:f* spread of fires which could result from serious failures of the steam generators. 'The turbine-genérator, condenser, feed-water heaters and boiler feed pumps are adjacent to the steam generators, as in a conventidnal.steam power plant. T:3.0 SHIELDING The dfiplex-compartmented shielding design for a reactor plant of this type depends greatly upon the method of maintenance which is to be followed. The basic premise’fér maintensnce in this design is that personnel will be permitted access to the containment vessel only during periods when the reactor:is shut down. . Based on this premise, the primary shield surrounding the reactor is /36 [] TaNK EXPANSION INTERMEDIATE AT HE EXCHANGER ——FLOWMETER CHECK VALVE SHUT-OFF VALVE 4 SAFETY VALVE & POSITIVE ACTING CHECK VALVE MOTOR OPERATED CONTROL VALVE -Of Y o2X Z COLD TRAP GLOBE VALVE DIAPHRAGM VALVE NORMALLY CLOSED VALVE EXPANSION' TANK ORNL-LR-Dwg. -15263 CONDENSATE STORAGE TANKS EMERGENCY COOLING CONTROL CONDENSATE EMERGENCY COOLING " STEAM RELIEF VALVE TURBINE o TUR BINE- [ .. p STEAM S OM LVE GENERATOR e . e , ) / N —DESUPERHEATER UNIT b ~ [l N “_o }—de NCE — THROUGH ] STEAM GENERATOR —)‘—N—N-d)gbo-i = A I STEAM - I MP l g‘:sm PRESSURE-TEMPERATURE l » O ! REDUCING STATION 4 J1 1 LOW LEVEL CONDENSER CONTROL: DESUPEREATER- JD Dé:::i-fl T surenreaten PIEDWATER i €7 . - HEATE 6] CONDENSATE L) $ PUMPS HIGH LEVEL CONDENSER X CONTROL: L FEEOWATER I e | REGULATING L A N ek ae VALVES o Ak Y 3 ATURAL 50 AJ_.'.LH F'... ' 1® I_i ® ©® CIRCULATION K s > RO W N o O oy DL BOILER PO o +— NOTE: FEEOWATER HEATERS OEMINER FOR PURPOSES OF ILLUSTRATION - A ONCE -TMROUGH 'STEAM GENERATOR FEEOWATER HEATERS ALIZER I SToRAGE AND A NATURAL O IR AR . " . . . = 1a ~ &, - D YR Y :_:A;: . S A . S 3 . o WESL IR s g T 2 v AT BE D o ~p @ 4T % FROM : o\ fo:]" ,v._/ g oLt % ."L > I ]——-)ruowmn ) N ¥ R« (5 TSI I SROURICN SR CARCRL S ® NEATER o. e." Y- . _'[ 2. i I‘ FUEL PUMP MQTORS e i s =N . g o . ° AR - . a - By .- W . ’ | j:" LA e P a8 ._flm NP AS . " VAL . % o 4 “No PuMP | e ok E:’\nowlumou:nvz SODIUM DRAIN - - o5 h AND STORAGE TANK REACTOR ’ e R & gy bl i Pl . - Ak 0 o' 20 o'd R ol T e b o sl N ) | \ | N [ ] LIt Sl PRSI I - 4-.‘ o '.-'.". ‘ y L - 0T ‘._' AT R BRI (T R ] A ek o 5 CONDENSER GROUND ‘ LEVEL 29ZCT- "Sma-¥d1-"INYO “LEY 3 No PUMP YRR WA P SR I oo 'e D 4 B £ /.9 r - . . \. “« P, 5 2 o R & ., T e s W S0 - o o™ X -5 TURBINE GENERATOR STEAM DUMF SYSTEM b ) L L L v ~—td el | N . . 3 » - B . 1 . > ey \u § oY B, o Al ‘o b e 3 Tl Na PUMP A ) Ot - O. . g % STEAM GENERATOR FIG. 73— PLAN VIEW OF POWER PLANT R TN et 7o conoenserf , < ! ) X3 7 VMAIM STEAM LINES c9zGI- ‘Sma-y¥T1-TINHO QT 139, designed to protect personnel against decay gammas while perforuing direct- maintenance within an intermediate compartment on the intermediate heat . exchanger or primary loop pumps without'draining the fuel from the reactor vessel. In addition, the primary shield serves as a neutron shield so that the equipment in the compartments does not become radidactive° Provisions are made for draining and flushing the sodium from individual, intermediate, heat exchauge? ¢ircuits so thet no radiocactive sodium is within the compartment where maintenance is to be performed. The radial‘compartment partitions serve as shields and provide protection from decay gammas from the .radiosctive sodium in adjacent compartments. This eliminates the necessity of draining all the primary circuitsfwhen-oniy one cifcfiit requires maintenance. The secondary shield,.located on the outside of the containment vessel, protects plant personnel from excess radiation exposure while the reactor is in cperation. The shield over the reactor will be made as a rotatiflg plug with an offset removable section for access to the primary heat exchanger and fuel circulating pump for maintenance. With this firrangement, the huge mass of the top shield will not have to be handlgd before performing wmaintenance. An aiternétive method would be to fabricate the top bhield;ifl sections and . remove only the sections necessary to provide access to the primary heat exchanger or fuel circulating pumps. The shielding shown in Figures 7.2 and 703.18 intended to be schematic only. T.4.0 MAINTENANCE Maintenance of equipment outside the containment vessel will be conducted in a conventional manner'since none of the equipment or materials to be handled will be radioactive. Special provisions, however, are necessary to allow - . ‘-].-lfl'-’ maintenance of equipment inside the containment vessel and shielded areas. The major items of equipment which will require special techniques are the - fuel circulating pumps, primary heat exchanger tube bundles, intermediate heat exchanger tube bundles, and primary circuit sodium circulating pumps. 7.4.1 Fuel Circulating Pumps Access to the fuél circulating pumps and primary heat exchanger tube bundles is through an opening in the top shield. This shield, which covers the entire reactor vessel compartment, can be rotated until an offset removable plug is centered over the desired location. The plug can be lifted with.the Erane, and the pumps can be removed by remote handling eq_uipment° Spare pumps will be kept on hand to replace faulty pumps. . 7-4.2 Primary Heat Exchanger Tube Bundles In the event that a leak should develop in a primary heat exchanger tube bundle, the sodium in the circuit will leak into the fused fluoride salt and dilute the fuel concentration. The rate at which uraniufi is added to the fused salt could serve as an indicator of tube leakage. 1t is felt that a means of locating the leaking tube bundle can be developed. Since extra heat transfer surface has been designed into the primary heat exchanger, ofie or more leaking tube bundles can probably be blanked oéf and left in the reactor, with little loss of power, until time is available to remove and replace the faulty bundles. The bundles will be taken out of service by cutting the feed ' lines connecting the tube bundle to the ring headers and blanking them‘off. The cutting and welding are to be done remotely, since the radioactivity will be too intense for direct maintenance. The bundles will be removed through the top gshield, in the same manner as described for removing the fuel circulating pumps. lhl o 7.4.3 Intermediate Heat Exchangers and Primary Sodium Pumps The intermediate heat exchangers and primary sodium circulating pumps are located in shielded intermediate compartments. Priér to méintenance, the radioactive sodium in the piping, pump and heat exchanger, in one compart- ment, will be drained and flushed out with non-radioactive sodium by remote operation. The equipmént can then be maintained directly. With the shielding provided, the sodium in adjacent compartments wlll not have to be drained. Sections of the top shield will be removed so that the crane can be used during maintenance of, these componen‘fis° T.5.0 PARTIAL LOAD OPERATION The ultimate removal of the heat energy from the reactor system requires the coupling of the fuel circulationvto the primary, intermediate and steam loops. Operation of the system at some steady state partial load can be achieved by variation of the full load design flow conditions in one or more of the loops, use of control rods in the reactor, change of fuel concentration with load, bypass control of the loop flows, fise of a supplementary heat exchanger, or any combination of these procedures. It was apparent that this multitude of fiossible combinations precluded a complete mathematical solution of the problem in the time available for this study. Therefore, the decision vas made to approach the problem of partial load operation in such a manner as to give results in the time available. This was accomplished mathematically by only studying the effect of varying the flow rate in the intermediate loop. Another approach was used which involved the reactor simulator facilities of the Oak Ridge National Laboratory. A description of the use of the reactor simulator, the circults and the results are given in Section 7.5.2 and Appendix J. 142, T.5.1 Mathematical Approach In a8 reactor system such as .the HRT or PWR, partial lcad operation is achieved by aliofiing the steam pressure to increase as load decreases. This method of control could only be used in a very limited manner for those reactor systems in which high steam pressures are achieved at full load. Of the many possible methods of control available, it was decided to'attempt a mathematical solution of the system with a variable flow rate ih_the intermediate loop. The reasons for this choice were the following: 1. Time limitations prevented anything but a single, variable study. 2. With the exeeption of the intermediate loop pumps and the.interu mediate sodium coolent flow, all of the pumps and loop flows would operate at constant values. 3. Temperatures would be held to reasonable values. L. The problem of safely varying fuel concentration with load is eliminated. 5. The solution thus obtained is also'useful if-bypass control of the intermediate loop flow is utilized. 6. Control rods are not necessary for the achievement of steady state . partial load operation. 7. The mathematical equations obtained would serve &as & basis for further parameter studies of this system. In order to obtain the mathematical solution to this problem, several simplifications and assumptions were made. It is believed that individually, or in the aggregate, these simplifioatioos.do.not invaiidate the analysis over wide load~range5uf The mathematical analysis i1s given in Appendix I . and is based upon the following considerations: 1. The flow rate of the fuel in the resctor and the sodium coolant in - the primary loop remain constant w1th load -2, Physical properties of the heat transfer fluids and the materials of construction remain constant for the temperatures of operation..: 143, 3. Sodium heat transfer resistances are not the governing items in the over-all heat transfer coefficients and do not vary with load. This assumption is the most limiting restriction of the analysis given in Appendix I and affects the load range, for which the derived equations can be applied with reasonable accuracy. 4. Based on the‘foregoing,.the_0veffall héat_transfer coefficient in the primary heat exchanger, the intermediate heat exchanger and the boiler remains constant. The constant over-all coefficient in the boiler further presumes constant average boiling and scale coefficients. 50- The final steam temperature and pressure arg constaht with load. 6. The average fuel temperature is constant with load} T, The feed water temperature is h50°F at all loads. 8. At very low loads, some of the steém generated is bypassed directly to the condenser. This does not enter the mathematical analysis directly, but is an important element of control. ' ' Equations describing the behavior of the system at steady state, partial load operation have been obtained for both the naturallcirculation boiler with separate‘superheatér and the'once-through boiler. These equations are given in Appendix I. It is noted that for the natural circulation boiler witfi sepa- rate superheater,‘consideratibn (5) above, involving the constancy of the final steam temperature with lcad, was abondoned at the end of the analysis for this unit. It was found that a &ingle variable control cannot séfiisfy all of the restrictions placed on the systefi in this section. 7.5.2 Reactor Simulator Analysis In order to determine the partial load and'fransient.operétifié conditions of this péfier plafit design; the circulatihg fuel find:sodium loofié were simfilafed on the Analog Reactor Simulator at thé Cék Ridge National Labdratory.' The analysis was limited to the design utilizihgfihe'qnéefthrofigh steafi generator. The time constants and heat capacity of the reactor and sodium coolant loops are shown on Figure 7.4. The simulator constants were computed by 1hk, -Dr. E. R. Mann. These constants and the simulator diagram are included in Appendix J. ‘The purpose of this simulator analysis was to determine whether the . reactor would be stable at partial load and during operating transients, and to determine the minimum load at which the final steam temperature could be - maintained at design conditions. The analysis indicates that the reactor is stable down to at least 17 percent of full load, the lowest partial load con- dition investigated. A final steam temperature control device was simulated by providing an afixiliary heat exchanger in the intermediate sodium circuit, where heat is transferred from the hot leg of this circuit to the cold leg. By varying ':the flow through this heat exchanger, the heat input to the steam generator can be controlled. The results of the simulator analysis indicated that the design temperature of 1000°F could be maintained down to 50 percent of rated load. Below this load, the temperature increased gradually, but this excess temperature can be controlled by the use of an attemperator, which is provided in the steam cycle. Figure 7.5 shows the steady state, partial load temperatures without the use of the auxiliary heat exchanger mentioned above. The steam temperature - increased from 1000°F at rated load to 1110°F at 17 percent of rated load. The feed water temperature entering the steam generator is MSOOF at rated load and decreased as the load is decreased. The resulting temperature differ- ential between the sodium and water at the cold end of the steam generator is considered to be too great from the standpoint of thermal stresses in the tubes. For these reasons, it appears desirable to include the auxiliary heat exchanger _ in the circuit. © 1k, Figure 7.6 shows the steady state, partial load temperatfires with the auxiliary heat exchanger in the circuit. The control signal used to adjust the bypéss flow through the auxiliary heat exchanger circuit was set for plus or minus 2°F of design steam temperature. The exact setting of these limits would have to be studied carefully to eliminate hunting in the system. The close tolerances used on the simulator are not required for satisfactory steam turblne ovperatlon. The curves on Figure 7.6 indicate that thc steam temperature remained constant to below half load and increased to only 10909F at 17 percent of rated load. 