IBRARIES [l ORNL-1721 This document consists of 107 pages. Copy 7,,,2,::4 231 copies. Series A. Contract No. W<7405-eng-26 AIRCRAFT REACTOR ENGINEERING DIVISION ORNL AIRCRAFT NUCLEAR POWER PLANT DESIGNS A. P. Fraas W . Savolainen A, May 1954 DATE ISSUED ROV 10 1854 OAK RIDGE NATIONAL LABORATORY Operated by CARBIDE AND CARBON CHEMICALS COMPANY A Division of Union Carbide and Carben Corporation Post Office Box P Oak Ridge, Tennessee AR 3 U4sk 03L0LS9Y B om‘qpr.n;hww_—- >-TMERZOCSOSFPEIPCINIEOPOTDOIP wihmggemmP A41vE . Abbatiello . Batch . Boyd . Bussard . Cardwell . Center . Charpie . Clewett , Cottrell . Cowen J. Cromer UmIPrPpMEIZTMIZ P> . L. Culler . B. Emlet . P, Fraas C. Gray R. Grimes . Hollaender ., S. Householder . Jordan . Keim . Kelley Lane . Larson . L.a¥erne . Livingston . Manly . McQuilken . Meem . Miller 74, 75. 76. 78. 79. 80-91. 92, 93. 94.98. 100. 101, ORNL-1721 Special INTERNAL DISTRIBUTION 20. K. Z. Morgan 31. E. J. Murphy 32. J. P. Murray 33. W. G. Piper 34, H. F. Poppendiek 35. P. M, Reyling 36. H. W. Savage 37. A. W, Savolainen 38. R.D. Schultheiss 39. E. D. Shipley 40. M. J. Skinner 41. A, H. Snell 42, R. 1. Strough 43, J. A. Swartout 44, E. H, Taylor 45, D. B. Trauger 46, J. B. Trice 47. F. C. Vonderl.age 48, A. M. Weinberg 49, G. C. Williams 50. C, E. Winters 51-60. 61, 62-67. 68. 69, 70. 71. 72-73. X<10 Document Reference Library (Y-12) Biology Library Laboratory Records Department Laboratory Records, ORNL R.C. Health Physics Library Metallurgy Library Reactor Experimental Engineering Library Centra! Research Library EXTERNAL DISTRIBUTION Air Force Engineering Office, Oak Ridge Air Force Plant Representative, Burbank Air Force Plant Representative, Seattle . Air Force Plant Representative, Wood-Ridge American Machine and Foundry Company ANP Project Office, Fort Worth Argonne National Laboratory (1 copy to Kermit Anderson) Armed Forces Special Weapons Project (Sandia) Armed Forces Special Weapons Project, Washington (Gertrude Camp) Atomic Energy Commission, Washington (Lt. Col. T, A, Redfield) . Babcock and Wilcox Company Battelle Memorial institute Bendix Aviation Corporation iii 102. 103-105. 106. 107, 108. 109. 110. 111. 112, 113, 114. 115-119. 120. 121, 122, 123-125. 126. 127-134. 135. 136, 137. 138-139. 140-143. 144.145, 146-147. 148. 149, 150. 151. 152155, 156. 157-158. 159. 160. 161-162. 163-165. 166, 167-173. 174-175. 176-185. 186-187. 188. 189. 190. 191, 192. 193-194. 195-196. 197. 198.203, 204-215. 216-230Q. 231, Boeing Airplane Company Brookhaven National Laboratory Bureau of Aeronautics (Grant) Bureau of Ships Chicago Patent Group Chief of Naval Research Commonwealth Edison Company Convair, San Diego (C. H, Helms) _ Curtiss-Wright Corporation, Wright Aeronautical Division (K. Campbell) Department of the Navy - Op-362 Detroit Edison Company duPont Company, Augusta duPont Company, Wilmington Duquesne Light Company Foster Wheeler Corporation General Electric Company, ANPD General Electric Company, APS General Electric Company, Richland Glenn L. Martin Company (T. F. Nagey) Hanford Operations Office lowa State College Kirtland Air Force Base Knolls Atomic Power Laboratory Lockland Area Office Los Alamos Scientific Laboratory Materials Laboratory (WADC) (Col. P. L, Hill) Nuclear Metals, Inc. Monsanto Chemical Company Mound L.aboratory National Advisory Committee for Aeronautics, Cleveland (A, Silverstein) National Advisory Committee for Aeronautics, Washington Naval Research l_aboratory Newpert News Shipbuilding and Dry Dock Company New York Operations Office North American Aviation, Inc. Nuclear Developiment Associates, Inc. Patent Branch, Washington Phillips Petroleum Company (NRTS) Powerplant Laboratory (WADC) (A, M. Nelson) Pratt and Whitney Aircraft Division (Fox Project) Rand Corporation {1 copy te V. G. Henning) San Francisco Field Office Sylvania Electric Products, Inc, Tennessee Yalley Authority (Dean) USAF Headquarters U. S. Naval Radiological Defense Laboratory University of California Radiation Laboratory, Berkeley. University of California Radiation Laboratory, Livermore Walter Kidde Nuclear Laboratories, Inc. Westinghouse Electric Corporation Wright Air Development Center (WCSNS, Col, John R, Hood, Jr.) Technical Information Service, Oak Ridge Division of Research and Medicine, AEC, ORO FOREWORD Formal Air Force interest in nuclear propulsion for aircraft dates from October 1944, when the head of the Power Plant Loboratory (WPAFB), Col. D. J. Keirn, upproached Dr. Vannevar Bush on the subject. Subsequent to that and other discussions, the NEPA group was formed in 1946, The NEPA group moved to Qak Ridge in 1947, and by 1948, ORNL had begun to provide assistance in research and testing. The ORNL effort gradu- ally expanded, and the ORNL-ANP General Design Group was formed in the spring of 1950 to help guide the program and to evaluate and make