'These temperature conditions are considerably better than those obtained without the use of the auxiliary heat exchanger. .The reactor power and tempe;ature transients, with the auxiliary heat excfianger 1n_the infermediate loop, are shown ifi Figure 7.7. During this S test, the load demand of the steam turbine . was reduc;d frém rated load to half load in approximately 15 seconds. The initial effect of this load demand 1is éeen in the rise of the sodium temperature entering the steam generator. It is interesting to note how the reactor power and other tem- s peratures varied and leveled out in a relatively short period. A series of these tests vas:‘performed and all showed similar results, which indicate that the reactor and system are stable during load changes. It should be noted that the rate of these load changes is conaiderablylin excess of that normally occurring in conventional steam power plant practice. These same system transients withbut the auxiliary heat exchanger in the intermediate loop are shown in Figure 7.8. 7.6.0 ECONOMICS A complete economic study of the proposed power plant has not been attempted. It must be realized that the large capital expenditure for a nuclear power plant' 74 1150F 1.3 Sec. 1150F 1100F L Sec. 1100F 1000F . - ‘ ’ A -[1200F VWA | —— 1.6 Sec 1.1 Sec. 1l sec. 3 sec 1125F , - : X Vo 1050F S — v, 1000F 2.3 Ssec. 1000F 675F 5 Sec. 675F L5OF * REACTOR INTERMEDIATE HEAT EXCHANGER BOILER-SUPERHEATER _ ShelX Tube- Shell Tube AT =150 F AT =150 F = hos F AT = 425 F ' = = - | = 887.5 F T, avg - 1125 F Tavg = 1075 F Tavg 887.5 F Tavg 7.5 trgel (in core) 9.3 sec M © 21,990 BIU/ F tNa (exchanger) 1 sec t sec ®ruel (in exchanger) °°° 5¢ Hopel1 = 170 BTU/F - Na (boiler) ~ S | . ' t _ % el (in transport ) 2.8 sec Hpipe = 9,190 BTU/F Na (transport)= 9 sec Hovel (core) = 397290 BTU/F “xa (tuve) = L6 =€ Hy, : 1320 BIU/F 6,020 BTU/F ’ - Ha (exchanger) / ty (transport) 3.6 sec Hy o = 190 BTU/F " = 1,800 BTU/F 2 SN S.S. (exchanger) - t e (she11) = - sec = 20000 BIU/F = 20,800 BTU/F pipe Hevel (exchanger) . FIG. 7.4 TIME CONSTANTS AND HEAT CAPACIT® OF COMPONENTS JN ONCE THROUGH STEAM-GENERATOR SYSTEM = 550 F AN 6LTQT="*3Mq=TT~TNHO am =1 W 1 : . e — | L7, 1600 |IBEE ARE TR | Figure 705 S SR SEE - 500 to Steam Generator Sx . Final Steam Temperature oy DecS SRS e ORNL~LR=Dwg.~18180 ?:2'%' 1100 TR LLL INERN EENI Y51 0 Y O { Reactor Load - MW Temperature of Sodium from Reactor - Partial Load System Temperatures' - Without Auxiliary Heat Exchanger 1 O B g 1%} S S L Ligg O : O I o) & O =~ Gb 3HUH ] 3 B " & g & o H =) ] G 2 E+ ] o i ¥ 1 le 19 1 - b - e L i I . 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IT..x::...O.DL ® ey T (annn H EEas wum T OHE OHH G T () » 8 w ~ O o~ 0 u.%.HH /nnw HE TN O HF WNWHF O HH m/nnu WHHH J.nhu/nn o~ B+ % () +2 F Tt A TTT 5..Hnuu Snnuu. ummm I“Tnuuu 3Hn_n_ 3.n.m.MMm 2.Hu o _ . &4 o A Y 1 | u r T 01 G 66 1 £y O g i HHH HHY , - 3 o o) ; o 5 ® o g : ® o H + = 3 4 K ” o~ M e 8 3 £ - : : O o 3 > : + g Sagks! O L~ m g % a5 H & - o w g ; 834G o i ¥ o P & -~ 5 A ; Yt g 2 . = e HHH Le) Sy o 8 o o sia 5 fhie e s 3 " o 3 g 3 gt = STt L ¢ Ry ..m A m © T M FHHH m : (o] : ...ub n | m m m a.on./u.* o _ [ o - @ | G | | . Hned sy @ b i o o _ 8£A g HE O + a | © @S ©T KA I X i m. =k Bl A9 =t £ . B 2Rl H e 8 : I H 2 m + 42 T -m m.. : o o m o n T ® g i SESSS HE Qe : . ‘ S B e HHFH L . M inaa 82! H . s s L ENE i .Hw [Em. = N Mog HHHETH 5 0 L et o O T T B mmm 1N HHQ HES HHEQ HES O o HHE HHHH _n HED BSOS HEW HES HHED HH S QEESHR TS, e e : = O :j___T-TL_B HH O~ t~ 5O O T HHH H-H : | | :“: A = eanjexedute], EEE i51. is one of the major factors which influences the cost of generating electricity. These capital costs would have to be evaluated for various arrangements of the basic design parameters in connection with the resulting reliability, ease of maintenance, safety, efficiency, etc., and a careful analysis made of the re- sults before such a study would have any real meaning. 7.6.1 Factors Requiring Additional Investigation These studies would have to include the following considerations: l. The selection of internal or external primary heat exchangers would have to be evaluated carefully. It is considered that fuel inventory and space requirements have been minimized through the use of the internal arrange- ment, but these desirable features would have to be balanced against the over=-all comparative costs of the two possible arrangements, their reliability, ease of maintenance, etc. 2. Two sodium loops have been included in this study and are considered a necessity, resulting from the internal arrangement of the primary heat ex- changers. It is possible that the intermediate sodium loop could be eliminated if external primary heat exchangers were used. However, again, such an analysis would require considerable investigation to determine the lowest cost and most reliable system. 3. If the reactor type using internal heat exchangers, as proposed in this study, is considered to be desirable, further studies would have to be made to determine the optimum arrangement of intermediate heat exchangers and steam generators. It is quite possible that a unitized arrangement using an equal number of intermediate heat exchangers and steam generators would result in many desirable operational and economic features. This question could not be resolved without a complete analysis of the entire system. 152. 4. As mentioned previously, the steam cycle arrangement also would require much investigation. The question of steam conditions, reheat or non-reheat, tandem compound of cross compound turbine generators, feed water ) heating cycle, turbine exhaust pressure, etc., would have to be settled from a careful economic analysis. » 7.6.2 Approximate Economic Analysis In order to examine the approximate economics of this plant, Reference 37 was used to estimate the capital costs. The capital cost for various types of nuclear power plants listed in this reference varied from $183 per kw to $450 per kw. It appears that the fused salt reactor and steam power plant discussed might cost in the range of $250 to $300 per kw. These two values will be used in arriving at a powér cost. For 240,000 kw generator output, 80 percent load factor, the total kw-hr per year is 1.7 x 109. There will be no operating cost for fuel processing. It is shown in Section C.7 that the original fused salt charge can be used for at least 19 years before it is economically desirable to add a new charge of fused salt. During this period, there is no fuel processing and uranium is added to over- ride the effects of fission product poisons. Using Reference 37, a figure of 1.2 mils per kw-hr for operation and main- tenance is cited for 10 mil per kw-hr total power cost. Since this is the highest value used in the reference, it will be used in this report as a conservative figure. The inventory charge is based on 150 percent of the hot, clean critical mass of U-235. An additional 50 percent is allowed for poison override and . fuel on hand. (See Section C.6.) Using 4 percent interest charge and an 153. assumed cost of $20 per gram as the cost of U-235, the inventory charge is 0.2 mils per kw-hr. Since some atoms are destroyed and no heat is derived by resonance absorp- tions in the uranium, the fuel burn-up will be greater than 1 gram per megawatt day. This figure is estimated to be 1.3 grams per megawatt day for this reactor. Based on this value, the fuel burn-up at full load is 780 grams per day. For 80 percent load factor, the annual burn-up is 227,760 grams. At an assumed cost of $20 per gram, the fuel cost per year is $4,555,200. This amounts to 2.6 mils per kw-hr. These data are summarized in the following table: TABLE T.1 ESTIMATED POWER GENERATION COSTS $250/kw $300/kw mils/kw-hr mils/kw-hr Capital cost 5.4 6.4 Fuel processing 0.0 0.0 Operating and maintenance 1.2 1.2 Fuel burn-up 'l 25T Fuel inventory 0.2 0.2 15k, APPENDIX A PHYSICAL AND THERMAL PROPERTIES Figures A.1l through A.10 Tables A.1l through A.3 o ORNL~-LR=Dwg, ~18184 55 I L 5= 1 { 51 1T T | 1 1 m) w ] - ) : ) f, H H i - . T e i 1 H ) mE 4 E ] = I 3 g = L o m H By En sl DU kasE mnil SAPP — T o o/owrry I QEu = | (@] H o, A Mr &8 g i & 5§ A N = L ; 1y 0 O ~ D @ I H N 2+ | — 5L 9, m ee | = __ | o — 2 O : m O - rHH &u W i L) ] 1 R HE 3 .A - 11 7/ & O B E MS; Y I o S Ee > S o 3 ! 1A - 5t R v.(.@w~ 0 o SESE £ o g =) | 1 = . 1] LD % & T .— LM | o) ! B i 0 o SERRsna TEseaEaE & S _ H il 9 S T O & EEgEmams : + , AmBHEnNEaE m i i tH iy | & mmw_ SEESESEEEEISE _ P ¢ - S _,:WHW Sne SdE ffi—.l o [Lxfl.mL(HL.. B ._,E “ , H 4v 1 rif.“-. 1 __ %..— o\ { ENE ot : e B~ 1 T _, 1 BREEE S i Ll o BEEERERRNE GRERE ) A1 H AR A EERA N \ No) i \ \ \ o \ 5 o ,inn,r- HH \ t ] 2REEELARNR Bp i B I T 1 0w SN \ A = 1 1 1 i { t i\ “ ~ A\l - S - | (35 e 4 o [ Lfm. __ w i ] n nam - t L \ ~H \ 3 T e 1] I | { -+ ) - - - N HEH | H i ST9915 SSaTUT®Lg ] EEE ul L { [ i . T Ann H HHH t - - B gaseE mEE b i Ho O o w0 O « 3 & = ee R S a __;_—__b__u,,_H_J‘L_\L________~_____vLLI_\_LLI_LI-.r____;__._._LV__.>A._IIT~»IL P e Y/dy 34 *IH/1IE-3TATIONPUO) TWISUL JO 3USTOTIIE0 ORNL~LR=Dug,=18185 156. ::::?I : = / x # i A St R f. Sahasa F : H = ;’ 1 l ;_ i ; i | H o e . asi £ : . Figure A,2: Source: Thermodynamic Properties of Steam by 3 Keenan and Keyes; John Wiley and Sons 0 i o 157. ORNL=-LR=Dwg,.=-18186 600 . S S W1 T [ 1 I stila T HH{[II'IIII( Ty RS SESENANERE l'i—H':}AII] 08 19 I O McGrawH ) 5 O I O 0 O | 0 1 A I A O 0 W ) T 1 I H {0 0 N U 6 0 O I A T THE We , page 484 I i I . McAdans, (0 I I 0 1 5 0 0 I T T T {Figure A,3 500 ty 1 i 1 1 + T 400 HE 1 1 NN T ==l e Viscosi uctivity - Jele 300 rnal- Co: The 200 a 3rd. Edition L] VISCOSITY & THERMAL CONDUCTIVITY OF SATURATED WATER Source: Heat Transmission, 100 EEEE M T A b T 1 T i |-Fi- T 1 | T + HHh T o N s i - 4 | i ) R FrEren 063 HF d p *QT-£3TSODSTA - e d, *3d tIg/mag £yTATIONpUO) TRUILY] Tenperature-oF —_— : ORNL=LR=Dwg,.~18187 158. DENSITY OF SODIUM, POTASSIUM A SODIUM POTASSIUM ALLOYS Source: Mine Safety Appliances Co. report dated Dec, 15,1953 curve @-202 58 56 ($) o, udl Q’Q’% N N .{‘%Q\%&Q I Yo Q" K75 O0r'9, PAes AR (D %, % 52 AR 78S o .94_966‘,}.0 ! < N %0 M . 48 G, 04 & ): 0,28 ¢§b = | - ; .b » E 50 F?JI/‘ HES oo "“‘%0 Qo‘? [ : N R = 1 1 % a‘ i ‘;E:: B éQ (] oD N i K> S Ky (?Q’ 5‘5 g 2 O R, 25 K [} % &> X, R4 g \e N 00 Q’é : ¢ 1,8 2 RS0 Y (=] Q Q\? T 0 o3 < 4 % (°4 Jo2 <, W, 6 log f = Q\,IA QJ & Q‘[ O) Q, § QJ E 16 o do : 4 08, . ¥ Qf Iy ! Figure A.l I-I-Z i 0 1 2 3 L4 5 6 7 8 9o 10 11 12 13 1 Temperature-100°F ORNL=LR=Dwg,.=18188 l . VISCOSITY OF SODIUM, POTASSIUM AND SODIUM POTASSIUM ALLOYS Source: Mine Safety Appliances Co, report dated Dec, 15, 1953 « L6 curve G-203 - 2nd Ed, Liquid Metals Handbook ) Journal of Metallurgy and Ceramics, Issue #3 May 19L9 ')-Il-l— E L NRL report # C-28I L2 = . Lo ! - 38 ? ) : : - 36 | : i i | e e .3 , N . 32 521 128 }*\'O' [:;:% i « 30 o e I : i o ERR IL E .28} @ \ A i + IR & 26 3 o o F nhH 5 . 2)4,[ 1 SEAtaEmEs E B i o i: 22 F O b F O B Y HH - o H - ot LT I 020 fit - 2 .18 I ! 5 J{ - b 1 - 3 e .16 HHEE £ N H 5 & 1 B ] NG : (B , s e FH - ¢12 B set . odium * iss AT 2 : ; - HHH g i 156 NakK of Figurs 4.5 = 578 NakK FEEdE Potassium o HHHH IR ER R S s S T e e T L o] Temperature-100°F Thermal Conductivity-Btu/hr, sq. ft. °F/ft. Temperature-100°F ORNL=-LR=Dwg.=18189 160. THERMAL CONDUCTIVITY OF SODIUM, POTASSIUM H AND SODIUM POTASSIUM ALLOYS = 50‘ Source: Mine Safety Appliances Co. report EE; dated Dec, 15, 1953; curve G-200 - L8 ' 1.6 Iy L2 110 38 Sodium- 2nd Ed, Liquid Metals Handbook HHH KAPL lemo LFE-6 ’ 19 51 MSA Tech, Report i 36 SEnaaE: H H 34F ! 32 H 30 : 28 1 £ " +Potassium~ 2nd Ed, Liquid Metals Handbook 2 ] HE " 2 22 i - 20} s Figure A,6 18 b o 16[0T 2 56 NaK i | - SR 8 NaK 1y MSA Tech, Report # 2l l2o0 2 2 3 L 858 6 7 8 9 10 1 12 13 1 SPECIFIC HEAT AT CONSTANT PRESSURE- BTU/[°F LB '/Z/' % ' SPECIFIC HEAT OF STEAM AT CONSTANT i IR ERENE NN NN EEENE N SN NN : ' PRESSURE FROM STEAM TABLES. | L ot 7 i = T - NN N SN SN SNEEE N i 1] - Figure A,7 ........................ Sources Thé;nddinanic Propefties of Stean by 0 100 200 300 400 500 600 700 800 900 (000 1100 TEMPERATURE- F 19T 0618T="* Anq=gT~TNHO THERMAL CONDUCTIVITY - BTUju — . THERMAL CONDUCTIVITY OF STEAM BASED i ON CALCULATIONS OF GRANET AND GOULD.. 004 0. 0025 : a0 | s : 0.02C 00§ 0010 0005 FlgureAB f 2 T Source: Trans., ASME, Aug. 1950, Vol., 72 pp. 767-778 0 100 200 300 400 500 600 700 800 900 K000 |00 | ) TEMPERATURE-F 1200 *291 T6TQT=*3Ma=YT~TINHO ol olo - Q0 o =7 m‘ g7, Too? - ] i -4 O.C" = ! L4 >— = . "J REEI >4 ’.—. I Nk —QG 28 T ) : O TP O oo : 2 t > | 003 8 uJ 1 - : H 30.02 : Q 80.0’ HHH < Figure A,Q [ VISCOSITY OF STEAM BASED UPON DATA OF TIMROTH. o FHHHFRHH R P FPPFFFPRFEEEEERET - Source: Petroleum Refiner, May 1953, Vol. 32, HHi No. 5, pp. 179-181 g O 100 200 300 400 500 600 700 600 900 000 10O 1200 TEMPERATURE-°F €91 26TQT=" SMq=NT~"TNHO 164, ORNL~LR=Dwg.,=18193 i rl a2 X by =l < — Pl o = 0 o 4 ! Q 0 mc 9 8 © “ 39 = R L ok o I g &8 9] 4 & o wy - il hoo = m o+ e E . m_m mman B S5 e = O o o = © g4 munuo 4y A ©° =9 Z . o o m m ,HH..M e N HH g L & 0 +HH g mu.. mzan ] 2 H © m ~ HH & 3 o w3 m “ o~ L o &8 8 8 $ § 8 8 8 % g8 8 n ~ ~ HH ~ ~t o on o on o0 saaun o UOUT/d;=»%d *IH/M3d ~£3TAT00pUO) TemISYY 1 ui Source: TABLE A.1 . Specific Heat Bx/1b - OF COOOO0 C.O0 00 oRoNoNoN o] OCOO0OOCOC OO COCOC ooNeoNoNe OCO0OOOO0 - 3117 .3115 .3113 -3110 3108 .3106 310k 3102 3100 3098 .3096 .3094 .3092 3090 .3088 . 3086 . 3084 .3082 .3080 .3078 - 3077 - 3075 .3073 .3071 . 3069 . 3068 . 3066 . 3064 .3063 .3061 . 3060 .3058 .3056 3055 3053 THERMODYNAMIC PROPERTIES OF LIQUID SODIUM Lee, John F., "Thermodynamic Properties of Liquid Sodium", NUCLEONICS, Vol. 12, No. 4, page T4, April 1945 Enthalpy Btu/lb 1k1. 143, 1l , 146, 148, 149, 151. 152, 15k, 155. WMo\ OWV\VO F® }—l = N ONOWVOF WVEFODWE® WU —JHOGHW 165. Table A.1 (continued) Temp . Specific Heat Bx/1b - OF 0.3052 0.3050 0.3049 0.3048 0.3046 oNoNoNeoNe OO OCOO COOO0O0 COO0OO0O0O W (7Y (WY) (oY O o o o | = o n o — n \O oNoRoNoNe () O O 3 166, Enthalpy Btu[lb 195.7 197.3 198.8 200.3 201.8 167. Table A.l (continued) Teup . Specific Heat Density Enthalpy Op Bx/1b - OF 1b/ft3 " Btu/1b 875 0.3004 52.21 256.2 880 0.3003 52.18 257.8 885 0.3002 52,13 259.2 890 0.3002 52.09 260.8 895 0.3001 52 .04 262.3 900 0.3000 52,00 263.8 905 0.2999 51 .96 265.3 910 0.2998 51.92 266.8 915 0.2998 51.87 268.3 920 0.2997 51.83 269.8 925 0.2996 51.78 271.3 930 0.2995 51.7T4 272.8 935 0.2995 51.70 27h.3 9ko 0.2994 51.66 275.7 9hs 0.2993 51.61 277.2 950 0.2993 51.57 278.8 955 0.2992 51.54 280.2 960 0.2992 51.48 281.7 965 0.2991 51 .44 283.2 970 0.2991 51.40 o84 .7 975 0.2990 51.36 286.2 980 0.2990 51.32 287.7 985 0.2989 51.27 289.2 990 0.2987 51.23 290.7 995 0.2988 51.19 292,2 1000 0.2988 51.15 293.7 1005 0.2987 51.10 295,2 1010 0.2987 51.06 296.7 1015 0.2986 51.02 298.2 1020 0.2986 50.98 299,7 1025 0.2986 50.93 301.2 1030 0.2985 50.89 302.7 1035. 0.2985 50.85 304.1 1040 0.2985 50.81 305.6 1045 0.2985 50.77 307.1 1050 0.298L 50.72 308.6 1055 0.2984 50.68 310.1 1060 0.2983 50 .64 311.6 1065 0.2983 50.60 313.1 - 3 O o g w \n o \n \Nn (VN = = O Table A.1l (continued) Temp » Specific Heat Density Enthalpy Op Bx/1b - F 1b/ft3 - Btu/ib 1075 0.2983 50,51 316.1 1080 0.2983 50.47 317.6 1085 0.2983 50.43 319.1 1090 0.2983 50.38 320.6 1095 0.2983 50,34 322,0 1100 0.2983 50.30 323.6 1105 0.2983 ' 50,26 325.0 1110 0.2982 50,22 ' 326.5 1115 0.2982 ) 50,17 328.0 1120 0.2982 50.12 329.5 1125 0.2982 50,08 . 331.0 1130 0.2982 50 .04 _ 332.5 1135 0.2982 49.99 334.0 1140 0.2982 | 49.95 335.5 1145 0.2982 49.91 337.0 1150 0.2983 49,87 338.5 1155 0.2983 49,83 340.0 1160 0.2983 49.78 341.4 1165 0.2983 : 49, Th 3L2,9 1170 0.2983 49,70 34k, 4 1175 0.2983 49,66 : 345.9 1180 0.2983 49,62 474 1185 0.298L ' 49,58 348.9 1190 0.298kL 49,54 350, 4 1195 0.2y84 49.50 351.9 1200 0.2984 | Lg.45 o 353. L 1205 0.2985 , 49.41 354.9 1210 0.2985 49.37 356, U4 1215 0.2985 L9.32 357.9 1220 0.2985 . 49,28 359.L4 1225 0.2986 : L9 .24 360.8 1230 0.2986 49.20 362.3 1235 0.2986 49,16 363.8 i240 : 0.2987 49.11 - : 365.3 1245 0.2987 49,07 366.8 1250 0.2987 49.03 368.3 1255 ‘ 0.2988 - 48.98 369.8 1260 0.2988 48,94 371.3 1265 - 0.2988 48.90 ~ 372.8 1270 0.2990 48.86 37h.3 1275 0.2990 48,81 375.8 1280 0.2991 W8.77 377.3 1285 0.2991 48.73 378.8 1290 0.2991 48.68 380.3 1295 0.2992 L8 .64 381.8 Type Analysis Carbon Manganese Phosphorus Sulphur - : Silicon Chromium Molybdenum - Other Elements Minimum Physical Properties Tensile Strength, psi Yield Point, psi Elongation, % in 2 in. Max Brinell Hardness Creep Strength pounds per sq in. -~ Rupture Strength 169, load in pounds - psi which lead to rupture in - 100,000 hrs c—— TABLE A.2 ' SELECTED PROPERTIES OF STAINLESS STEELS Source: Pressure Tubes and Piping - Timken Roller Bearing Company Type 304 Type 347 Type 316 0.08 max 0.08 max 0.08 max . 2.00 max 2.00 max 2.00 max 0.030 max 0.030 max 0.030 max. 0.030 max 0.030 max 0.030 max 0.75 max 0.75 max 0.75 max 18.0-20.0 17.0-20.0 16.0-18.0 - Cb lo X C min, l:o max MO 290“’390 Ni 8.0-11.0 Ni 9.0-13.0 Ni 11.0-14.0 75,000 75,000 75,000 30,000 30,000 30,000 35 35 35 200 200 200 Rate of 1%, hrs Rate of 1%, hrs Rate of 1%, hrs 100,000 10,000 100,000 10,000 100,000 10,000 800°F a-- - - -—- ——— -—- QO0CF --- --- - -—- --- —— 1000°F 10,700 18,000 --- - 1k ,750 2k ;500 1100°F 7,900 13,000 14,250 21,000 12,000 21,000 1200°F 4,300 8,000 1,750 5,300 6,800 14,500 1300°F --- --- 660 2,000 4,300 9,200 1500°F 1,450 2,850 -—- -—- 1,800 4,200 " Load Load Load 900°F -- -- - 1000°F -- -- - 1100°F -- 16,000 --. 1200°F 7,200 6,200 12,500 © 1300°F 3,700 2,800 8,900 1500°F 1,700 -- 1,600 Source: 170. TABLE A.3 SELECTED PHYSICAL PROPERTIES OF "L" NICKEL Private Communication, Jr. J. J. Moran, Jr., The International Nickel Couwpany : Creep Properties Condition: Cold Drawn-Anncaled Stress, 1000 psi, to Produce a Secondary Creep Rate of: 0.00001%/hr 0.0001%/hr 1%.0 25.0 7.5 13.0 L.3 7.3 3.7 6.3 2.5 h.5 2.1 3.7 1.2 2.1 0.5 0.9 Ruptiure Properties Condition: Cold-Drawn Annealed Stress, 1000 psi, to Produce Rupture in: 1,0C0 hrs 10,000 hrs 100,000 hrs 27. - - - 20.5 -- - 17.5 14.0 11.5 - 1.5 11.3 8.7 11.3 8.7 6.7 7.8 5.4 3.7 - b7 g APPENDIX B DESIGN CURVES Figures B.l through B.k 17 32— HEAT TRANSFER TO STEAM OR SUBSATURATED WATER IN TURBULENT FLOW (INSIDE A TUBE) : 600 . T B e Gk et i | F H x®" 2 e 4 H b { .2 - o E & na®° = ¥ ( w000 a )° $ : > £ 100 o®” . W - fluid flow rate - 1bs,/hr,t el e d - diameter - inches i b‘:;]j a - flow area - square inches | o H Derived from Equation 6,2 300 I]ll[I.IIIIIlllll(llll"ll|lil!llLlIllllll!‘lllllll!ll[llIlllllllllll :'GOO Physical Properties from Appendix A S iyin e J. . %*’ [ o Go %bo t 1 % i e Do My Yo 9,(’ 1 & 'ZO u‘i: Gj TN 4 200 K)oc T %2 i "{I Ll‘oo p Q‘-{ - }%9'4 T »“’l.’ 7_iymi. % .?ie “ --1 'I,. St LT 1500 pgi1g T Bepassaansett : 58 - aasail 1_!' j:"Zo H_H_IL JE EEE ] ii”i H““.“ } 100 EEEEEEE WS Average Bulk Fluld Temperature Fl‘igureB.'T!‘. 200 300 }:00 500 600 700 800 900 1000 1100 1200 H6TQT="3na-UT=TNHO ‘el «18195 ORNL~LR=Dwg. 173 o — t » T L r T % m ¢ 111! TEIS s HE e i Rk kR S S8as] sasssasaat istasias 3 2 '8 Babdv S r* t e + 2 r.m . WL 1] .‘.TT ss 1SS bs et g et 8 Eucas T HH G R EHAHS | H 2 v+ Lv.l H & m# I‘ T»'n“” _'MH 4 mrH fik. HX‘ y asd .H - T o 2 sz do 3 it R 520 01 T 2 6 83 e8s B H H WAL 1] H11 11 1] 13 A FROY EASRRERER: 22T 8235325288+ $3858 | spsgsdatal fhisy g 1 f cOu o d + 1+ » f s it r*m... m. H b b 5 LTY,P 4 : T : T ? ) greg o A~ m T ._Yw . 8 B 11 ] .PUm NMI]ULIJ .‘II it UH IM.. t HA” : JEBES 8 @ (1 a3 s - +4- 44 4 [ (1] 4 Lfi ‘4 ‘ALfixvv i a8 u + “ m a v .11.” $ b4 4 2 H-41 11 =g 8 +11 T H SesEang § 1 o H Y 1 9 sEistatad kil : P Ht AR : m £ i : ABEE THE . R il + i 8 s i L : m 1 T ug g “% S5 apar .M ab @ O ] o 17 II‘.T BuE a8 - r ‘ ‘ 1 .mu-u ‘ ”“r 1 ° SpE0E 898 ‘W & 8 o o 7 - o w r ufl @ 0 x 588488 sggsase ¢ 2 e ) ° ....111+ e 2 e 4*1 -L. w -4+ = t n fl . B t . rfi 4 4 a T I g L7 » 83 sEadsgasid & [HHH Baisd T g “ HH 49 © ; .m% i Tf s 88 : .W..flm T s fus: M ! w : 4 e o on sEss duiag bet .m [T - 111 H . Tt By iSufuas ez G e ’ w o = L gefges m (111 H HTH ey 8 s =& B HHEIEH R R o o B : H : i b - H b HE == HHH 5 @ 1 (o) w T H 88 §s s M - sass r F 3] i 0flp4” H a3 BSEE: 1t T Q SHHH Y H 53525 8SRaE sasas shunk o1 . - : S THE 8 e PRI L S 3 111 SA T e H - o 8! - . 4L PR R psand agws lww ] LH] . M ® - - 1A ] H %u |8 gRBET Ba T 1 ReShfasoateass: g r.nm 0 T s FHH H ;fi,, I HHET rmL ame: L4 fi:\ L + m.. 1 S sga 1 1t SEeS aut s H rb - R Y 2 T | T ¢ Jaibbsl sy ax TR HTHEH H i H A HH HH E LR 3 T » HHH HHtt Shusy Bpfnan ans LABRE B i R 3e s ) P 4- 41 a8 i n qstdateess H w Q0 i H . isdudbs L A S : N Az £ ghos ¥ + .ne aT WIL.I rHLT. AI (11l nL [o10} HWHw uvA E JIHmI X ] m S3i2eerapasgs o g o B HH s 2228 548 b4 40 Tt ..AM A E e E n *‘rv.fi 1 - ° 1] Xfiv I ] IBEE 6 o 1 f 41 :.mm w T o) aESH § 1 4444 an ] Sagndunn o .fl» T 2533 SRR IRNLSE: H- — oo HH L : ] S 55 S A 3 e . f _ SisEstisnEt ke 22 rL Mun [ fi t ] b ® B 0 B = ISy & 1 2 B 1 ) *M. H H 'm i Iur LB 5588 . Edgss T o s U $ N.L. Yw;.. IRER! ‘m an + sygas BERSRS F Y — 4 «.n“ - w%: ww #WL HH © m ssus HHT H T E k5 1 ' o r u ufi . L [ 1 fESpE suEs + 8 IU“ w... . ..*“I LL. w = fil i E vg 11 - r 11 pyunet 4 [ 11 L -1 HHH featas 11 T ] -+ttt Ht 11 aw e R 251 s el i e =352 geafiaadisssates HH 1 Heks - ade 1 sdsggsss 1T + N z t4 4+ .*.‘L 1171 (4 e . L +4 e ® 1H ,&A } ] 1 @© 11T ] sduuns [ -Ww H £ H o THE L H- sgusan ass ' _P 3 A + 4 u e L S ..m. 1 H ' 1313 -4 4- S080 p L i Ik EEHidhasast Echi s - i T i 1 . w ] . 0 e s s ¢ugns b l g ! 2 1 1 414 34 “ -+ 40 1 m “v ”. | rL ..Afi lfl .HXfL 4 p 8 f 1 £ T T L Eifaet i g21) ; i SE i | S e e ‘ i , [EaRss & s HHHIH IR 8 e H Bt L st et s R ek 4 t . ] 11 1 4 -+ k ] 5 44 44 a8 - _:. fiflmflm as . L iy a8 ++4+1 +11 r.I. + 4| A Es H = . T ] T : i ! 1T HH agndsen Ht ....m A 1 M fl H I 82 £ i Da l220 380 0 JigR 228" o H H - ~yaktes + - LA} \ A2 [ + Jm [ : T y - 1 1 % 17k. ORNL'LRPM ° .18196 * T TIT T wt._jl:, ASE] XA TE RS IREST TURR PR RN (R 4R : 23 SR ERSEL RRRRERaEas AEARE ARSH HaCSi T ] ft e544 L1141 301 1501 DSEEE FIST] ISOR) LIRS RSP EE: {8 M@ e b 5 RBEE D L UL o +11 “ =0 fos + ~+rm ! - . e RiE R e .“.0“”” ' e ¥ H uma “3 IB3aE Ease e 11 Y ALUTW fiffim 1141 Aw““ T L SRERY B i Sl et 1ed0 B 18 024 EERR Pabs M e s HH T 2 R T et 358 SET EREK B e : T H T 8 254 E2i0d Ehawd LONRE ENRRY KRR ERRLY ERERS RETH ERE N ! e . ks T a3s s o jseed Said Shacd HERESRRRT MES H i L_m g = K 1 1 1] 1L ] * 1 ! 5 perrig ot d HOR B ) R i brd 24 o . ¥ s : R e Hit ] e e H v - : SISO FAAFRANSEe Ses i = . e S ans .t 4 L H o . ! = b 'Mf .m. Jf Pty .w i ik ! 0 @ ! . 1.l = : . 0o BREd o “ 28 : FREL Gl Tt U T kv i 45015451 isTHRERER) tskcARinE B e N0 Nt _ £ : K sspujgpany oe - abaa s sl E=Rot 18! 14 r _ . o — 0 + u SulgE *.. w\w 1 A* wA m + = ; 11 Ht ; A o sasalS3and I2s %u e SO | i H 8 eRESEA LY Pl ¢ 1 Yo = ssags RdEd bul ssmss o e w) - suses gauns g 111t 8 S AORA (i 1 1 ! = m W ) P D TR o ke cea: HHHEH x. L = " & ‘ ”&H;H .MML ..x,rufiru.v leHJH L w ...r 1 AYHL J J 4 : b 1 :AA nm U w “rm m e "flw m : u s #1411«,& “m HAYTW ® r r“.l.‘ 111 - fiv Mj AI.To...,w '@ B %Aw v ." w” u i »w s r. N - —— 9._Im dBtazduii iadasiotel basinnian: SR fsasta: 134 E5ekt A1 SO0 FSE RO R UG i AR +i- | TR S 33 pa3st aaby ! w .Wa.M.. HHETE HHHHH HEHH | ge88s HE2S o : , o T - Y IR Fi 2 Bow CJ _ +HH - TR AR . Lwfl f fi: 1 I By b o B |t Aol .va SSSENERIST ERSRE RRR - all g8 se s HH BTG P - AEHEEN e kBT L . ._- et st M g & [t L nTesetrr is o P “ PR BRI A (o o ,. H Ll lo; ™ 1t -+ lfim ' m w fi hvwv. { “ t | il 0 O TS . ) . ....l.w.L.Lm " L & gagse H flw. HAHHH 1 3RRR25800 i b , | 13 3 ERed LEns 4 L ..A 2 THHEITH L Had] jpdae 1 ! -t P AR o6 S dip R .m @ [if B TH LM H1] seadsasd i ; = ” Seadrioiseesy pery ST R G I | i o Bl RSt LR Ei Sl e e O m 1 = 4 - - n - - 1T ITHE H T =8 EEREYFERS! 12 , H T H . un ngssdud SREAR FER (I8 o .Mu HH 1 m T HHHTTH . E_ 158881 ISEL1 HERS) FRNEIR- . | e lJI 3! R e eE BEE: T vl ] “Hnfl T .Ww w* .”M ; : m sastlass i TR e T 23005500 MRARROREE CEzed dnuzt punni Atl IR - T RS- WSS SRS S o -t . T Wonm. 2 : Lx.fx T. T I RHERE .;‘ & I as % 11 :i H i sge & o + - Il T LY pEs 25 + H A TS0 AN UE RO > aal pid 1 3 igezl m‘ M S HHEHHH ] 1H nh HHHF saeifesaeeast teqaylices sdnddsads! fiq 7 T T ol ¥ 3 .Av... - H l n 17... o T“l“rLH. 1 J11 Alfl I 1 1t IAPr.‘v Aivv va i A $ Afiwl. 1 Jn r‘ - = + m Lfivv = w ERL seeassieatis 1 8 o H &8s gigdpas: R PR T LU - Sy TR ans 25| : 24 B r C Hit 38 b sSSasns ssbau iss 1 5 a8 8 \r ] 111 T 5 H1H 12833 eadh pRekdachd skianiases Rndases Safpsghespease L L : &) s g & g f aed abid ASY 184 S e araics 2o LSy sy (-4 S P Pog BE H1 [ e iead gun . 19 Ae 3 -4 H-11 ] | i 1 T TT Sannit s 1 -1 h4—r aEsa ..%w oL n.wfi THIHTTE o » : H‘ wfi M 1 I o 4 ] L Ll H y FRET - - +1 . ] 4 + 44 T el 1 : a8 sskéa - T - Ht1 8 agw L : : 1t X wTT: aguss = a8 1 m s 83 % ] SEENSEEs 1 t 88 T snagey e fif T R T L L 2 \ » 5 i +1-11 Avlfifl o0 A Tl ly -1 [ - T . SNpNNNERERE RS Ilfix ] M 3 ‘JL L TT 4 = +- gzes ) g T R B R T H H H B R R © v ssatnattibuaclialee ResdindetInishInatsaaqisaseds taseinss + THTH . dusshdedas . u T e 1 ssafupud ffuns seiad Afl H H sessduns CE T HW P S agqEs ‘r..i +4 lhfl.flv o L 111 1T mma fi.w AHA.,‘.).I . 1 . HE e A : sass 28382483, it SEassIups 4 . T 111 1] TH B il SRS TIT] Ifi firdy T B T T R B H T R R D e B T L E 4441411 1 1 1] 4 oL 4444 44 +4 1l T Sy r1 1 T $ii ] vvh vg H ffi. H fi_ 1] urfi.f. 1111 Wfi H H ImUI @ H |- . -t -+ 4 5 ) -4+ 4+ H4+1 4 -t 117 [ ] g & 'l ANBABREE -1 H L : § mAj sds sbepspuans sagas s HHHTT : dedgandatanstininaRis sE39128230458 Hit = H4 J.1. » B @ L4 /B iR i R R R R R R IR R o FH H 11 26688 pREng | aggs 8 H444 o ; H C S wags * e s ,:.Axv 1f1.fi. vI”A Av e 1 A I + 11! $ “ + 1 Um.w .A ;vA.w f_.v TTJ 888 snaus 11 ] ] 1 “ | H» H M Ll _* : W ok 4% i ! l 1 p 3 R R fl.o t ~ .ui. ~w G d s sl R AR A e =T A Nl , i - HEH AT H L 4 1T ERSS ARG U2 BN NEQEPIAREN SERAY BT ..Lh _ " 1? THH [T - 3 1 [ o8 CRRSE R238 - $223 SR RS LR 828! P H 33 °ay/nyg=-uai LVIH ; Wl* LT | | * i 4+ + N /75— TUBE COUNT FOR TUBES SPACED ON EQUILATERAL CENTERS ! (%ube on center) Source: Unpublished derivation by I. Granet, 8/55 L619T="3rq=uT~TNHO *GLT g , H L 1 8 2 5 T | : = 2 "R-Radius to center of outermost tube P-Tube pitch, center %o center of tubes 1 0 O O =% Figure B.l e H- 3 P R | - APPENDIX C REACTOR ANALYSIS CALCULATIONS C.1l INTRODUCTION The solution of the three group, three region (3G3R) problem has been coded for the Oak Ridge National Laboratory's electronic digital computer, ORACLE (Ref. 28). The fast and intermediate energy group microscopic cross section data used to calculate the input data for the code were obtained from the "Eyewash" code (Ref. 27) and from BNL-325 (Ref. 30) by assuming a 1/E flux distribution in these slowing-down groups. The thermal cross section data were obtained from BNL-325 by correcting the tabulated values to the energy corresponding to the average design temperature of the reactor, 607°C. Table C.l is a summary tabulation of the microscopic cross section data used for each energy group. The thermal energy cross sections from BNL-325 were corrected to the average neutron temperature and for a Maxwell-Boltzmann flux distribution as follows: o = 0.885 5 (To)% = 0.48 o, T where, o and T, are the BNL-325 values for microscopic cross section and the standard temperature of 293 A (20°C); and T is the average design temperature, 880°A (607°C). The intermediate energy group was selected to range from .Ql15 ev to 100 ev. This upper limit of the thermal group and lower limit of the intermediate group wvere selected to approximate the energy at which the discontinuity in the neutron flux distribution occurs (Figure C.l). At this energy, the neutron flux changes 177. from a near Maxwell-Boltzmann distfibution to an approximately 1/E distribu- tion. This discontinuity occurs at an energy which is approximately six times - the average thermal neutron energy. ©Small changes in this group enérgy limit' (+ 1 ev) have only a negligible effect on criticality calculation results. Maxwell -Boltzmann e Distribution Neutron Flux 1 Distribution E E | | [ l | I | T Neutron Energy —> Figure C.1 NEUTRON FLUX DISTRIBUTION The upper energy limit of 100 ev for the intermediate energy group waé selected because the uranium-235 resonances cease at about this energy. The upper energy limit of the fast energy group was selected to be 2 mev. The 3G3R code calculation assumes a monoenergetic fission neutron source; and 2 mev is approximately the average energy of the urafiium-235 fission neutron energy spectrum. | Following the notation convehtion of Reference 28, the designation of the fast energy group becomes Group 1, the intermediate energy group becomes Group 2, and the thermal group becomes Group 3. The geometric region designa- tions are: the central core is Region 1; the l-inch thick.baffle between the central core and the heat exchanger is Region 2; and the heat exchanger, other baffles and the downcomer homogenized to be Region 3. The numerical subscripts | [ ' 3 I i |5 174 TABLE C.1l BASIC MICRCSCOPIC CROSS SECTION DATA USED TO CALCULATE 3-GROUP, 3-REGION CODE INPUT Element — _Fast Group _ — o Intermediate Group _ —— Thermal Group (bighly enric§8d) g6_'01 . O—trl 0?1 ral f Tty 30trp 5—;2 6;2 1 3 J—1:1:‘3 0—5‘3 fi3 Fluorine 0.36k4 11.30 .