use of the information being obtained. Four years of work at ORNL on the design of aircraft nuclear power plants have dis- closed much of interest. In a project so complex and so varied it is inevitable that many of these points should escape the attention of nearly all but those immediately concerned or be forgotten in the welter of information produced., Some of this material is buried in ANP quarterly reports, and much has never been formally reported. ' Many reactor designs have been prepdared, but each design has represented an isclated design study, and the issues have been much confused by variations in the assumptions made in the course of each reactor design. This report is intended to provide a critical evaluation of the more promising reactors on the basis of ¢ common, reasonable set of design conditions and assumptions, CONTENTS FOREWORD & v evvvveeenennnenss, e BT v SUMMARY i it ittt i e et n e st s e s e s e s e e e e | PART I, DESIGN CONSIDERATIONS MILITARY REQUIREMENT S & i it ittt it s tr e s ssn o s e oo neosanononsasos 2 PROPULSION SYSTEM CHARACTERISTICS & . i vttt i i i it s aastannsnssnsssansss . 2 Vapor-Cycle Compressor-Jet . o v v i i i st it onavsonaesnnonesnsesensssessss 3 Gas-Cycle Compressor-Jet . . i v it o it ittt et it in oo onsetnonesenosoasonssans 3 T U O e v vt v it it i e e e e e e e e 3 Specific Thrust and Specific Heat Consumption .o v v it it iv v vunas b s es e 4 Chemical Fuel as a Supplementary Heat Source ... ..... S e e e e ae i e 4 REACTOR TYPES .. ittt ittt et st osannsnssnan C et e e et 4 AIRCRAFT PERFORMANCE ............ ettt e e e e e e e 6 FfHects of Reactor Design on Aircraft Gross Weight ... .. o i ittt i it it enennsn 6 Effect of Chemical Fuel Augmentation .............. f et et e e n et e s H SHIELDING ..ttt i i i ittt i isnenn e e n et e e e e e e e e e 16 Units of Radiation Dose Measurements . v o v v vt vt t ot nnnsansnnssesssonssonssnsas 16 Permissible Dose Rate for Crew . . . o it i it it it it it i s st s ot s asmo s nonononon s 17 Radiation Damage to Organic Materials and Activation of Structure . .o o4 L. e 17 Ground-Handling and Maintenance Problems . .o v v v v v s i i ittt s s i e i8 Shield Weight « v v it v it et et it e e s st e ottt saaneaears 21 NUCLEAR PROPERTIES & it it ettt sttt asassnssetsnsastossoenssasesess 33 Moderating and Reflecting Materials + . v o v v i it i i i it i i e 35 Effect of Moderating Material onDesign . . s i i vt i it o s ettt nesnnsas 36 Reflector-Moderated Reactor .. v v v i v ot o s v ot v s et s sososvososnsnassossosasssas 38 REACTOR CONTROL 4t v ittt i it i et snasssoaasananssanssosstoisanosnarssss 39 MATERIALS o v v vt vt et es v snonnncsosn et e e e e e e e ey 40 SHUCTURE & « 4+ s e o s o s v v s o o oo s o e o ssosooeonnsstnsasasanorsansassseesnsseass 41 Solid Fuel Elements o v v v e s o v ot o v o v am o s o s 0 s o s asosonusnsesncensnsosscnsaas 42 High-Temperature Liquid Coolants and Fuel Carriers . .. ..o v v i i i oo iin s, 48 HEAT REMOVAL i et it e e et e e 51 TEMPERATURE GRADIENTS AND THERMAL STRESSES . . o vt it i s i e ittt e e 55 TEMPERATURE DISTRIBUTION IN CIRCULATING-FUEL REACTORS ..o vt v it it iien s 59 PART Il. REACTOR STUDIES COMPARISON OF REACTOR AND CYCLE TYPES ... .o it ittt et i oo 63 REACTOR, HEAT EXCHANGER, AND SHIELD ARRANGEMENTS . .. . ittt enneenn 64 Shield and Heat Exchanger Designs . . . o v v i i it i vt i ien s i s s i et o ancsan 67 Reactor Core Configurations . v v v v v vt s e v s n v s s et sasaccsnsnsas S e e s 76 Vil DETAILED DESIGNS OF REACTORS --------------- R a8 4 8 & B B F s 8 6 a s s SN e oo 79 Sodium-Cooled Solid-Fuel-Element Reactor . .. v ii it i i it in it it et te v in e ensennn 79 Circulating-Fuel Aircraft Reactor Experiment . ., .,..... S et s e et et e e e 83 Fluid-Moderated Circulating-Fuel Reactor .. i i ittt i ittt ittt it ettt ensanans 83 Reflector-Moderated Circulating-Fuel Reactor .. ., i i e i i i ii it i ii it i it e eanaans 87 SECONDARY FLUID SYSTEM 1. i ittt ittt ettt it ettt onssansnnsaatonens 93 MAJOR DEVELOPMENT PROBLEMS L. it it it ittt it e sttt aosttanansens 96 vili ORNL AIRCRAFT NUCLEAR POWER PLANT DESIGNS A, P, Fraas A, W. Savolainen SUMMARY The detailed design of an aircraft nuclear power plant poses an extraordinarily ditficult set of problems. V*2:3 |t will be found implicit in this report that the problems are so intimately inter- related that no one problem can be considered independently of the others; yet each problem is sufficiently complex in itself to be confusing. In an effort to correlate the work that has been done, a tentative sef of military requirements for nuclear- powered aircraft is presented first and accepted as axiomatic, . The types of propulsion system that might be used are discussed next, and the turbojet engine is shown to be the most promising. Aircraft performance considerations are then presented on the basis of a representative power plant, and the shield data used are validated in a section on shielding. It is shown in these sections that the reactor should be capable of a power density in the reactor core of at least 1 kw/em® and, prefer- ably, 5 kw/ecm®, and it should operate at a suf- ficiently high temperature to provide a turbine ajr inlet temperature of at least 1140°F for the turbojet engines. The effects of nuclear considerations M The Lexington Project, Nuclear-Powered Fiight, LEXP-1 (Sept. 30, 1948). fi 2R\epor’r of the Technical Advisory Board to the Tech- nical Committee of the ANP Program, ANP-52 (Aug. 4, 1950). 