- 0028 0.352 9.85 o 00T 11.7 -- .0051 Sodium 1.03 36.3 -- 0 0.286 9.70 -- .0102 12.0 - 0.258 Zirconium 0.163 22.5 -- 0005 0.135 18.6 o= ,0113 18.6 - 0.092 Uranivm 0.175 36.4 5.31 7.12 0.5 35 34,3 50 30 301 356 Nickel 0.459 36.3 -- 050k 0.592 49.6 - 275 L9 -- 2.35 § Etl 321-.1-1 Zfl & ai § 21:2 3 gtra z 2 b3 a2 3 2“‘3 s 3 233 *Stainless 0.024 1.67T -- 0.005 0.036 2.937 - 0.017 2.89 - 0.1k0 Steel (304) * Note, macroscopic, not microscopic, values for S.5. QLT - - - iEE& on the nuclear property notation refer to the energy group and geometric region, respectively. For -example, 25312 "is the macroscopic absorption cross section for eénergy group 1 apd geometric region 2. C.2 SAMPLE CALCULATIONS Calculation details of macroscopic group constant® for group 1, region 1, for a 0.01 gm/cc uranium concentration are given as sample caleulations below. All the cases for which calculations were made, and the geometric and uranium concentration parameters of each, are given in Table C.2. A summary of all the macroscopic code input data calculated for each case is given in Table C.3. The fused salt fuel composition: 43 mol % ZrF), 57 mol % NaF ~0.1 mol % U-235 The variation of the density of this fused salt composition with tempera- ture is given by the equation, € - 3.65 - 0.00088 x (°c) Then at the average design temperature of 607°C (1125°F), ¢ = 3.12 gg/cc The weight percent of each elemental constituent of the fused salt was calculated to be: Fluorine 45,49 Sodium 13.T% Zirconium 40.9% 180. TABLE C.2 CASE DATA AND COMPUTED MULTIPLICATION CONSTANTS (k) Uranium-235 Core Radius R, Concentration Case No. ft-in. cm - | gm/cm3 Remarks Computed, k "RO CO by - 6" 137.16 0.020 Basic Case 1.34 1 " " 0.010 (Baffle Plate 1.13 : as Region 2) 2 " " 0.005 0.86 3 " " 0.030 1.42 R1 CO 4t - O" 121.92 0.020 1.31 1 " " 0.010 1.09 2 " " 0.005 0.83 3 " " 0.030 , 1.39 R2 CO 3 - 6" 106.68 0.020 1.26 1 " " _ 0.010 1.05 2 " " 0.005 0.79 3 " " 0.030 1.3% R3 CO 5t - o" 152.40 0.020 ' 1.36 1 " " 0.010 . 1.16 2 o " 0.005 ’ 0.88 3 " " 0.030 B 1.44 RL CO 3 -0" 9L.uk 0.020 | 1.21 1 " " 0.010 0.99 2 " " 0.005 0.7T4 ROSCO Lt - 6" 137.16 0.020 Baffle Plate 1.35 as "shell" ROOCO br - 6" 137.16 - 0.020 No Baffle Plate 1.3k 5,68x102cm” 181. TABLE C.3 MACROSCOPIC GROUP CONSTANTS FOR 3-GROUP, 3-REGION.CODE Conc. of U Region 1 Region 2 Region 3 - in gm/cc (Core) (1" SS Baffle) (Heat Exgha.ngerL Group 1 Dy 0.906 cm Dyp 0.50 cm Dy3 0.783 Fast £ixy] 3-04x10-3ecm-1 £y, 2.79x10-3cm1 ax]_g 2°72"1°'3°“_’;-/l S are 2.21x10"%em-l Sai, 4.9x10-3cm™1 S a1 1.29x10 3cm-1 =005 { - gffi 1.68x10-%cm 1 1 IS fig 5.86x1072cm =L a1 3.12x10 " eml Sags 4.9x10-3cm1 b2 1.32x20"3em ™t -010 Z VEfyy 3.35m10 ket 2Tl VE £13 L11x0 en" T ~3em-1 8x10-3cm~1 s }.95x10" *cm k.9x107Fcm = 1.38x10 Jem_ -020 v E ?i’i 6.,"{0x10'1’cm'1 ‘ alz ——- S ?ig 2..35x10"‘*cm s 6.77x10 %em"L £ 4.9x10"3cm™1 < 1.45x10 3cm*1 +030 { "’Zgii 1.00x10"*em1 12 --- -yi?ig 3~52x10"hcm”l Group 2 Dsy 1.42 cm Dnn 0.308 cm . Dpz 0.818 cm Intermediate fixgi 3.80x103en™t £ xgg 7.48x10-3cm ™1 é‘_.xag 3083x10_’3cm’l -3 -1 -3..-1 -3 -1 Sayy 1.06x10 3cm Eapn, 1.86x1073cm Zapz 5-5x1073cm -005 { 1)2;21 1.08x10"3cn ™+ a22 -—- VE 23 3.78x10-%em"1 1.70x1073cm™l £aop 1.86x1073eml £gnn 5.74x10"3cmL -010 .ué ?Si 2.16x10-3cm™ -—— vE, f:3 7o55x10"l‘cm“‘1 2.98x10 3em~1 < 1.86x10-3cm~1 £,apz 6.18x10"3cm"1 .020 {‘U g::i 49322{10‘3(:"['1 322 - _VZ f23 1. 5]-.:(10‘“3cm'°:! - 4.26x10"3en"! S, 1.86x103em™r < .. 6.63x10"3em™* -030 é -.fg‘ ror 6.48x1073em=l T BR T 2 Z«;gg 2.27x10 3cn" Group 3 1331 1.22 cm D3p 0.312 cm D33 0.788 cm Thermal 3 2‘ 1 = 5 1 < 8.41x10"Jem-1 0.155 cm” 4. 76x10"cm” o { V< :gi 9.48x10"3cm~1 "32 --- ’Uzggg 3032fl0-3cm-1 . = 1,30xio--2cm'l Zna, 0.155 cm L < ans 4.92x102cm-1 -010 {—ys :gi 1.90x10~2cm -t 832 - Vs ;gg 6.65x10 3em "1 < 2.21x102em~t &£ ___ 0.155 em~1 £ aas 5.24x10-2cm~1 a020 {'}’E;gi 3079x10-2cm-l ,. 8.32 - Wé-fgg 1033Xl0-2(3m-1 1030 3.13x20"cm~1 2832 0.155 cm~t Zoa33 5,56x10-2cm ™1 .= fgS 1.99x10-2cm 1 182. The negligible quantity of sodium and fluorine in the (NaF)QUFh fuel con- centrate was not included in the calculations. The atomig concentrations in Region 1 are: Npq =z £ (W% of F in salt) Ng = hb9 x 1022 atoms/cc atomic wt of F 1.12 x 10°¢ atoms/cc ] NNa 1 8.42 x 10°! atoms/cc NZr 1 Ny 1 = (uranium conc.) Ny = 2.56 x 1019 stoms/cc @ 0.01 gm/cc 235 The total macroscopic absorption cross section, ' F Na zr U Zau =NF10;1 + Nya 1 Zal ‘”'NZrla;:L +N 1% 3.12 x lO"'h cm 1 ] The total macroscopic fission cross section, multiplied by -/ , the average number of neutrons released per fission (2.46 for U-235), =k fission neutrons/cc @ 0.01 gm/cc U conc. vEg, =¥y (VI7 ") = 3.35 x 10 The diffusion coefficient for Group 1, Region 1j l = l ‘ o Na "y U 3Zir 11 Np 33%aF + Mya 3% - + Ngp 13 % 2Ty 130 0.906 cm The lethargy change in the fast group, ev) = 9.9 units Aul;ln 2x:l.06 100 ev 183. The lethargy change in the intermediate group, Auy = 1n (100 ev) - 5.3 units, 0.5 ev ' | The effective removal cross section (Efx) was approximated as follows: lethargy interval of group AL 'E = average change in lethargy per collision f_-TZs ¢ average no. of scattering collisions/cc-sec Au/f avg no. of scattering collisions for neutrons to pass thru u Approximating 238 by E;t’ the effective removal cross section; or effective cross section, for a neutron to pass entirely through a lethargy interval, fi, becomes, The effective removal cross section for Group 1, Region 1, = Ng 1(§ o—t) F 1t NNa 1(§ o_t)Na 1+ Nzr 3 fd’t) Zr 1"'_1_“‘1_;- '1(3 U'QU 1 Aul | ixl 1 - 3.04 x 1073 emn~? C.3 POWER DENSITY For a central core radius of 3 1/2 feet and a core height of 10 feet, Total volume of the central core - 1.09 x 107 cc Assuming all 600 Mw of heat are generated in the central core, Average power density in the core - 55.2 w/cc 18k, C.4 NEUTRON FLUX The average neutron fluxes in the central core for the hot, clean + eritical condition were computed as follows, ) average rate of fission = 55.2 w/cc x 3.1 x 1010 f’issions/wasec 1.71 x 1022 fissions/cc-sec average rate of fission also equalszzf 11¢11 +2f 21 ¢21 + 2-} 31 ¢31' From Figure 4.3, B, =088, Py =038, The values of 2 ¢ for the hot, clean critical uranium concentration for a 3 1/2-foot radius central core ;, 0.0088 gm/cc, were computed to be, ] b oL g gy = 120 x 10 cm Sy =763 x 107 em”l Seqs 6.80 x 1073 cm~1 Making the appropriate substitutions and solving these equations; one - obtains for the hot, clean critical conditicn of a 3 1/2-foct radius central core, average fast flux in the central core, 100 ev to 2 mev: ¢ll = 1.2k x 1015 neutroms/cc-sec 'avera.’gel intermediate flux in the central core, 0.5 ev to 100 ev: ¢ 2] = 5.96 x 1014 neutrons/cc-sec 185, average thermal flux in the central core, up to 0.5 ev: §j31 = 1.62 x 103* neutrons/cc-sec the percent of the total fissions occurring in each energy group: =+ A x 100 . Se Bx 100 total fissions/cc-sec 1.7l x 101 Fast group -—- % Intermediate group -~=-- 27% Thermal group --- 64% C.5 FUEL BURN-UP Approximately 1.3 grams of-uranium-235lare consumed to produce 1 Mw day of energy. At 600 Mvw power, the uranium burn-up will be appfoximately 780 g/day. This is equiva}ent to burning up a éuantity of urapium approximately equal to the total uranium inventory in the fuel system every year. C.6 EVALUATION OF FISSION PRODUCT POISONING An extended analysis of ihe fission prpduct nuclear poison buildwpp for conditions similar to thosé\ef\phis design estudy has fieen perfofmea (Ref. 33 and 34). This analysis is for a reactor using ufanium-233 as fuel. The fission product yield for U-235 is so nearly equal to that of U-233 that this factor was neglected. Thi§'analysis assumes couplete removal of all.ggeeous fission products, which should very nearly be the condition for the reactor of this ‘design study. The results of this reference are presented as the "poison fraction" build-up with time of operator for the reactorvconsidered° Poison fraction, P = ééz = £(t) . EF The fission product production, and therefore EE:P, is directly proportional to the power density of the reactor fuel system, which in turn is proportional to 186. the product of neutron flux and fission cross section, fo}?: If it is fiasumed that the primary means of fission product removal is by neutron abscrption and only a small portion is removed by radioactive decay, the quantity of fission product poison, Zp, present is approximately inversely proportional to neutron flux,Sd° | | The poison fractiom build-up with time is then seen to be very nearly independent of operating conditions of different néutron flux, fuel concen- tration and power density, P:2p =f(t) (£f¢)(—}3) = £(t) >t E The plot of poisoh fraction increase with time given in Reference 34, page 14, should, therefore, be applicable to this design study's reactor as é first approximation. | The data presented in Reference 34 extends only to 150 days of operation. A visual extrapolation of the curve for total fission product poisoning would indicate that the curve asymtotically approaches a poison fraction of about 15 percent for very long periods of operation (>5 years). This is probably a conservative estimate. Any plating of the metallic fission products in the low flux, heat exchanger region (see Section 2.6.0) would help to reduce this value. Assuming little or no corrosion, and therefore corrosion product nuclear poisoning to be small, the value of 15 percent poison fraction may be used to approximate the negative reactivity introduced in this design study'’'s reactor after very long periods of operation (:>5 years) . 187. Converting poison fraction to units of reactivity (Ref. 8, p. 262), z = Z_-u = 2 (for this design study) Zn reactivity, €= - P z ’-E' - 091'5 x 2 = -0.10 | 1+ z l1+2 From Figure 4.1l of this report, it may be seen that it will require a uranium concentration of approximately 0.0115 g/cc to override'this negative reactivity and maintain criticality after a very long period of operation. This is equivalent to about 131 percent of the hot, clean critical inventory. C.7 REPROCESSING CYCLE TIME The excess uranium fuel inventory that is added to override fission product and corrosion product nuclear poisons should be the economic bptimum that considers the costs of fuel reprocessing. Thé following econofiic analysis is based on the fuel reprocessing method available today (Section 2.,6,0) which requires contaminated salt solvent to be discarded and replaced. Cost of replacement salt = $5 per 1b . Cost of recovering uranium = $1 per 1b from contaminated salt by UF), volatility process (Ref. 38) Total = $6 per 1b Wt of salt in fuel system = $1.85 x 10° 1b - Cost of replacing all of the & e solvent salt in the fuel system = $1.1 x 10 — . 188, . Assuming an interest charge of 4 percent per year on the cost of uranium fuel inventory, Optimum cost of uranium inventory in = $l.1 x 106 = $27.5 x 100 excess of that for the hot, clean 0.04 x Y Y critical condition, so that the accumu-~ lated interest charge on this inventory just equals the cost of fuel reprocess- ing (replacing salt) where, Y = cycle time, in years, to replace all the fused salt solvent in the fuel system. Asguming $20 per gram for highly emriched uranium-235, tkis is, Optimum uranium inventory in excess of = 1.4 x 10° g that for the hot, clean condition . Y Volume of the fuel system = 2.7 x 10( cc Optimum uranium concentration in excess == 0,051 g/cc of hot, clean critical condition after Y ' ! reproces8sing equilibrium is attained Uranium concentration for hot; clean = 0.0088 g/cc ceritical condition Equating the sum of these uranium concentrations to that required to override the fission product poison (Section C.6), The optimum cycle time to replace all = 19 years the solvent salt in the fuel system ' It would initially require a period of time equal to the time of one reprocessing cycle, during which no fuel reprocessing is performed, to attain -this balance of uranium inventory and fission product poisons. APPENDIX D PRIMARY HEAT EXCHANGER DESIGN D,1.0 DERIVATION OF FORMULAE FOR PARAMETRIC STUDY The following basic formulae were used considering sodium coolant on the tube side and fused salt fuel on the shell side. D.1.1 Continuity Equation w - Cav | - (D.1), D.1.2 Flow Area on Shell Side for Triangular Pitch »> 88 1/2 PT x 0.867 Pp - 1/2 rrao2 per 1/2 tube (D.2) R p— . 5 o = B867TP°2 - 7 a per tube / T 3 ° ' 'D.1,3 Equivalent Diameter for Triangular Pitch for Shell and Tube Sides A Shell Side: de = 4 ° flow area = 4 x l/2[(!86722r2 - QEL-doiZ | (D.3) ' wetted parameter 1/2 d, : - vhere, Tqr =d, +s - de = b 6867PT2 - I d02> o Tdg K Tube Side: 2 - . de = 44 | ‘ | (D.&) . 190. D.l.4 . Over-all Heat Transfer Coefficient for Duplex Tube Based on Outside Surface 1 =1 41, /1lnrg + [ry 1n 1y To Ni + l, (D.5) U bg k rijng |k ri/ss To 55 bya [Ti ‘ To l =1 +ro Inrog+ 10 1N T+ T'o Up by kyy ry kgg 11 hyg Ty D.1.5 Film Coefficient between Fluid and Surface Shell side - Dittus-Boelter Equation Nu = h de = 0.023 [deV )0-8[ cp/0.h (D.6) B3 P k - Tube side - Martinelli and Lyon Equation Nu = h dg = 7 40.025 (eig_‘_’_)(LB (31)4?0«:8 | (D.7) & Iz k | D.1.6 Heat Transfer Equation q = Uy S (Atgp) - | . (D.8) D.l1.7 Surface Area Required for Heat Transfer So =Trd, LN ~ (Dp.9) D.1.8 Combining the Above Basic Equations into Working Formulae Then from (D.l) and (D.2) with constants for the conditions of this design, the equations become: D.1.9 Salt Flow . wWg = 65[867 PT? ‘JL;L doajvs x N (D.10) 191, D.1.10 Sodium Flow , - "Na = eNa _71?[_ d3” VNa N (D.11) D.1.11 Salt Heat Transfer Coefficient From (D.6), | hg = 0.023 k (eac Vs)o -8 (Ef)o i , (D.12) 0.4 0.8 . hs-0023k C&)OB( ) v . oy vg® S | then, bg = Cq Vg0 | ¢ 5 C, = 0.023 k_[Pde 0.8 4 A0k ' (D.1k) | de A K | - D.1.12 Sodium Heat Transfer Coefficient from (D.T) hyag = 7 }_c_ 4+ 0.025 L edev 9"8 S.Ef‘o.a | (D.15) de ‘ de /( k . - hNa = C2 + C3 VQ-8 then, =7 k_ and C3 = 0.025 .lf_ Cdg 0.8 32_@0"8 ' : | (D.16) . de de M k , S D.1.13 Over-all Heaf Transfer Coefficient in Terms of : the Number of Tubes From (D.8) and (D.9), q =UoTTdg LN (Atgyg) (D.17) 192, For a éiven Pp and d,, and for the flows required for heat transfer, from (D.10), VS = Ch. N and YNa = 05 N Cy = __¥s . ] where, CS = wNa oo from (D.17), Ut.J N L =z Cg where, C6 = a T, (BEeg) from (D.5), (D.14) and (D.15), = 1 + Ryy + Rgs + 1 Uo cl VSOoa RNyi = ro 1n 1o RSS = _1;9-_ln'rm . C7 +Cg Vyal*O (D.18) (D.19) (D.20) (p.21) (D.22) (D.23) To CB :C3 I"i To Rewriting (D.24), inserting (D.18) and (D.19), 1 Yo Cl c_hOos' c7+08 Eé 0.8 N : N Therefore, U, is a function of N since, 1 = f (N), then also U, = f (N) Uo But from (D.22) and (D.30), then, £f (N) NxL zCg 193. -~ (p.27) (D.28) ~(D.29) (2.30 Now, as L is evaluated at some arbitrary value to agree with the requirements of the problem, f (N) N - Cg L r - (D.31) which is known, then there is only one value of N to satisfy the conditions of the two simultaneous equations (D.22) and (D.29) with L established. These two equations can be solved either by iteration or by graphical means. latter procedure was selected for this study. - The 19k, D.l.14 Calculation of Pressure Drop and Horsepower Due to Friction Sodium pressure drop, Ap = (2.0+ 2L ) VW2 € . (D.32) de 2g 14k combining with (D.19),; ADp = (200 +Clo) 09 (D°33) N2 where, ¢ | | (D.34) 9 > ‘1L % 2g Cig = fL x12 =202 (D.35) de de vhere L - 20 feet, do in inches. Horespower, Pnp =Ap x 144 x VNa | , (D.36) CRa 250 : where, Cyq = b4 x wyg (1b/sec) (D.37) Cyva 5% | Salt pressure drop, combining with (D.20) and (D.32) Ap = [&.o + 013] C1ow -~ (p.39) where, Cip = C° @ - | | | (D.40) 195. Ci3 =fL_x 12'= 2ho t (D.k41) where L = 20 feet. Horsepower, ' where; Cqp = 14h x fis (1v/sec) ' | | (D.43 } g 550 D.2.0 PRIMARY HEAT EXCHANGER SAMPLE CALCULATIONS Solution of the two simultaneous equations discussed in Section D.1.13 results in a heat exchanger with 14,600 5/8-inch 0.D. duplex tubes with an 1/8-inch ligament and .065-inch wall as the optimum and most applicable to this design. However, actual layout of the heat exchanger tube bundle indi- cated that 2L éxchangers with 650 tubes each was a practical éf?aqééfienton Nevertheless, the following sample calculation is based on é heat Efghéfiger of 14,600 tubes with 0.065-inch duplex wall with 0.Ok2-inch Type 30hn??gin= less steel and 0.023-inch Type "L" nickel clad. D.2.1 Properties Fuel Mixture Side Salt No. 3%, 47 mol % NaF and 53 mol % ZrF), Average Properties C ' N 195 1b/ft3, /“= 19 1b/hr-ft, C, = 0.285 BTU/1b-°F, 1.30 BTU/hr-ft2-OF /£t Inlet Temperature = 1200°F Outlet Temperature = 1050°F - Sodium Coolant Side - Average properties © = 51.0 1/£t3, 4= 0.53 1v/hr-ft, C, = 0.30 BTU/1b-°F, K 36,2:BTU/hr-ft2-9F/ft 196. Inlet Temperature - 1000°F Outlet Temperature = 1150°F Inlet Enthalpy = 338.5 BTU/1b Outlet Enthalpy = 293.7 BTU/1b D.2.2 .Flows Sodium 6 Wya = 600,000 x 3413 = 45.7 x 10° 1b/hr = 12.7 x 103 1b/sec 338.5 - 293.7 Salt WS a C, Ebfid wvg = _ 600,000 x 3413 = 48.0 x 10® 1b/br = 13.3 x 103 1b/sec 0.285 (1200-1050) D.2.3. Velocities Sodium ‘ 2 v = VNa 4 =TT d N - No, of tubes Na ETKI?E Ay e i VNa = 45,7 X 106 x b x 14k - 12.8 ft/sec 3600 x 51.0 x (.625 - .13)2 x 14,600 ‘ Salt . VS = Vg Ap = .867 P T2 - Trd°2 C AoN V. - 8.0 x 106 x 14k = 3.7 ft/sec s = - 3600 x 195 x (.867 x .752 - _%;_x .6252) 14,600 197. D.2.4 Equivalent Diameter Sodium N de = 44 = .625 - .13 = 0.495 in. Salt de = 4 (Ay) =4 (867TP2 - /4 a ®) perimeter ‘ T dg de = 4(.867 x .75° - 7T /4 x .6252) - 0.368 1n. - T x .625 - D.2.5 Reynolds Number Sodium i Re = €de VNa /-( Re - 51.0 x 0.495 x 12.8 x 3600 = 183,000 12 x .53 Salt Re - ede VS Re - 195 x .368 x 3.7 x 3600 = 4200 12 x 19 D.2.6 Heat Transfer Coefficients and Wall Resistance Sodium hya de = 7 + 0.025 (Re)-8 (pr).8 ’ X Re x Pr - Pe hya = 36.2 x 12 [7 4 0,025 (183,000) -8 (?30 X ,5€)°8;]; 10,750 36.2 198. Salt hg dg = 0.023 (Re)-8 (pr)-* X h =1.3x12 [0.023 (1:,2oo)°8 285 x 19 T = 1360 - 368 13 ' Wall resistance based on external tube surface * Kgg = 12.9 BTU/hr-rt°-F/rt kg = 335 BIU/hr-ft°F/ft Ryy =T Inrg =z .310 1o .312 z 0.595 x 107 kyy Tp 33-5 x 12 .289 Rys = 1 N 18,700 RSS - I'o ln I'm - 0310 ln 0289 - 3011" X lo-ll. kg ri 12.0 x 12 257 Rgs = _L D.2.7 Over-all Heat Transfer Coefficient Based.on External Tube Surface - -l - , 1 =1 +BRy +Reg+ 1 = 1 4.595x10"% 43,1k x10"+_ 1 (.2h7 U, bs LTS hyg Ty 1360 - 10,750 (.312 ro —— 1 =7.35x 10™* + .595 x 107 + 3.14 x 10°% + 1.18 x 10°% Uo 2 U, = 815 BTU/hr-ft°-F D.2.8 - Effective Length of Tube Required‘for Heat Transfer q = Uy Sg (Btyp) Sg = Mo N L L = q = 600,000 x 3413 x 12. U, Td, N (atgg) 815 x T x .625 x 14,600 x 50 e i = 21 £t P.2.9 Pressure Drop Through Tube Bundle - Sodium - (200 + fL )VNaa . e . T 2 1 1 5.' & o o ' = 183,000, £ = 0.016 (Ref. 16, p. 15). D> d = (2.0 4 £L) vg? de 2SI& _(20+.0h—0x2lx12)372xl95-85psi 368 2g 14 D g D.2.10 _ Pumping Power Required for Heat Exchanger Only Php = aop x 14k = 9.1 x 144 x 12.7 x 103 = 594 hp AP Na %550 51 x 550 o vPhp =aAp x 14l wg = 8.5 x 14l x 13.3 x 103 = 152 hp : . € 550 195 x 550 D.2.11 _Fuel Holdup in Tube Bundle Salt area flow Ayg = [867]? - Tr a 2]n ’ Salt volume Vol = Agg x L, assuming L = 21.0 £t Vol -_-' .867 X .52 - 1T .625] 1k ,600 x 21 ‘ B 1Lk Vol = 385 £t D.2.12 Sizing the Fuel Circulating Pumps Pressure drop in tube bundle 8.5 psi Outer flow periphery 1.0 Check valve 2.8 Contingency 2.7 Total 15.0 psi Total pump power With 8 pumps Assume pump efficiency at 60% Pnp = AD X 1k wg wg = 13.3 x 106 1b/sec 550 % Php = 15 x 14k x 13.3 x 103 = 446 np 195 x 550 x .60 Horsepower each pump - 56 hp Specify 60 hp motor driven pumps 200, hg] e 201. APPENDIX E CALCULATIONS FOR PRELIMINARY DESIGN OF INTERMEDIATE HEAT EXCHANGER E.1 DESIGN PARAMETERS The intermediate heat exchangers are designed to transfer 600 Mw of heat from radioactive sodium in the primary loop to non-radioactive sodium in the intermediate loop. The design 1f based upon the following, previously - selected sodium temperatures and heat exchange rate: ) ~ Primary Na entering “tpy .11509F Primary Na leaving thé 1000°F Intermediate Na enterifig tél' 6?59F Intermediate Na leaving tc2 1100°F Total heat to be transferred --- 2.048 x 109 BTU/hr Heat losses from the reactor vessel, primary ldop piping and immediate heat exchanger have heen ignored. The basic assumption was made to use six intermediateflheat exchangers. E.2 PROPERTIES OF SODIUM AND 304 SS Sodium } Temperature, OF 1150 1100 1000 675 Density, 1b/ft3 49.87 50.30 41.15 53.92 ) Enthalpy, BTU/1b 338.5 323.6 293.7 195.7 Specific Heat, BIU/Ib - 0.2983 0.2983 0.2988 0.3052 Thermal Conductivity, 35.7 36.4 37.7 42,1 . BTU/hr-ft2-F/ft Viscosity, 1b/ft-sec x 103 0.136 0.141 0.152 0.201 304k Stainless Steel Thermal Conductivity 13.2 13.0 12.6 11.3 BTU /hr-ft°-F /£t E.3 SODIUM FLOWS a. Primary loop flow - q = 2.048 x 107 BTU/hr - = 45.7 x lO6 1v/hr OH (336.5 - 293.7) BTU/1b b. Intermediate loop flow =z g = _ 2.048 x 10° BIU/hr _ = 16.0 x 10° 1v/hr AH (323.6 - 195.7) BTU/br E.4 SODIUM VELOCITIES IN HEAT EXCHANGER Preliminary calculations indicated that 1100 1/2-inch 0.D. tubes with 0.042-inch wall appear to be satisfactory. a. Primary sodium velocity (shell side) Assume that the tubes are in a triangular pattern with a 3/4-inch pitch Using Figure B.4 for 1100 tubes, Radius of tube bundle = R = 17.h4 Tube Pitch P R = (17.4)(P) = (17.4)(3/4) = 13.05 in. Shell I.D. = tube bundle dia + tube dia + 2 in. free space f (2)(13.05) + 0.50 + 2.0 28.6 in. (Use 30 in.) Shell cross section area - (O°785)(§(_)_)2 - 4,91 £t° 12 Total tube cross section area = (0°785)£005)2 (1100) = 1.50 £t2 144 Shell side flow area = 4,91 - 1,50 = 3.k1 £t2 6 7.62 x 100 1b/hr Na flow per exchanger = 45.7 x 10 . 6 Average Na density = 49.87 + 51.15 = 50.51 1b/ft3 | 2 (7.62 x 106'lb/hr) Shell side velocity (50.51 1b/£t3)(3.41 £t2)(3600 sec/hr) 12.29 ft/sec ' Y 203 o b. Intermediate sodium velocity (tube side) Flow area per tube = 0.1358 1n® Tube side flow area - (0.1358)(1100) = 1.037 £t° =T Na flow per exchanger = 16.0 x 106 1b/hr = 2.67 x 106 1b/hr . = Average Na density - 50.30 + 53.92 - 52.11 1b/ft3 ) (2.67 x lO6 lb/hg) (52.1 1b/£t3)(1.037 £t2)(3600 sec/hr) Tube side velocity = 13.73 ft/sec E.5 EQUIVALENT DIAMETER FOR SHELL SIDE de = 4 x shell side flow area = (4)(3.41 £t2)(1kk in.®/ft) - 1.137 in. - wetted perimeter (1100)(3.14)(0.5 in.) E.6 REYNOLDS: NUMBER Primary Sodium (shell side) Average Na viscosity =(b,136 +0.152) x 1073 - 0.14k x 1073 1b/ft-sec 2 ./ Re = X#aL = (12.29 ft/sec)(1.137 in.)(50.51 1b/ft3) = 3.75 x 10 Z< (12 in/ft)(0.0001kk 1b/ft-sec) Intermediate Sodium (tube side) Average Na viscosity = £0.141 + 0.201) x 1073 : 0.171 x 10~3 1b/ft-sec N 2 J Re = #d€ =(13.73 ft/sec)(0.416 in.)(52.11 1b/ft3) = 1.45 x 107 )3 (12 in/£t)(0.000171 1b/ft-sec) _ E.7 HEAT TRANSFER COEFFICIENTS Film Coefficient Primary Sodium (shell side) Average specific heat 146.2983 4+ 0.2988\ = 0,2985 BTU/1b-F | \ 2 Average thermal conductivity =(35.7 + 37.7\= 36.7 BTU/hr-ft2-F/ft — 20k, Pr ='&';F a (0.2985 BTU/1b-F)(0.00014k 1b/ft-sec)(3600 sec) = 0.00425 " (36.7 BTU/hr-£t2-F/ft)(hr) Using an empirical correlation for shell side of unbaffled liquid metal . heat exchangers (Ref. 15, p. 285), Nu - 0.212 [(de)(Re)(Pr)]O°6 = 0,212 [(1.,13,7)(30*5 x 105)(o¢,00‘t+25fl°“6 . Nu - Eg_: 19.08 X hy = 19.08 (E_)= (19.08)(36.7 BTU/hr-£t-F)(12 in/ft) - 8402 BTU/hr-ft°-F de (1.137 in.) Film Coefficient Intermediate Sodium (tube side) Average specific heat = 0.2983 + 0.3052 =z 0.3018 BTU/1b-F 2 Average thermal conductivity = 36.4 + 42.1 = 39.3 BTU/hr-ft-F 2 Using Lyon-Martinelli relation for liquid metals and a uniform wall heat flux, Pe - AXCC, = (0.416 1n.)(13.73 ft/sec)(52.11 1b/t3)(0.3018 BTU/1b-F)x3600 soc X (12 in/ft)(39.3 BTU/hr-£t-F) Pe - 685.7 Nu = 7 + 0.025 peQ-8 Nu = by = 7 + (0.025)(685.7)0-8 _ 11.62 _ K h; = 11.62 (k) = (11.62)(39.3 BIU/hr-ft-F)(12 in/ft) - (€)) (0.416 in.) by, = 13,173 BTU/br-ft2-F Tube Wall Thermal Conductivity Average thermal conductivity = 13.2 + 13.2 + 12.6 + 11.3 ]+ . 12.52 BTU/hr-ft-F 205. E.8 OVER-ALL COEFFICIENT OF HEAT TRANSFER do 1 dy +dp.1n3 +1 ' Uo Mdy Kewe Bo | 0.50 - (0.50) + (0.50} 1n EII?TG'F 1 (13,173)(0.Lk16) (12.52)(12) 8402 it 0.0000912 + 0.000612 + 0.000119 0.0008222 U, = 1216 BTU/br-£t°F - E.9 . HEAT TRANSFER SURFACE AREA Logarithmic Mean Temperature Difference At = (th2 - tep) - (tny - tep) In thy - te; (1000 - 675) - (1150 - 1100) 1n /1000 - 675 (1150 - 1109 = 147°F Total Heat Transfer Surface q '.'."Uo So At . Sy =_ g =__(2.048 % 107 BTIU/hr) = 11,457 £t? U, Ot (1216 BTU/hr-£t2-F)(147°F) - Heat Transfer Surface per Heat Exchanger S = 11,b57 _%_._ E.10 EFFECTIVE TUBE LENGTH - 1910 ft° L S - (1910 £t2)(12 in/ft) = 13.27 £t - (N)(Tube perimeter) (1100 tubes)(0.50 in.)(3.1l4) 206. E-11 SODIUM PRESSURE DROP THROUGH HEAT EXCHANGER Pressure Drop for Primary Sodium (shell eide) From Section D.6, the Reynolds number for the shell side is 3.75 x 107 and the corresponding friction factor is f = .01k, From Reference 39, Chapter 1.5, the pressure drop is given by, AP = £ L x Cy2 de 2g AP .(.,011+ x 13,27\::(50,51 x 12.299) 1.137 /] | ek x 12 J NP = 1.61 p{ai Pressure Drop for Intermediate Sodifim (tube side) From Section D.6, the Reynolds number for the tube side is 1.45 x 107 and the corresponding friction factor is £ - .0l7. In this case, the pres- sure drop is given by Equation (5.8), AP = £ £f1L\ x £V? de 2g where 2 1is the number of velocity leads allowed for the tube sheet entrance and exit losses and the 180° bend. AP =/2 4 .017 x 13.27 x 12\ x{52.11 x 130739\ 316 oh.k x 164 J OP - 9.02 psi 207, APPENDIX F STEAM GENERATOR CALCULATIONS . F.l NATURAL CIRCULATION BOILER WITH SEPARATE SUPERHEATER From superheater heat balance, 8boslers = 16,000,000 (127.9 - 42) = 1,372,000,000 BTU/hr Assume two bollers genergting saturated steam-at 2,000 psié, steam on the shell side and sodium on the tube side. Say désign is for 1000°F - and 2000 psi external pressure, the fhickness of a tube for external pres- sure can be calculated (Ref. 35, paragraph UG-31). For 30k sS, "S" allowable = 8800 psi (Ref. 35, Table UHA-23) . X ¥ .19 (min) ".For 3/k in. 0.D. tubes, X min = .143 ' Use .165 average wall (8 BWG) For a Single Wall Separating the Fluids Wall Resistance, ro Inrg = 0.375 1n .750 = .00Lk6 "k ry 12.hx 12 20 .Water Side For present purposes, the assumption will be made that thé unit opefates in the nucleate boiling region and latér on this assumption will bé verified. In order to detérmine a reasonable value of the average boiling and scale coefficients for design, resort will be made to both the-PWR and HRT designs. mRY - From ORNL-LR-Dwg-2795A, there are 251 3/8-1nch 0.D. x .065-in. wall type 347 SS tubes of 20 ft effective length. e 208, Surface (actual outside) = 77 x .375 x 251 x 20 - 494 sq ft 15 On Page (8) of Spec. HRT-1004, the fluid properties, flow rates and temperatures are given, 572 - T 4ok .5 471 ) — Atol = 7705 = 5302 In 101 35 U(Design) = 1.71 x 107 = 652 kol x 53.2 Tube Wall ro Inry = .1875 1n .375 = .00625 k rf 10.6 x 12 245 Primary Fluid hd = .023 ( ) (dV€ Flow area = ( .245)2 x 251 - .0822 sq ft Ny Velocity = 1.79 x 102 = 11.k fps ) 3.6 x 103 x .0822 x 53.2 .339 x .023 x 12 (1 23 x 266\“ ..