37. A. Sims, Final Status Report of the Fairchild NEPA Project, NEPA-1830 (no date). on the size, shape, and composition of the reactor core are presented, and in the light of the preceding presentation, possible combinations of materials and the limitations on the materials are discussed. The effects of the physical properties of several representative coolants on the maximum power density obtainable from a given solid-fuel-element structure is determined on the basis of a consistent set of assumptions. Design limitations imposed by temperature - distribution and thermal stress are also examined. | From the data presented in the section on air- craft performance and in the sections on nuclear materials and heat removal considerations, it is shown that the reactor types having the most promising development potential and the greatest adaptability to meet the wide variety of military requirements are those in which a liquid removes heat from the reactor core at temperatures of 1500°F or higher. Designs fer several high-temper- ature reactors are presented, and their advantages and disadvantages are discussed. The problems involved are too complex to permit anything approaching an Aristotelian proof to support a choice of reactor type, but it is hoped that this report will convey something more than an oppreciation for the various decisions and compromises that led first to the circulating- Hluoride-fuel reactor and then to the design of the reflector-moderated reactor type recently chosen as the main line of development at ORNL. ' PART I. DESIGN CONSIDERATIONS MILITARY REGQUIREMENTS The potential applications of nuclear-powered gircraft to the several types of Air Force mission are quite varied. Robot aircraft, ram-jet and rocket missiles, and unmanned large nuclear-powered tugs towing small manned craft have been suggested as a means of avoiding the shielding problem in- volved in the use of nuclear power. As will be shown later, it is probable that even in missiles some shielding would be required because of difficulties thot would otherwise arise from radi- ation damage and rodiation heating.? Furthermore, from the information available, it appears that these applications, while possibly important, either would not justify the large development expense of the nuclear power plant required or, in the case of nuclear rockets, would represent such an ex- trapolation of existing experience as to be very long-range projects, A number of different missions for monned aircraft with shielded reactors are, however, of such crucial importance as to more than justify the development cost of the nuclear power plant, All these missions involve strategic bombing, Studies by Air Force contractors have indicated that the aircraft should be capable of operation (1) at sea level and o speed of approxi- mately Mach 0.9, or (2} at 45,000 ft at Mach 1.5, or (3) at 65,000 ft at about Mach 0.9. A plane of vnlimited range that could fly any one or, even better, two or three of these missions promises to be extremely valuable if available by 1965, |In addition to the strategic-bombing application, there are important requirements for lower speed (Mach 0.5 to 0.6), manned aircraft, such as radar picket ships and patrol bombers. The problems associ- ated with supplying a beach head a substantial distance from the nearest advance base indicate that a logistics-carrier airplane of unlimited range would also be of considerable value. In re-examining these requirements, it is seen that a nuclear power plant of sufficiently high performance to satisfy the most difficult of the design conditions, namely, manned aircraft flight at Mach 1.5 and 45,000 ft, would be able to take care of any of the other requirements, except those involving rocket missiles., Because of the rapid 4R, W. Bussard, Reactor Sci. Technol., TID-2011, 79- 170 (1953). rate of advance of ceronautical technology and because of the inherently long period of time re- quired to develop a novel power plant of such exceptional performance, it appears that develop- menta! efforts should, if at all possible, be centered on a power plant of sufficiently promising develop- mental potential fo meet the design condition of Mach 1.5 