%2*_5 x 53.4 x 11.4 x 3600\° - 5800 245 .339 / 266 / h (based on outside surface) = 5800 x .245 = 3790 | .375 ' e 1 = .000264 3790 209. The over-all U for the HRT was based on using 3/4 x U clean. R | =.00115 = 1 652 x 4/3 U clean " (wall and primary resistance) = .00890 R | = 000890 = .00026 . U clean h (Boiling) - 1 = 5100; say, boiling coefficient -00026 x 3/ of 5,000 A similar calculation on the PWR design (Ref. 36) indicates that these units héd a degign boiling plus scale coefficient of the order of magnitude - of 2000, This indicates a scale'coeffiéiént of the order of magnitude of 3300 average was used on the steam side of these units. For the present design, an average value of scale plus nucleate boiling coefficient of 2000 is fised. .This should be conservative for design purposes, since the bigher . average temperature differentials in this unit should yield higher bo;ling coefficients than either the PWR or HRT designs. For a natural circulation boiler, the steam water mixture will be external to the tubes and the sodium will be internal to fhe tu.beg° On the basis of preliminary calculations, it appeared that a reasonable size and pressure drop would be realized in two units with approximately 2000 3/k-inch 0.D. tubes to carry in each. i Sodium Side hd = 74 .025 (d_vf cp>'8 *® ~ d = 0.42 ft x 12 K = 40 BTU/hr-£t2-OF /ft Cp = -302 (avg) BTU/1b-°F V - 21.9 fps /<= .178 x 3600 x 1073 - .64 1bs/ft-hr 52,7 lbe/ft3 o < m 11 h =12 x 40| 7 + .025 h2 x 21.9 x 52.7 x 3600 x 30%) A2 Lo h = 15,700 | h based on outside tube surface = 15,700 x .42 = 8800 .15 1 - .000113 00 Flow area = 7T x (.42)° x 2000 = 1.93 sq ft N 1hh V = 8,000,000 = 21.9 fps 3600 x 1.93 x 52.7 1 - (Boiling 4 scale) = .0005 h 1 = 0.002073 U0 Uy = 483 Say, U, for design = 48O 960 675 635.8- Atog = 285 = 135 in 324.2 39.2 ' Surface Area (Boiler) = 686,000,000 = 10,600 sq ft j4»8021{135 211. 377 x 2000 x L effective = 10,600 - L8 L effective = 26.9 ft. Say, total length (including tube sheets) = 28.5 ft Reynolds number = L%g x 3600 x 21.9 x 52.7 :‘229,000 - : qan £ - .015 | | Ap - [2 4 .015 x 28.5 x 1%] x (21.9)% x 52.7 psi _ _ [j 42 4 - 2g 14 AP = 38.6 psi Check of assumption of nucleate boiling, At inlet At - 324.2°F (q/s)inlet - 324.2 x 480 - 156,000 As indicated (Ref. 4 and 6), this heat flux is still in the nucleate boiling region. As a further check, assume the tubes to be. clean. 1 - .00208 L8o 1 - .000303 3300 (1/Uy) clean = .00LTT7 v Uy clean = 564 - | q/S inlet clean - 324.2 x 56u = 182,500 At £ilm = 182,500 a 36.6°F 5000 This is less than the critical At of 45 tc®50 F (Ref. 6, p. 386). There- fore, the assumption of nucleate boiling is Justified. Size of steam generator: R/P : zu.9 (Fig. B.h) Diameter of channel: 20.9 x 2 x 1 + 2.75 = 44.6, say, 47 in. at ID of channel - - 212, A preliminary check of the circulation characteristics of the steam generator indicated that a reasonable number of risers and downcomers can be provided to give adequate circulation for these units. The steam drums A would each be approximately 66 inches ID x 45 feet over-all length to give commercial steam purities with commercial type feed water treatment. The center line of drums would be approximately 30 feet above the center line of the exchangers. Drum supporting steel Qill be provided and the support "~ saddles on the heat exchangers will be designed so that one saddle is fixed - to the foundation and one is let free to move to accomocdate the expansion of the unit and/or the piping. Over-all Heat Balance: Boiler and Superheater Assume boiler at 2000 psia and superheater outlet to be 1000°F and 1800 psia. Total heat load: 600 Mw Q BTU/hr - 600 x 3413 x 1000 H superheated oteam @ 10N0°F 4 1800 psia - 1480.8 BTU/1b H feed @ 450°F + 1800 psia = 430.9 BTU/1b AH - 1049.9 BTU/1b Total 1bs of steam produced = 600 x 3413 x 1060 = 1,950,000 lbs/hr - 1049.9 Temperatures of intermediate loop are as indicated on heat balance diagram Figure 6.2. H sodium @ 1100°F - 323.6 H sodium @ 675°F = 195.7 < Ah - 127.9 BTU/1b 600 x 3413 x 1000 = 16,000,000 1lbs/hr sodium (total) 127.9 AH Sodium/1b = 1,950,000 x 345.7 - 42 BTU/lb 16,000,000 AH superheater 1480.8 1135.1 AH = 3B5.7 A Eya entering superheater - 323.6 - 42.0 H = 281.6 t = 960°F At 1100°F, "s" is 7500 psi Table UHA-23 ASME Unified Pressure Vessel Code - Section 8. The minimum tube wall thickness required is, ‘%t = PRg (PUA-1) = 2000 x .25 = .0602(nomenclature according to ASME) SE + &P 7500 -+ 800 Use 0.065 inch minimum wall tube. If two walls are used to separate the water and sodium, and each wall is required to withstand the full pressure, then the o;xter tube required will be .6660D x .5001D with a wall thickness of 0.083 minimum. | A. Single Wall Separating Both Fluids Agsumed sodium flowing parallel to the tubes and that the unit is of the counter flow type, 1100 o : \ 960 E S t [NOOO | 635.8 At g = 324.2 - 100 = 191°F . . 32%.2 100 ' . Tube resistance = rg ln ry (based on outside tube surface) R 7 K - 12.4 BTU/hr-£t2-°F/ft ' Tube resistance - .25 1n .500 = .00050k4 12.4k x 12 « 370 21k, Preliminary calculations indicated that a reasonable size of super- heater could be achieved by using 1275 tubes to carry per unit. Two such units will be required. : o # Flow area/tube = %E_(°37)2 - .1075 'sq in. Water (steam) side coefficient, . . N o8 a2 zy (. ¥W__ 1000a . Where, W = 1,000,000 1bs/hr per unit b= _ 175 1,000,000 B 175 , (.370)-= 1000 x .1075 x 1275 .82 - o (7.3)°8 - 1046 BTU/br-£12-OF Based on cutside surface; this is, h = 1046 x .370 = TT75 . 500 Resistance = 1 = ,00129 175 Sodium Side The free area/tube = .866 x °752 - T (1/2)2 = .290 E ’ The perimeter/tube = 77 = 1.57 : 2 =4 x Area = 4 x .290 = O.Th .in. perimeter 1.57 - The equivalent dia For the sodium side heat transfer, the Liquid Metals Handbook indicates that the Lyon-Martinelli or some modification is satiéfactdry9i Sincé'the sodium is not the controlling resistance, and for the sake of consistency, the Lyon%fiartinelli correlation will be used. 215, hde = 7 + .025 /dgVP Cp -8 R R . De - 007,4‘ ft - 12 Pavg =20- 3'+ 51.67 = 50.9 lbs/cu ft RE36.U4 'f;_ 38,2 - 37.3 BTU/hr-ft2-OF/ft Cp = +299 BTU/1b-°F Based on main bank geometry, Free area (total) - 1275 x 0.29 = 370 sq in. ) W = f’x AxV Vv = 8,000,000 x 14k - 17 ft/sec 50.9 x 3600 x 370 | h = 37.3 x 12 /; + .025 7h x 17 x 3600 x 50.9 x °2950 i 373 - 8200 BTU/hr-£t2-CF Resistance = 1 = .000122 8200 1 = .00050% 4+ .00129 + .000122 UO Uo = 522 For design purposes, assume U, = 500 Surface required =z 1,950,000 x 345, 7 = 3520 5q ft unit ; 500 x 191 x 2 Active tube length = = 3520 = 21.1 ft ) : 1275 x e Say, approximate tube length (including tube sheets) - 22 1/L ft From tube count layout sheet, - + "R/P = 18.82 Dia to outer tube - 2 x 18.82 x 3/4 = 28.2 216, Approximate dia of channel = 28.2 + .5+ .5+ 1 1/2 = 30.7 For clearance; etc.; say, ID of channel 32 inches. Pressure Drop The following calculations are for the pressure drop in the exchanger. The pressure drop in the inlet and outlet nozzles is not included as part of the pressure drop of the components. This 1s estimated elsewhere. Sodium Side Reynolds No. = DVE = ‘% x 1T x 3600 x 50.9 = 356,000 < St f = .01k for smooth tubes AP (105+,01hx21,1x12\ v2 @ -63v2¢0 .T% ] 2g 14k 2g 1LL 603x%x%gg510psi For bend use 1/2 velocity head, OHP Bed - 0.79 psi say, ll psl for exchanger only Steam Side Assume flow and friction pressure drop are essentially that for am incompressible fluid. If the calculations on this assumption show a high pressure drop, then the variation of specific volume with tube length may have to be considered. The acceleration loss due to the change in specific _ volume between inlet and outlet will be considered. In terms of our nomenclature and the assumed flow conditions, the general pressure drop equaticn becomes, - 2 2 Py -Pp = f L Vayg Cavg + vi) 1 (Vo - V1) + say, 1.5 vel heads, d 2g 1hk4 A) (3600)2g 144 entrance and exit 4+ .5 vel heads bend 217. Cavg = 5»3';—203' - 3.8 v, =.188 v2 g 0M2 - 1,950,000 lbs/hr total =3 e - 010§5 x 1275 = .954% sq ft/unit 1 1,950,000 = 4.8 fps 3600 x .95% x 3.8 Vavg = Reynolds No. = DVC - Qigl x T.8 x 3600 x 3.8 = 48,000 u .068 f = 021 for smooth tubes 021 x 22,3 x 12 x (7 u 822 x 3.8 + 2(74.8)° x 3.8 AP .37 1L% 2g 1LL /1,950,000 e X 2 hh2 - .88 \3600 x 954 32 2 :172(7h8)2x38+(28h) x (k2 - .188) x 1 2g L 1R 32.2 39.4 + 4.4 = 43,8, say, 4k psi Since this represents but a small part of the pressure of the system, it will bé satisfactory to take the pressure drop as being this value. A more accurate calculation, taking into account the variation of specific volume with length, will give a higher pressure drop; but not significantly so. For system design purposes, asaumé a nominal 50 psi pressufe drop on the steam side, not including nozzles. B. Two Walls Separating Both Fluids " The assumption made in this calculation will be the same as for the single wall separating the two fluids. Because of the thicker tube assembly, the 3/4-inch pitch cannot be used. It will be assumed that a 1/k-inch clear 218. ligament will be required for welding of these tubes into tube sheets. It will further be assumed that a thermal bond can be realized between the inner and outer tubes,; but that this bond is not a mechanical bond. Thus, the failure of one fiall will not give rise to & complete failure of the unit. It is further realized that radial temperature gradients; in Some instances; may tend to cause separation of the two tube walls. However, in the superheater, it may be possible to use two different types of stainless steels with \ S e N / 05007 coefficients of expansion in such a ratio as to negate the temperature differentials. Pitch = .666 +-°250 = .916 in. For tube -sheet layout assume ;g in. pitch (.9375) ' 1 Log Mean Temperature Difference = 191°F For Preliminary Purposes, neglect sodium coefficient Assume 1500 tubes to carry .370 {Based on OD of outer tube) -] 510 BTU/hr-£t2-°F .(outer tube OD) < 666 > Steam Side h - 1046 x'(iél;)°8 x 1500 h = Resistance : 0.00196 Wall Inner tube Outer tube = .333 1 Uo .00050k x .666 = .00067 . 500 1n .666 = .0006k 124 x 12 500 1 - .00067 + .00064 + .000196 - .000327. 219, Uy = 306, say, 300. It will be necessary to compute the sodium side afl this point. [866 x_(.9375)° '%;L ( 666)"17& = 0.79 in. 77 X W = Cpx A x V A - 1500 x 413 = 4.3 sq £t 14k V = 8,000,000 - =10.2 fps 50.9 x 3600 x 4.3 hd, = 7+ .025 /a (eVP CID"S h - 0.79 37-3 h = 566 [7 + .025 (1780)] - 6500 1 = .000154 500 1 = .003k2 UO U, = 282 Say for design purposes, U, = Sreqy = 1,950,000 x 345.7 = 6400 sq ft/unit 275 x 191 x 2 ~.37°3X1?f/+( .025) (79x102x3600x509x 299)7 275 BTU/hr-fte-QF (outer tube outside surface) Surface per unit = 1500 x 7T (.666) x L effective = 263 L effective 12 s Leff s 6&00 s 24 ft, 5 in. 260 . Say total tube length = 25 ft, 9 in.. From tube count layout sheet, R/P = 20.3 Dia to L of outer tube - 2 x 20.3 x .9375 = 38;1 in. Approximate channel dia = 38.1 +-2 5 = 40.6 in. For clearance, etc., say ID of channe1==h2 in. Pressure Drop Sodium Side .79 x 10.2 x 356,000 = 235,000 i 17 Reynolds No. = f - 0015 AP = (? $+ .015 x 2k b2 x 12) x (?0,2 2 x 0.9 = 3.8, say, b psi -T9 2g 1 Steam Side Reynolds No. = 63.5 x 48,000 = 40,800 4.8 f - 0021 37 x 2g / | 1k Ap friction + Ap acceleration = 021 (25 .75 x ) ( 63.5)° +2635> x38+ g x (M2 - .188) x 1 1% 32.2 - 32.4 + 3.2 - 35.6 psi F.2 STRAIGHT-THROUGH BOILER 1. Steam FPlow Steam out at 1900 psia, 1000°F, H = 1W77.7 Feed water in at 1950 psia, 450°F, H = 430.1 Change in enthalpy - 1047.6 BTU/1b Steam flow = 600 Mw % 3%13 x 103 BTU/Mw-hr 1047.6 BTU/1b Steam flow = 1.955 x 100 1v/hr 1,950,000 2 .2 x 1275 1k 3600 x 1500 say, 4O psi for desigg 221. . 2. Sodium Flow Na Temp. in = 1100°F Y Na Temp. out = 675°F At average Na temp. (887.5°F), c, = 304 BTU/1b-°F 600 Mv x 3413 x 10° BTU/Mv-hr o Sodium flow = ' -304 BTU x L250F 1b-OF Sodium flow = 15.85 x lO6 1b hr 3. Change in Sodium Temperature through Preheat, Evaporator - and Superheater Zones The change in sodium temperature in a zone is found from a heat talance between heat gained by the water and heat given up by the sodium. (Na flow) cp'At = (steam flow) AH Dtyg = Woag WN& C’P The sodium temperature change calculation is tabulated below: zone = e s preheat 2%6 . « 3090 95 evaporator h73 « 3050 ' :190 superheater | 539 - - 3015 ” lég | 425°F 4, Calculation of Over-all Coefficient of Heat Transfer (U) in Pre- heat, Evaporator and Superheat Zones (Refer to Section 6.5.0) A. Preheat Section a. Since the water entering the boiler is subcooled, fiucleate boiling will occur. A combined coefficient for boiling film . " : and scale of 2000 B’I’U/hr £t2-OF was used (Section 6. 5 1). 222, b. Tube wall coefficient k - 11.2 BTU/hr-£t2-°F/ft for 304 S5 at T23°F ‘h =k = 11.2 - 1790 BTU/hr-£t2-OF ft .25 In .2 12 185 c. Sodium film coefficient k = 41.b BTU/hr-ftach/ft | h zk [7 + (0.025 DV f’cp)°°8] (Ref. T) d X d = 0.0618 ft V = 13.25 ft/sec - 53.6 1b/ft3 0 ' 0.3085 BTU/1b-CF Ca g " h = 41.4 {7+ 0.025 (.0618 x 13.25 x 53.6 x 3085)0"8 5400 BTU/hr-£t2 -OF 1= 1 + 1 4+ _1 = .001453 U 2000 x .185 5K00 1790 - .25 U = 690 BTU/hr-£t°-°F B. Bo@;ing Section a. Use a boiling film and scale coefficient of 2000 BTU/hr-fta-QF ! b. Tube wall coefficient k = 11.8 BTU/hr-£t2-OF /£t for 304 SS at 865°F h=k _ 11,8 = 1890 BTU/hr-rt2-°F £t .25 1n L2 ‘ 12 .185 c. Sodium film coefficient 39.3 BTU/hr-£t2-OF /£t w‘ 1) < " 13.25 f£t/sec oLy o7 .0618 ft AN ) 52.35 1b/ft3 p = 0.3047 BTU /1b -OF h =k |T+0.025 (va’cp)°°é] a k . | - 39.3 |7+ 0,025 {.0618 x 13.25 x 52.35 x. .30k7)"" ,0618 39.3 - 4460 BTU/hr-ft2-COF | = 1 + 1 4+ 1 =-0.001k29 2000 x .185 - 1890 4REO- - 25 U - 700 BTU/hr-ft2-OF C. Superheater Section a. Steam film coefficient hde = 0:023 [dev \O-® cp/4l/3 Kk ' = k| Steam temp. out = 1000°F Steam temp. in - 620°F Average temp. sz 810°F na®-2 . Y(w/1000a)°-8 W =2 x 10° = 625 1b/nr per tube 3200 Y - 165 d - dia, in. = 0.5 - 2 x 0.065 = 0.37 in. A = flow area 1in.2 = 0.1076 in.