at 45,000 ft either with or without the use of chemical fuel for thrust augmentation under take-off and high-speed flight conditions. [t has heen on this premise that work at ORNL has proceeded since the summer of 1950, PROPULSION SYSTEM CHARACTERISTICS Several types of propulsion system well svited for use with manned aircraft are adaptable to the use of nuclear power as a heat source for the thermodynamic cycle on which they operate. One of these is the turbopropeller system in which a steam or gas turbine is employed to drive a con- ventional aircraft propeller, with heat being added to the thermodynamic cycle between the compressor and the turbine. A second is the compressor-jet system, a binary cycle in which a steam or gas turbine is employed to drive a low-pressure-ratio air compressor. The air from the compressor is heated in the condenser or cooler by the turbine working fluid and then expanded through a nozzle to produce thrust. A third system, the turbojet, employs a gas-turbine cycle. Inthis system enough energy is removed from the air passing through the The balance of the expansion of the air is allowed to take place turbine to drive the compressor. through o nozzle to produce a relatively large thrust per pound of air handled. A fourth system, the ram-jet, will work well only at flight speeds above Mach 2.0, because it depends upon the ram effect of the air entering the engine air inlet duct; the ram effect provides the compression portion of the thermodynamic cycle. Heot is added after compression and the air is allowed to expand through a jet nozzle to produce thrust. Because it eliminates the relatively heavy and complicated parts associaoted with the compressor and turbine, the ram-jet system appears, on the surface, to be much the simplest mechanically, but in practice, serious complications arise because any given unit will work well only in the very narrow range of flight speeds for which it was designed. |t should be nofed that each of these four systems operates on a thermodynamic cycle that involves an adiabatic compression, followed by addition of heat at constant pressure, and then an adicbatic expansion. Qf the four types of propulsion system cited, only the compressor-jet and the turbojet look prom- ising for the applications envisioned. The turbo- propeller system is handicapped by the poor aero- dynomic performance of propellers obove high subsonic speeds and by the very serious problems associated with the high blade stresses inherent in such designs. The ram-jet power plant is use- fess for take-off and landing ond is so sensitive to speed and altitude that it does not look prom- ising for manned aircraft. Vapor-Cycle Compressor-Jet The wide-spread use of vapor cycles has di- rected attention to water as a working fluid for the thermodynomic cycle of an aircraft power plant. The principal difficulty associated with such a power plant is the size, weight, and drog as- sociated with the condenser. in attempting to establish the proportions of such o power plant, it soon became evident that only by going te high temperatures and pressures and by using the cycle in conjunction with a compressor-jet engine to give a binary cycle could o reasonably promising set of performance characteristics be obtained.® By superimposing the water-vapor cycle on a com- pres sor-jet cycle, the power generated in the steam turbine could be used to drive the air compressor, while the condenser that would serve as the heat dump for the steam cycle could also serve to heat the air of the compressor-jet cycle. With this arrangement, the air pressure drop across the condenser could be kept from imposing an intoler- able drag penalty on the airplane. Vapor cycles essentially similar to the water-vapor cycle have been proposed which use mercury,® sodium,” or rubidium as the working fluid. These fluids make possible much lower operating pres- SA. P. Fraas and G. Cohen, Basic Performomce Char- acteristics of the Steam Turbine-Compressor-Jdet Ajreraft Propulsion Cycle, ORNL-1255 {(May 14, 1952). 6A. Dean and 5. Naokazate, Invesfigotion of a Mercury Yapor Power Plant for Nuclear Propulsion of Aircraft, NAA-SR-110 (Mar. 21, 1951). Ty, Schwartz, [nvestigation of a Sodium Vopor Com pressor Jet for Nuclear Propulsion of Ajrcraft, NAA-SR- 134 {(June 25, 1953). sures than could be used with water af any particu- lar temperature level. Unfortunately, the weight of the mercury required per unit of power output for the mercury-vapor system appears to be too high,® while the sodium-vapor system must be operated at a temperature well above that feasible for iron-chrome-nickel alloys.” Gas-Cycle Compressoar- et A somewhat similar system has aolso been con- sidered which would use helium as the working fluid with a closed-cycle gas turbine.