° - h x (,'37)0°2 ;'165 (625/,0716)0*’8 h = 1100 BTU/hr-£t€-°F -Tube wall coefficient h =k = __ 11.8 =.1890 BTU/br-ft>-°F ft .25 In .25 12 165 ) 223, c. Sodium film coefficient h -k [Tr + 0.025 (dvfcp)°°8] d X d - 0.0618 £t V - 13.56 ft/sec € - 51.0 1w/rt3 k - 37.3 BTU/hr-£t2-OF /£t Cp = 0.3015 BTU/1b/°F h = 37.3 [? +0.025 (,0618 x 13.56 x 51.0 x 3015 x 3600)° .0618 8450 BTU/hr-£t2-OF 37-3 .2 . + .25 1n jigg-f 1 - .003358 37.3 1= 1 U 1100 x :185 025 U - 300 BTU/hr-£t2-OF 5. Calculation of Zone Surface Required 8450 y In a zone, the heat gained by water equals the heat lost by sodium and this heat is trensferred by the equation, q = UAQAt, where &t is the log mean temperature difference. re U 4608 x 10° 690 9252 x 107 700 6624 x 10° 300 6. Number of Tubes A 3,741 £t° 5,981 £t° 11,470 £t° 21,192 f£t° ‘hAs a result of checking pressure drop and steam velocities; a tube size of 1/2-inch OD and nominal 50?f00t length was decided upon. | 225, Total tube area required from (5) above = 21,192 £42 Area per tube,TTDL = 6.54 £t° No. of tubes - 21192 - 3240 tubes 6.5k 7. Pressure Drop in Tubes (Ref. 7, p. 36, l(5, 55, 69)(Section 6.5.2) The fraction of the 50-foot tube length required by the preheat, evaporator and superheat zones is proportionai to the percent of ares of (5) above. | | | 8., Preheat zone 26,800 = 26,800 = 112,000 Re = /A~ 239 £ = .019 Op = FV2 L e 2g de | Ap = .019 x 5.14% x 8.85 x L47.1 6.4 x .0308 | - Ap = 105 1b/£t° - .73 psi b. Evaporstor zone {two phase flow) Re = 26,800 = 26,800 = 132,500 M .202 f - .018 ) AP = -018 x 6.13° x 14,1 x 39.5 bh.bk x .0308 fl APo = 190.4 lb/ft2 (single phase) Appp = APy (APgps) + r (G)° (Ref. 7, p. T2) AP, g Appp = 19C.4 (6) + .2 x (2u2)° 32.2 APpp = 1506 1b/f1:2 z 10.5 psi (two phase) c. Superheater zone (Ref. 7, p. 69) Re - 26,800 = 26,800 = 413,000 /"{ 0065 d. 226, £ - .015 AP = G2 flv +§2 (V2 - Vl) 2g de g _ Ap = (242)2 x .015 x 27.05 x .305 ¢ (242)2 (.4165 - .1950) Bk x L0308 3.0 AP = 4059 1b/ft° - 28.2 psi Total tube pressure drop The total drop is the sum of a, b, and ¢ above. (a) 0.73 (v) 10.50 () 28.20 Total Dfop = 39.43 psi M 227 ° APPENDIX G CALCULATIONS OF STEAM PLANT HEAT BALANCE . The calculations below determine the‘heat balance and gross electric & povwer output of the steam cycle described in Section 6,8°O‘of this report. Thé over-all cycle and the conditions of the working fluids are shown in Figure 6.2, - G.l THROTTLE STEAM FLOW The steam flow at the turbine throttle is calculated from Equation 6.11. Steam flow = 600 Mw x 34:3 x 103 BTU/Mw-hr (1577.7 - 430.1) BTU/1b Steam flow 1.955 x 10° 1 hr . G.2- EXTRACTION STEAM FLOWS No. 1 Heater heat balance 1.955 x 10° 1b (430.1 - 364.2) BTU - Wy (1344 - 430.1) BTU hr 1b 1b Wy = .14 x 10° 1b hr No. 2 Heater heat balance - 1.955 x 106 (364.2 - 300.7) = .14l x 106 (430.1 - 364.2) + W, (1286 - 36h4.2) Wo = 1246 x 108 1b/hr No. 3 Heater mass balance and heat balance 6 s I. 1.955 x 106 = (.11 + .1246) x 10 +w3'+Fow, 6 II. 1.955 x 106 x 300.7 = (.141 4 .1246) x 10~ x 36L.2 + 1222 w3 +238.8 F.W. 228, The solution of Equations I and II above gives, Wy z .089 x 10° 1 hr 6 F.W. = 1.6 x 10” 1b (feed water flow leaving No. 3 Heater) No. 4 Heater heat balance 1.6 x 106 (238.8 - 178.0) = (1164 - 243.9) Wy Wy = .1057 x 10° 1 - hr No. 5 Heater heat balance 1.6 x 100 (178.0 - 117.9) = .1057 x 106 (243.9 - 183.1) + (1095 » 183.1) Ws W5 = .0984 x 100 1b hr No. 6 Heater heat balance 1.6 x 106 (117.9 - 59.7) = (.1057 + .0984) x 106 (183.1 - 122.9) + (1024 - 122.9) Wg WE = 0877 1b hr G.3 HEAT CONVERTED TO WORK IN TURBINE (Equatiop 6.10) Position Steam Flow (1b/hr) OH (BTU/1b) Work (BTU/hr) Throttle 1.9550 x 10° 136.3 266.5 x 106 No. 1 Heater 1.8140 x 10° 58.0 - 105.2 x 108 No. 2 Heater 1.689% x 10° 64 .0 108.1 x 106 No. 3 Heater 1.600% x 106 58.0 92.8 x 108 No. L4 Heater 1.4947 x 10° 69.0 103.1 x 106 No. 5 Heater 1.3963 x 106 71.0 99.1 x 108 No. 6 Hea£er 1;3666 x 108 82.6 107.1 x 106 Total of heat converted to work = 881.9 x 10° BTU/hr ¢ G.4t CORRECTION TO WORK FOR FEED WATER HEATING BY PUMPS a. Correction for Hot Well Pump » The work by the pump done on the feed water was fofind from Equation 6.14. Py is the pump inlet pressure of 1.5 inches Hg or .736 psi. P, is the pump discharge pressure and is the sum of the three items listed below:- No. 3 heater pressure 103.1 psia Elevation difference (cond. to’ 17.3 No. 3 Heater) Friction drop 40,0 - : P, = 160.4 psia The specific volume, Vi, is .0161 ft3. 1v Work = (160.4 - .736) x 1hk x .0161 / 778 Work - 476 BTU/1b The work put into the pump was found from Equation 6.15. Work = 476 - .676 BTU/1b o7 The heat gain of the feed water from the pump was found to be .2 BTU/1b by Equation 6.12. The reduction in steam extracted for No. 6 Heater was found from - -Equation 6.16. _ W = 1.6 x 10 x .2 | - (102h - 122.9) | | ’ W = .000355 x 10° 1b hr The increase in work at the turbine vheel was found from Equation 6.10. Work = .000355 x 106 (1024 - 9hioh) Work = .0293 x 106 BTU hr b. Correction for Boiler Feed Pump This correction was made in the same manner as for the hot well pump above. Work into feed water = (2000 - 103.1) x 14k x .01766 178 Work into feed water = 6.20 BTU/1b Work into pump = 8.86 BTU/1b Heat gain of feed water - 2.66 BTU/1b The reduction in extracted steam is, W = 1.955 x 106.x 2.66 (1286 - 36L.2) W = .00564 x 106 1b/hr The increase in turbine work is, Work = .0056L4 x 106 x {1286 - 9ki.h) - 1.9 x 10° BTU hr Work The turbine work calculated in Section G.3 was corrected as shown below. Uncorrected turbine work (Section G.3) 881.Y x 100 BTU/hr . Hot well pump correction + .03 Boiler feed pump correction + 1.9 Corrected turbine work = 883.83 x 106 BTU/hr This is equivalent to an ocutput of 258.96 Mw. | G.5 CORRECTION FOR TURBINE -GENERATOR LOSSES The losses found by the method of Reference 4l were applied to the corrected turbine wheel output of Section G.4 as shown below. P Corrected output at turbine wheel . 258.96 Mw Turbine exhaust loss , - 8.35 Mw : Mechanical and generator losses Del » : : Total losses 13.5% Gross Electric Power Output 2h§°h2 Mw 231, G.6 CALCULATION OF COOLING WATER PUMP POWER . The cooling watér pump power was found from Equations 6.14 and 6.15. A pressure drop of 10 feet was assumed in the condenser (Ref°’23.)° The cooling water flow was calculated from a heat balénce between the heat rejected to the condenser by the turbine exhaust and No. 6 Heater drain and the heat gained by the cooling water. In calculating the required cooling water.flow, a8 15°F rise in temperature was assumed. This could be greater, or less, depending upon the tefiperature and quantity of cool- ing water available and the type of condenser selected, i.e., single or double pass. | a. Heat rejected to condenser by iufbine éxhauat q = WAHR g = 1.3126 x 106 1b/hr (941.% - 59.7) BTU/1b un q = 1157 x 108 BTU/nr 131 b. Heat rejected to condenser by No. 6 Heater drain .2914 x 106 1b (122.9 - 59.7) BTU/1b hr q q = 18.4 x 106 BTU /hr Co Totalzheat rejected to condenser. The total réjected heat is the 6 sum of Items (a) and (b) above, or 1175.4 x 10° BTU/hr. d. Cooling water flow to condenser W o= 1175.4 x 10® = 78.% x 10° lx15 hr e. Work done by cooling water pump Work (P -P,) VW Work = 10 £t x 62.4 1b x .0161 £t3 x _ BTU _ x 78.4 x 108 1b £t3 1 778 £t/1b hr 232. Work = 1.002 x.10° BTU - .294 Mw hr f. Work put into cooling water pump. From Equation 6.15, Work in = o29h~ Mw - 011-90 Mw Q G.7 POWER REQUIRED BY PUMP MOTORS (steam cycle) The motor power was found by dividing the required pump power by the motor efficiency. a. Cooling water pump motor pover = 490 Mw = .515 Mw 95 b. Hot well (condensate) pump motor BTU 6 1b = Mw-hr 6 pover = 676 T 1.6 x 10 or X 3013 X 10° BTU =95 334 Mw pover ¢. Boiler feed water pump motor power = 8.86 %%H x 1.955 x 106 %% X %!fi%g x 106 BTU .95 power = 5.34 Mw The total motor power required by the above pumps in the steam cycle is 6.19 Mw, I ORNL=LE=Dwg,=18198 1000 IRV ENE N | 1 10 Absolute Pressure- Psis T e T T L e g I O 5 0P 1 O 0 A 0 O O )19 O (O U 0 W Y 0 L of | 4 10 0 T Pressure - Enthalpy Relation Turbine Expansion Line 1 O O O O S U Y O O O O O INEGFEN IEESNREEES EEE AR W NN AN NN . Figure G,1! 1 T ] *qT/n3g -Ldreyjuyg 23k, APPENDIX H SODIUM PIPING PRESSURE DROP CALCULATIONS Sample calculations: (Ref. 7T) H.1 PIPE LINE PRESSURE LOSS Primary loop: 8 in Schedule 40, 304 SS, 15 ft long T = 1150°F € - 49.9 1b/ft3 /A= .00013 1b/ft-sec E/D = .00006 D = 0.665 ft V = 30.5 ft/sec Re pv€ - .665 x 30.5 x 49.9 = 7.78 x 106 A -00013 L = sOkk AP = £ L V2 € psi = .011 x 15 x (30.5)2 x 49.9 = 1.25 psi D 2g 1Lk .665 ‘ 154 H.2 EXPANSION LOSS Hot leg to heat exchanger. Assume one velocity heat loss for expansion. AP = K vl2 x ¥ psi =1.0 x (31+,h)2 x 49.9 = 6.4 psi 2g 1Lk oL. L 1LL H.3 CONTRACTION LOSS Heat exchanger to cold leg. Assume 0.5 velocity head loss for contraction. 16 in Schedule 40, 304 SS pipe. T = 1000°F € 51.2 1b/ft3 1.227 £t 235, V = 145,700,000 1b x 1 br x 1 £t3 x 1 = 33.5 ft/sec - hr 3600 sec 51.216 1.227 x 6 ft°© ' OP = K V12 @ = 0.5 x (33.5)2 x 51.2 = 3.1 psi . 2g 1Lh oh. I 1L H.4 90° BEND LOSS @ Assume 0.1 velocity head loss per 90° bend (Ref. 16). AP =K V2 € psi = 0.1 x (33.5)2 x 51.2 = .6 psi 2g 1LL ' 6L, L N 100°F 51.2 1b/ft3 33.5 ft/sec T e V' K - Ool H=5 VALVE LOSS o Assume 0.6 velocity loss per swing check valvé (Ref. 16). T = 675°F 53.96 1b/£t3 = 26.2 ft/sec € v AP =KV2 € psi £ 0.6 x (26.2)2 x 53.96 = 2.4 psi 2g 1L AT T1hL 236 TABLE H.1l Sodium piping pata ) o VALVES 90~ bends Description No.of lines Length Pipe Schedule Design Number Type Pressure Number Pressure Line Pressure Total Pressure ft. Size Number Velocity Drop Drop Drop Drop ft/sec pai psi psi psi A, PRIMARY LOOCP 1. Primary heat exchanger, 20 | 5/8 in| .065 1n] 12.8 15,7 tube. side wall 2, Feed line (hot)* 24 15 8 ir | 40 30,5 1.5 1,3 2.8 3. Ringheader (hot) 1 75 18 ir 30 26.4 - 8.3 ave . P swing N 4e leader (hot) * 6 30 116 im | 40 344 check 3.8 3,3 1.5 8.6 5. Intermediate heat 14.8 12,2 12,2 exchanger, shell side ‘ swing _ 6, leader (cold) * 6 50 16 ia | 40 35.5 check 3.7 4e3 2.5 10.5 7, Ringheader (cold) 1 75 18 123 30 25,7 8,7 8.7 8, Feed line (cold) 24 9 8 in _ | 40 29,6 1,4 0,7 2.1 7.5 10,5 26.9 68,9 ‘ 9, Pump Pescription: 18,55D gpm., 200 [ft.head,| 176C rpg,, 1000 F, 1335 HP B, INTERMEDIATE LOOP l, Intermediate heat exchanger, tube side 1.8 1,2 i 065 in 15.0 wall 2, Hot leg 6 50 10 in, | 40 27,1 1,6 2.6 4,2 ] 3, Secondary leader (hot) 2 65 18 in | 30 27.7 gate 0.4 1,2 2.8 A J 4, Boiler, shell side 50 . 9.6 11.6 5, Secondary leader (cold) 2 80 18 in | 30 26,2 gate 1,2 2.8 2.9 9.3 swing ' check 2.4 6, Cold leg 6 40 1C in 1 40 25,5 1.5 2,0 3.5 4.0 7.1 10.3 48.0 7. Pump description: 18,350 gpm,, 140 Ft.head,|17€0 rpn., 675 F| 950 HP -, ] APPENDIX I | ' PARTTAL LOAD OPERATION (MATHEMATICAL APPROACH) The assumptions used in deriving the équations of this appendix are explicitly given in Section T.4.1 of this report. In addition to the nomen- clature given at the beginning of these appendices, other symbols are used for the derivations of this section. These arc as follows: - Subscripts - Genera} 1 Full load value of a variable X Partial load value of a variable p - Value of a variable b Value of & variable in the boiler 5 Value of a variable in the superheater f Value of a variable in the fuel loop Subscripts - Temperatures 0 Cold mixed mean temperature of fuel leaving the primary heat exchanger 2 Hot mixed mean temperature of fuel éntering the primary heat exchanger - 3. Inlet sodium temperature to primary heat exchanger L Outlet sodium temperature from primary heat exchanger ) 5 Inlet sodium temperature to intermediate heat exchanger 6 Outlet sodium temperature from intermediate heat exchanger s T Outlet sodium temperature from superheater L Load fraction ratio actual load full load W Pounde of steam generated per hour at full load M Weight flow of sodium in the intermediate loop, 1v/hr zZ Logarithmic mean temperature difference I.1l NATURAL CIRCULATION BOILER WITH SEPARATE SUPERHEATER Since the average fuel temperature and flow rate are comstant, Aty = 150 L | - (I.1) and tp average - 1125°F {1.2) Solving Equations (I.1l) and (I.2) yields the following relations, 1125 - 75 L | . (1.3) Q " " to, = 1125+ 75 L (I,h) As a consequence of the assumption of constant over-all heal transfer. coefficient and constant flow rate of sodium in the primary loop, Equations (1.5), (1.6), (I.7) and (I,B)'result, Zpx = 50 L - (1.5) tyyx - t3x .= 150 L | ‘ (1.6) tox = t3x = 50 L (I.7) toy -~ tyx = 50 L (r.8) The simultaneous solution of Equations (I.6), (I.7) and (I.8) gives the equations for tj, and t3x’ t3x = 1125 - 125 L (1.9) 1125 + 25 L (1.10) Thx The sodium in the intermediate loop transfers its heat to the supexr- heater and the boiller, respectiveiy° At partial load, the over-all transfer rate in the‘superheater will be, 1 = .000710 + 1 . | | (T.11) Uox 775 (L)"8 where the constants in Equation (I.11l) are based on a design value of U, 1 of 500. 20 RC | 1000 SN §t7 . 63508 Therefore, (1.12) st = L x 500 X Zsl .000710 + . 1 . , a(téx - 1000) - (b'-(x - 635»8) 775 (L)-° 1n [ t6x - 1000 trx - 635.8 From the design conditions at full load, The assumptions made for partial load operation include the constancy of the over-all coefficient of heat transfer in the boliler. Based on this and the fact that the surface is constant, Equation (I.13) follows directly. a A 635.8 135 L = by - toy - (1.13) th - 635.8 The last equation needed to specify the conditions in the loops at - partial load is obtained by considering the intermediate heat exchanger. 147 L =(t1+x - t6x) - @x - tey) | (14&) in [tux - b6x G3x - tSX) Substituting the values for t3; and tj,, from Equations (I.9) and (I.10) into (I.14) and simplifying, ylelds Equation (I.15), WTL =150 L - (g - toy) (1.15) In (125 ¥ 258 - 6z 1125 - 125L - te The weight flow of the sodium in the intermediate loop is determined directly by a heat balance, L = M, 6‘6)( - toy) - - (I.16) M) 425 Equations (I.12), (I.13), (I.15) and (I.16) completely determine the steady state fiartial loadfioperation of the system based upon the simplifica- tion and assumptions given in Section 7.4.l. However, a superheater heat balance has nof been utilized in the foregoing. The superheater heat Salance gives an,additionél:independent equation (I.,22)° The ability to write more independent equations than there are unknowns indicates that too many con- straints have been placed on the system by the assumptions made in Section 7.4,1. Therefore, the final steam temperature will not be taken to be con- stant with load and the control of the final steam temperature will be accomplished by using an attemperator. Equation (I.12) has to be adjusted to reflect the change in this assumption, and instead of 1000°F, the'final steam temperature should be indicated to be variable. In this case, Equation (I.22) has to be written in terms §f the enthalpy of the steam side corre- sponding to the unknown final steam temperature at the superheater ofitlet. It is further noted thaf for a variable final steam temperature leaving the superheater, Equations (I.11) and (I.12) are not correct as written. A v N 241, 'pumerical solution is necessary where the final steam temperature is assumed for a given load. From a heat balance of the steam loop, thé weight of steam is obtained and on is calculated. The procedure is then to solve the appli- cable equations simultanecusly, and finally to check the assumed final steam temperature for the load under consideration. A procedure such as outlined above is tedious and time consuming. 