® Helium could be compressed, passed through the raactor, expanded through o turbine, directed through a heat exchonger to reject its heat to the gir stream of the compressor jet, ond returned to the helium The exiro power obtained from the helium turbine, over und above that required to COMPIressor.. drive the helium compressor, would be employed to drive the air compressor of the compressorsjet cycle. This system would have the advantage of using helium to cool the reactor and thus would avoid any form of corrosion of materials in the reactor. ' Turbajet Several cycles thot use air us the thermodynamic working fluid have been proposed. The first of these would employ the redctor to heat the air directly by diverting it from the compressor through the reactor before directing it to the turbine of the turbojet engine.® With this arrangement the only large heat exchanger in the system would be the reactor core, because, with an open cycle, no bulky condenser or cooler would be required. A versatile variant of the turbojet system. is based on o high-temperature liquid-cooled reactor that could serve as the heat source for not enly a turbojet but for any of the other propulsion systems menticned, that is, turbopropeller, compressor-jet, or ram-jet. Versatility would be obtained by com- pletely separating the air that would serve as the working fluid of the thermodynamic eycle from the reactor and by using a good heat transfer fluid to corry the heat from the reactor to a heat exchanger placed ot a convenient position in the propulsion system. While heat exchangers would be required with systems of this type, they could be kept B4 Schwartz, An Anolysis of Ineri Gas Cooled Re- actors for Applicotion to Supersonic Nuclear Aircrofs, NAA-SR-111 (Sept. 8, 1952). relatively small because they would operate ot a high temperature with superior heat mediums. transfer Specific Thrust and Specific Heat Consumption In evaluating the merits of any particular pro- pulsion system, it is convenient to work in ferms of specific thrust and specific heat consumption because the size and the weight of the power plant depend on these two parameters, The higher the specific thrust in pounds per pound of air handled, and the lower the specific heat consumption in Btu per pound of thrust, the smaller and lighter the power plant will be, The most important factor that offects these two parameters is the peak temperature of the working fluid in the thermo- dynamic cycle.?:? in the binary cycles, such as the supercritical-water and helium cycles, the peak temperature in the air portion of the cycle is also a very important factor. A comprehensive pres- entation of the effect of temperature on specific thrust and heat consumption can be found in the report of the Technical Advisory Board,? which shows that the specific thrust is dependent mainly on the peak temperature of the thermodynamic cycle, irrespective of whether a compressor-jet or a turbojet is employed. This is a very important conclusion, since it indicates that compressor-jets and turbojets give substantially the same perform- ance for the same design conditions, except insofar as the weight and drag of the machinery required is concerned. Chemical Fuel as a Supplementary Heat Source The use of chemical fuel as a supplementary heat source has important implications. The foremost among these is that the chemical fuel could be used to sustain flight in the event of a nondestructive reactor failure. Another very im- portant application would be the use of chemical fuel for warmup and check-out work when operation of the reactor would present radiation hazards to ground personnel. Yet another important possi- bility would be the use of chemical fuel for inter- burning to raise the air temperature just ahead of the turbine in the turbojet engine or for afterburning following the turbine. Either arrangement could be used to obtain increases in thrust of as much as %A. P. Fraas, Effects of Major Parameters on the Il:’ée;—f{c;rmnce of Turbojet Engines, ANP-57 (Jan. 24, 51). 100% with little increase in the weight of the machinery required. Such arrangements would be most attractive to meet take-off and landing or high-speed requirements. The use of interburning or afterburning would not be practical with the vapor or helium cycles because the low pressure ratio of a compressor-jet engine makes it inherently insensitive to the addition of extra heat from a chemical-fuel burner. Similarly, the large pressure drop through the direct-air-cycle reactor would make the air cycle less responsive to the addition of heat from a chemical-fuel burner than a high- temperature-liquid turbojet system would be. While separate engines operating on chemical fuel only might be employed, a lighter power plant and a lower drog installation should be obtainable by the addition of burner equipment to the nuclear engines. REACTOR TYPES Each of the various types of propulsion system described in the previous section could be coupled to one or more of a wide variety of reactor types. The most promising of the reactor types can be classified, as in Table 1, on the basis of the form of the fuel, the manner in which the moderator is introduced, and the type of fluid passing through the reactor core. The materials considered for each design are also given in Table 1, together with the type of propulsion system to which the design is best adapted. References to the studies of these reactor types are given. The only reactor types for which studies have not been made have been the boiling homogeneous reactor and the stationary-fuel-element liquid-fuel reactor cooled by either a boiling liquid or a gas. Studies were not made of these types because, at present, there are no known combinations of materials that would give good performance in these reactors. Many factors influence the selection of a reactor type because many different requirements must be satisfied. The various limitations imposed on the reactor design by aircraft requirements, nuclear and heat transfer considerations, materials prob- lems, etc., are discussed in the following sections. The information brought out in this way is then applied to a critical examination of detailed de- signs for reactors representative of the more promising types. TABLE 1. AIRCRAFT REACTOR TYPES FLUHD FLOWING PREFERRED TYPE FFERE E REACTOR TYPE FORM OF MODERATOR THROUGH REACTOR | OF PROPULSION SYSTEM REFERENCES Stationary Solid fuel (sintered UO2 and Circulating H,O Supercriticalewater— 10 fuel stainless steel in a stainless compressorsjet steel-clad compact, graphite- NaOH High-temperature liquid— 11 UOZ' SiC-UOz, cermets) turbojet Stationary Ligquid coolant (Liz High-temperature liquid— 1,2,12 {Be, BeC, C, Be, C) Na, Pb, Bi, fused turbojet : fluorides) Boiling coolant Sedium=vdapor—compressor- 7 {Na) jet Gas coclant Helium, gas turbine, 8 (air, helium) compressor-jet Direct-air-cycle turbojet 1,2,3 Liguid fuel (static fluorides Circulating NaOH High~temperoture ligquid—~ Not reported in tubes) turbojet Stationary Liguid coolant High«temperoture liquid— 13,14 (Be, BeO, C, BezC) {Na, Pb, Bi) turbojet Boiling coolant Not studied Gos coolant Not studied Circulating Homogeneous {fuel dispersed or Boiling Neot studied fuel dissolved in liquid moderator) Nonboiling NaOH-U02 slurry High-temperature liquid— 15 turbojet Li7GH-N00H°U02 High-temperature liquid—~ 15,16 solution turbojet Separate moderator Solid (Be, BeO, ) Fused fluorides Highstemperature liquid— 17,18 turbojet U-Bi High-temperature liquid— 19 turbojet Liguid {HZO, MNaOH, Fused fluorides High-temperature liquid— 17,18 NaOD, Li’0D) turbojet AIRCRAFT PERFORMANCE Quite a number of different approaches have been made to the problem of determining the feasibility of nuclear aircraft. Most of the NEPA studies were devoted to fairly detailed designs for a few particular aircraft to meet certain specified con- ditions. Both the Lexington Committee and the Technical Advisory Board did some parametric survey work, but, because of the limited time and information available, there were many questions left unanswered. North American Aviation, Inc., followed the same general approach as that used by the Technical Advisory Boord, but again, be- cause of the limited information available, their survey was The Boeing Airplane Company has done a fair amount of parameiric survey work, but the bulk of that published has been devoted to the supercritical-water cycle, The design gross weight of an airplane is a good indication of its feasibility partly because a high incomplete, gross weight with a low payload indicates a marginal aircraft, and partly because it is doubtful whether o craft of more than 500,000-1b gross weight would be tactically useful if it could carry only a small payload. Further, the costs of con- struction, operation, and maintenance of aircraft are directly proportional to gross weight. Any difficulty that required for its solution a small increase in component weight over the value assumed for design purposes would require a large compensatory increase in gross weight, Therefore it is important to know the effects on aircraft gross weight of the key reactor design conditions, ]ONuclear Development Associates, Inc., The Super~ critical Water Reactor, ORNL-1177 (Feb. 1, 1952). ”K. Cohen, Circulating Moderator-Coolant Reactor for Subsonic Aircroft, HKF-112 (Aug. 29, 1951). 120 B, Eilis (ed.), Preliminary Feasibility Report for the ARE Experiment, Y-F5-15 (Aug. 1950). IsR. W, Schroeder, ANP Quar. Prog. Rep. Mar, 10, 1951, ANP-60, p. 28. YR, C. Briant et al,, ANP Quar. Prog. Rep. Dec. 10, 1950, ORNL-919, p. 22! 15¢, Cohen, Momogeneous Reactor for Subsonic Aire craft, HKF-109 (Dec. 15, 1950). Yw. B. Cotirell and C. B. Mills, Regarding Homogene- ous Aircraft Reactors, Y-F26-29 (Jan., 29, 1952). 