1.2 ONCE-THROUGH STEAM GENERATOR The basic difference in the mathematical approach to the problem of partial load operation of the once-through boiler as compared to the natural circulation boiler with separate superheater is that the sum of the surfaces of the boiler and superheater sections is a constant, but each can vary indivi- .Gually with load. Equations (I.3), (I.4)}, (I.15) and (I.16) are applicable to this case, However, certain of the other relation§ were modified, and, where necessary, new equations were developedol In general, the loéd fraction (L) can be written for a given component as, L = Uy Z, Sy | | (1.17) Uy 21 5, For the boiler and preheater section, Equation (I.18) is obtained by combining Equations (I.17) and (I.13), Spx = L x 135 x SBpy _ln*(t"{x - 635°9 ' (I,iS) \th - 63508 ' (‘b"{x - th) ’ The same procedure follows for the superheat section, where Equation (I.17) is combined with Equation (I.12). However, the design values for the once- through boiler are different than for the separate superheater and the con- stants in Equation (I.12) were adjusted accordingly. The design value for 242, the steam side coefficient was 1100 and the over-all coefficient was 300 BTU/hr-£t°-OF. - | (1.19) 1100 (L) LxU; x191 [.002k2 + 1 %fisl = (t6x - 2000) - (b7x - 635.8) Sex. 1n tgx - 1000 ‘t7x - 635.8 Solving for S, (1.20) s L x 300 3 242 1 t6x - 1000 gx = x 3 x 191 [00 _ + 1100 (L).B (Ssl) in t7x _ 635.8 (t6x - 1000) - (7x - 635.8) The sum of the surfaces at any load must be a constant. This sum at design load is 21,192 square feet vhere, Seqy = 11,470 sq £t and sbl' = 9722 8q ft Therefore, . 1 tgyx - 1000 L x 300 x 191 (.oozhz + = 8) 11,470 1n ) 1100 (L)* t7y - 635.8 (t6x - 1000 ) - 6;7,( - 635.8) | 4+ L x 135 x 9722 ln(th - 635-8) = 21,192 (1.21) . 'bsx - 635.8 - . A heat balance of the superheater ylelds Equation (I.22), ~ My = 1ko L (I.22) fiI t6x - t?;)' Equations (I.22) and (I.16) are then combined to yleld a single equation with the flow rates eliminated, 66;( - tox)e %2% (tsx - tr(; | | (1.23) .1“":‘ B 2k3. The solution of Equations (I.15), (I.21) and (I.23) simultaneously gives the desired steady state partiaL load operation of the system for the once-through boiler. . It is noted that the final steam temperaffire can be set at 1000°F and the system is cbmpletelyldete;minate'fiith only inter- mediate loop sodium flow rate control; this is not the case for the natural circulation boiler with separéte superheater. ok, APPENDIX J SIMULATION FOR 600 MW FUSED SALT REACTOR AND STATIONARY POWER PLANT by E. R. Mann Instrumentation and Controls Division Oak Ridge National Laboratory Figures (7.%), (J.1), (J.2) and (J.3) are the elementary flow sheet for the power plant and the road maps required for simulation, with the excéption of that for the reactor kinetics which is described in Ref. 5k. On Figure (J.1l) is the road map for the fused salt loop and the first or primary sodium loop. Amplifiers 1, 2, 3 and 4 provide a unity gain cir- cuit for the fused salt loop. Amplifiers 7, 8, 9 and 10 provide a similar clrcuit for the primary sodium coolant. Amplifiers 5 and 6 generate the coupling between these two flulds as the difference between the mid tem- peratures of the two fluids in the heat exchanger. The time constant of the metal in the heat exchanger was not included in these simulation studies. Accordingly, about amplifier 5 are shown a potentiometer setting "o£" and series resistor "R" to indicate that such a time constant should be included. | . This time constant can be calculated rather simply as follows. The heat capacity of the fused salt in the heat exchangér is given by the expression, Ce = P ST/St where P 1s the power given up by the fused salt in passing through the heat exchanger at design point steady state, and ST is the temperature drop of the salt during the time St 1s in the heat. exchanger. P 1s in 5 2 ’\ 245, BTU/sec, OT in degrees F, and St in seconds, and Cs is in BTU/F. Now the mass of metal in the heat exchanger times its specific heat gives the heat capacity Cyg g, of the heat exchanger. If then, [g¢ 1is the time constant of the fluid in the heat exchanger, which here has beeéen approximated by two first 6rder lags whose sum is the fluid tfansit time, then the time constant of the heat exchanger is, In transient, this circuit with the time constant 7:H,E.’ provided by amplifier 5, gives the heat transfer delay between fluids due to heating | or cooling the metal of the heat exchanger. | The road map on Figure (J.2) shows, through amplifiers 13 and 14, the intermediate sodium in the sodium-to-sodium heat exchanger. Amplifiers 15, 16, 19 and 20 show the sodium "going and returning” in a heat exchanger between the hot and cold legs of the sodium piping. The flow sheet, Fig- ure (7.4),shows how valves V1 and Vg will provide means tor returning part of the cold sodium through this auxiliary heat exchanger which drops the temperature of the sodium entering the boiler. This auxiliary heat exchanger is one fieans available to liwmit thg steam temperature on partial'loads to a maxifium temperature of 1000°F. It was selected for simulation here primarily because the analog facility re- quired less equipment io do thisrthan it would have required to vafy the sec&ndary sodium flow rate. | The road map, Figure (J.2), shows amplifiers 17 and 18 as the sodium in the sodium-to-wa£ép andyvapor heat exchanger. It was assumed that the load "L" was proportional to water flow, i.e,, the water entered the boiler at constant temperature for all loads. 246, Road wmaps, Figure (J.3), show by amplifiers‘ll and 12 how the driving function "4 " between the two sodium systems in the sodium-fo-sodium heat exchanger was generated, and by amplifiers 21 through 26, inclusive, how the regulator action tb keep the steam fiefiperature constant with varying loads "i" was obtained. The quentities ".&" and "R", with amplifier 11, are determined by the time constant of the metal in the heat eichanger for the sodium-to-water heat transfer. The heat capacity of this heat exchanger was not used ifi these simulations. -This time constant could have beén determined by the same procedure proposed for that of the metal in the sodium-to-sodium heat exchanger. The output of amplifier 27, Figure (J.3), at design point should be +.2,5 V. This output was recorded on a Brown recorder with a full scale reading of 5 V. The output gafie a midscale reading. Two limit switches about the midscale reading actuated a mofior‘to-turn.a ten-turn potentiometer "E", Figure (J.3). The motor was reversible and: its-direction was determined by the limit switch contacted b&‘deviation of the output of amplif;er‘ET;,.If the deviation was positive and caused the recorder pen to move upscale into the upfier limit switch, which would be true for decreasing load, then the motor turned the potentiometer setting up and thereby "valved" more sodium through the‘auxiliary heat éxcha.nger° Except for the relatively slow motor speed, or ?ate of motion of fhé glide wire on the potentiometer due to the géar reduction avallable, this device could hold the steam temperatfife con=- stant for large variations in power. The curves show deviations in steam temperatures durifig transients; but these‘comé from tfie relatively slow response times of the actuator mecha.nism° 247, Scale factors in the simulation studies were as follows: 1. Computer time was real time . -2, One volt represented 10 megawatts of power 3. One volt represented 20°F The circuits were not set up from a set of differential equations representing the system. The set of equations can be derived from the circuits by setting the currents at the amplifier inputs equal to zero. Then by converting electrical units to power and temperature units using the scale factors given above, one gets the equations of the system. Some of these equations will be dependent. +A,+2.5V +T,456.25V +A,+25Vv +2.5V 66181~ "8m-dT1-"INHO “gh2 ¢ + T, +44.37V +8,+9.375V M . M IM I G- +Tg & b ' I'M -C R ‘FIG.—J.Z- SIMULATOR CIRCUIT FOR INTERMEDIATE SODIUM LODP 00281- "Emg-¥T1-"INYO * 642 +B +9.375V ‘ ‘ ‘ -9.375V FIG. J.3-SIMULATOR CIRCUIT FOR DRIVING FUNCTION IN INT.ERMEDIATE HEAT EXCHANGER AND REGULATOR FOR STEAM TEMPERATURE CONTROL 10281~ " SmI-d1-TINYHO *062 J,\ 251. APPENDIX K TEST OF DOUBLE TUBE SHEET DESIGN FOR WATER SODIUM ISOLATION An experiment has been conducted to determine the feasibility of using brazed joints in the fabrication of a double header for the heat exchanger(s). Procedure Approximately 400 holes were drilled in a 1/2-inch thick type 347 SS plate to receive short lengths of type 304 stainless steel tubes of nominal 1/2-inch diameter, .035-inch wall. The holes were reamed to allow .005-inch . clearance on the diameter to permit easy assembly and satisfactory design for brazing. Although none of the currently available elevated temperature alloys have been found to be entirely satisfactory in elevated tempergture sodium corrosion tests}/, the alloy Cast Metals No. 52 (84 Ni-4 Si-2B) was selected as representative of the alloy system most iikely to yleld a éorro- sion resistant alloy for extended service. This alloy system, however, possesses a fundamental.disadv&ntage, in that Boron diffusion into the stain- less steel grain boundaries results in a serious loss of dfictility. The experiment was conducted; therefore; to merely verify the feasibility follow- ing development of a nickel-base corrosion resistant alloy. The alléy V&8 applied as & precast brazing ring on one face of the tfibe shee? afound eachn tube and fixed in place with a CM 52 pounder slurry. The tuhe;éheet was heated in -lOOoF'dew point hydrogen to 1050°C at a rate of lOSOOC/hour, held 1/2 hour at 1050°C and cooled at 100°C/hour. This cycle was selected as representatifie of commercial practice. 1/ CF-56-4-130, "Sodium Corrosion and Oxidation Resistance of High Temperature Brazing Alloys", G. M. Slaughter et al. 252. Results The results of thie experiment are illustrated in Figures (K.l) and (K.2), photographs of the underside of the tube sheet after brazing. Four tubes were found to be only partially brazed, the most serious‘flaw being shown enlarged in Figure (K.2). In view of these results and in view of the research still required to develop a brazing alloy suitable for extended service in sodium, the method of fabrication suggested by this study is at best marginal and not recommended. —— \ 253 Underside of Brazed Tube Sheet (Arrow points to brazing flow) -1 Figure K 254 ONE INC Figure K-2 Defective Brazing of Tube (Close-up view of tube indicated by arrow in K-1) A f(N) 255. SYMBOLS USED IN THE ENGINEERING CALCULATIONS Cross sectional area for flow of fluid ft2 Ai inside or tube side A0 outside or shell side Constant Specific heat at constant pressure, BTU/1b-F Tube diameter, ft di inslide diameter d outside diameter ‘ 4~ diemeter of interface of cladding and base metal d equivalent diameter; 4 x flow area/wetted perimeter Function of N | Friction factor, dimensionless Acceleratién due to gravity, 32.2 ft/sec?l Heat transfer coefficient, BTU/hr-ft°-OF hNa for sodium hS for salt Enthalpy, BTU/1b Thermal conductivity, BTU/hr-ft>-OF Geometrical length, ft Number of tubes, dimensionless Nuééeit number, dimensionless Pressure, psi Power delivered to the fluid, horsepower Peclet number, Pe - Re Pr, dimensionless Tube pitch, ft Prandtl number, dimensionless Rate of heat transfer, BTU/hr 256. 2 , R Resistance to heat flow, hr-ft~-°F/BTU Re Reynolds number, dimensionless r Tube radius, ft . S Area of heat transfer surface, ft2 S{ referred to inside surface of tubes So referred to ocutside surface of tubes K s Ligament between adjacent tubes; ft t Temperature, °F T Absolute temperature, %k o Uo Over-all heat transfer coefficient based on outside surface of tubes, BTU/br-ft2-OF - v Velocity of fluid, ft/sec v Specific volume, ft5/1b | Vol Volume of holdup, ft3 \Y Mass flow rate, 1b/hr - x Thickness, ft Greek AP Pressure difference, psi At Temperature différence, % Aty Temperature difference between bulk mean temperatures of two fluids , Atyy Logarithmic mean temperature difference - 72 Pump efficiency /7 Avsolute viscosity of ‘fluld, 1b/hr-ft " ¢ Density of fluid, 1b/ft> i; Notation signifying a summation, dimensionless ny, | 77 3.1416 Subscripts - . S ¢ Na Refers to sodium S Refers to the salt 257. SYMBOLS USED IN THE NUCLEAR PHYSICS CALCULATIONS Atomic weight of an element . Geometric buckling of the reactor, cm~2 Constant Diffusion coefficient, cm Multiplication constant of the reactor " Atomic concentration, atoms/cc No Avogadros number, 6.02 x 1023 atoms/gm atomic wt Poison fraction,CZ{p 2{;, dimensionless Radius of reactor's central core, cm or ft Time, seconds Temperature; oC‘or éF Lethargy, logarithmic energy decrement Time, years Greek § 7 ¢ o z Average change in lethargy of a neutron per collision Average humber of neutrons released per.figs;on Nefitron flux, neutrons/cm®-gec¢ Density, gm/cc Microscopic cross section, ém? or barhs Macroscopic cross section, em~t 258. BIBLIOGRAPHY Ref. No. 1. Shannon, R. H., APDA, Private Communication, June 20, 1956 2. Ledinegg, M., "Flow Instability in Natural and Forced Circulation"; "Die Warme", 1938 3. Leib, E. F., Discussion to paper by Van Brunt, Trans. ASME, May 1941 b, Jens, W. H., "What is Known about Boiling Heat Transfer", Metropolitan Section of ASME, April 7T, 1954 5. Morabito, J. J. and Shannon, R. 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