7y, B, Cottrell, Reactor Program of the Aircroft Nuclear Prapulsion Project, ORNL-1234 (June 2, 1952). 18A. P. Fraas, C. B. Mills, and A. D, Callihan, ANP Quar. Prog. Rep. Mar. 10, 1953, ORNL-1515, p. 41, WK. Cohen, Circulating Fue! Reactor for Subsonic Adrcraft, HKF-111 (June 1, 1951). namely, temperature, powser density, and radiation doses inside and outside the crew compartment. Effects of Reactor Design on Aircralt Gross Weight A parametric survey29 of airplane gross weight was carried out by using the quite complete set of shield-wesight data prepared in the course of the 1953 Summer Shielding Session2! and the turbojet- engine performance and weight data given in a recent Wright Aeronautical Corporation report.22 The shield-weight charts are reprinted here as Figs. 11 to 15 in the section on ‘‘Shielding.” These charts constitute the only consistent set of shield-weight data availoble for a wide range of reactor powers and degrees of shield division. The degree of shield division is a function of the location of the shield material. The more divided the shield, the heavier is the crew shield and the lighter the reactor shield. The shield-weight data are for shields made up primarily of layers of lead and water. The reactor shields of Figs. 11 to 15 were ‘‘engineered’’ for reflector-moderated circu- lating-fuel reactors to include weight allowances for reactor, heat exchanger, pressure shell, struc- ture, headers, ducts, and pumps. As will be shown in the latter part of the section on ‘‘Shielding,”’ the total shield weights given are representative for most reactor types, except air- or gas-cooled reactors, for which the large veids infroduced in the shields by ducts and heoders would cause major increcses in shield weight in comparison with the values given. The Wright data for turbojet- engine weight are representative of the propulsion machinery weight required for the most promising types of propulsion system. A set of tables was prepared from the reactor design dato to focilitate solution of the basic equation for aircraft gross weight. Studies have shown that aver-cll power plant performance is not too sensitive to either the compressor pressure ratio or the pressure drop from the compressor to the turbine provided the pressure drop does not 20A. P. Fraas and B. M. Wilnher, Effecis of Aircroft Reactor Design Conditions on Aircraft Gross Weight, ORNL CF.54-2-185 {May 21, 1954), 212 b, Blizard and H. Goldstein (eds.), Report of the 1953 Summer Shielding Session, ORNL-1575 (June 14, 1954). 22R. A. Looas, H. Reese, Jr,, and W, C. Sturtevant, Nuclear Propulsion System Design Anclysis Incorpo- rating @ Circulating Fue! Reacter, WAD-1800, Parts | and 11 {(Jan, 1954). exceed 10% of the obsolute pressure at the com- pressor outlet.?® Hence the engine compression ratic was taken as &:1 aond the pressure drop from the compressor to the turbine was taken as 10% of the absolute pressure at the compressor outlet, with one-half of this considered as chargeable to the radiators. The turbojet-engine data were taken largely from the Wright report. The specific thrust and the specific heat consumption were taken from Figs. 1X-1 through 1X-12,22 the engine, compressor, - and turbine weight were taken from Fig. 1-19, and the engine air flow from Fig. {18, Engine nacelle drag was taken from Fig. 67 of ANP-57,° except that 50% submergence of the nacelles in the fuselage was assumed, The weight of the engine tailpipe, cowling, and support structure was taken as 25% of the compressor and turbine weight. The total weight of the NaK pumps, lines, and pump- drive equipment was calculated from the estimafes given in ORNL-1515"8 1o be 38 Ib/Mw. The radi- ator cores were designed to give a turbine air inlet temperature of 1140°F with a 1500°F peak NaK tem- perature and an air pressure drop across the radi- ator core equal to 5% of the compressor outlet pressure, The radiator size and specific weight were determined by extrapolation of the experi- mental curves in ORNL-150923 for a tube-and-fin core employing 15 nickel fins per inch. These data were combined with the turbojet-engine data to obtain the propulsion machinery weight, and then the installed weight of the propulsion machinery and the reactor power output as functions of thrust for various flight conditions were determined. The results of these calculations are presented in Tables 2 and 3. The basic equation used to relate aircraft gross weight to the weight of the aircraft structure, the useful load, the shield weight, and the weight of the propulsion machinery was the same as that used by the Technical Advisory Board, North American Aviation, and Boeing: Wg = Wst + UL + Wsh + me , where Wg = gross weight, Ib, W_, = structural weight (including landing gear), Ib, UL = useful load, Ib, 23w, S, Farmer ef al., Preliminary Design and Per- furmance) of Sodium=to-Air Radiators, ORNL-1509 {Aug. 26, 1953). W., = shield weight (reactor shield ond crew shield), Ib, Wom = propulsion machinery weight, Ib, The weight of the structure was taken as 30% of the gross weight. While the value would probably be closer to 25% for subsonic aircraft (except for gircraft using power plants with low specific thrust, such as the supercritical-water cycle), the value used seemed representative ond adequate for the purposes of this analysis, The solution for aircraft gross weight was ob- tained graphically by preparing charts such as Fig. 1. The weight of the propulsion machinery plus reactor and shield that could be carried by an airplane after providing for structural weight and useful load was plotted agoinst gross weight to give o family of steeply sloping paraliel straight fines., The weight of the shield and the propulsion machinery required for each of a series of gross weights was then plotted on the same coordinates, the aircraft gross weight being taken os the product of the thrust and the lift-drag ratio. The solution for the gross weight is defined by the intersection of the curve for the total power plant weight re- quired with the line defining the power plant weight that could be carried with a porticular vseful load, The liftedrag (L/D) ratio estimated for each flight design condition would not be the optimum lift-drag ratio obtainable with the airplane becouse take-off, landing, ond climb requirements would necessitate wing loadings lower than those for minimum drag. The L/D values used are given in Table 4, These L/D ratios are for the airplane configuration without nacelles, an allowance for nacelle drag having been deducted from the specific thrust given in Table 2. Thus the L/D ratio with nocelles would be lower than that indicated, par- ticularly at high Mach numbers. The useful load was considered as including the crew, radar equipment, armoment, bomb load, and other such items. Since the shield weights used were for a dose rate of 1 ¢/hr in the crew compart- ment, the useful lood can alsc be construed to include any extra crew shielding required to reduce the crew dose to less than 1 r/hr, For the purposes of the study, a useful load of 30,000 1b was selected as typical. The aircraft gross weights obtained were then plotted against dose rate at 50 ft from the center of the reacter (af locations cther than in line with TABLE 2, CALCULATIONS FOR POWER PLANT SPECIFIC QUTPUT Compressor Pressure Ratio — 6:1 Ratic of Radiator OQutlet Pressure to Inlet Pressure = 0,90 (3600°\d h ‘ . d b c d e { g=f — |- h i=— i R=1i+] \3413/e ¢ ] Specific e . . . Propulsion Turbine Specific Turbajet Engine NaK System Mach Altitude Inlet Specific Thrusi Less Heat Consumption Installed Weight Weight Machinery Thrust Nacelle Weight No. (1) Temperature (Ib-sec/Ib) Drag Btu/sec-lb kw /b Ib-sec/Ib ib /b (“;‘/]b of (1b/1b of CF) (Ib-sec/1b) of thrust of thrust of air of thrust thrust) thrust) 0.6 Sea level 1140 25.7 25.2 6.22 6.69 15.25 0.605 0.494 1.099 1240 30.7 30.2 6.07 6.5 15.0 0.496 0.481 0.977 1340 35.5 35.0 6.04 6.46 14.81 0.424 0.478 0.902 0.4 35,000 1140 40.8 40.3 5.35 571 56.6 1.405 0.532 1.937 1240 45 44,5 5.54 5.91 55.9 1.255 0.550 1.805 1340 48.3 47.8 5.6 5.97 55.1 1.154 0.555 1.709 0.9 Sea level 1140 19.6 18.6 6.86 7.62 12,05 0.648 0.549 ¥.197 1240 24.8 23.8 6.73 7.40 11.89 0.500 0.533 1.033 1349 26.3 28.3 6.55 7.15 11.73 c.414 0.515 0.929 1540 37.5 36.5 6.35 6.88 11.40 0,313 0.496 0.809 0.9 35,6000 1140 35.8 34.8 5.63 6.1 43.5 1.250 0.555 1.805 12490 49 39 5.75 6.22 43,0 1.103 0.5¢5 1.668 13490 43,5 42,5 5.8 5.26 42.4 0.998 0.570 1.568 1540 50.8 49.8 5.85 6.29 41.4 0.831 (.572 1.403 1.5 35,000 1140 24,5 20.0 6.30 8.14 24,6 1.231 0.635 1.866 1240 28.5 24.0 6.26 7.34 24.3 1.010 0.612 1.622 1340 33 28.5 6,20 7.57 24.0 0.843 0.5¢90 1.433 1540 40.5 36.0 6.16 7.32 23,6 0.656 0.571% 1.227 1.5 45,000 1140 24.5 20.0 6.30 g.14 28,6 1,979 0.748 2.727 1240 28.5 24.0 6.26 7.84 39.0 1.625 0,720 2.345 1340 33 28.5 6.20 7.57 38.6 1.353 0.696 2.049 1540 40.5 36.0 6,15 7.32 38.0 1.055 0.673 1.728 *j = £{0.038 NaK piumbing weight + specific radiator weight). TABLE 3. PROPULSION MACHINERY WEIGHT AND REACTOR OUTPUT FOR VARIOUS THRUST REQUIREMENTS Ratio of Radiator Qutiet Pressure to Inlet Pressure = 0,90 Compressor Pressure Ratio = 6:1 TURBINE THRUST (ib) MACH | ALTITUDE INLET 10,000 15,000 20,000 25,000 30,000 40,000 50,000 60,000 NOG. {ft) TEMPERATURE o W o » prx 1 W r W P W P W P W F W P W P {“F) pm pm pm pm pm pm pm pm 0.6 Saa level 1140 10,99 § 66.9 | 16.48 | 100.4| 21.95] 133.8 27.45 | 167.21 32.951 200,71} 43,95 267.6| 54.95 [334.5| 65.90 | 401.4 1240 9.77 1 65,1 | 14,65 ] 97.61 19.53] 130.2 24.40 1 162.8} 29,30} 195.3 1 39.10| 260.4| 48.85 {325.5] 58,60 390.6 1340 9.02 | 64.6 | 13.52 ] 96,9 18.03; 129,21} 22,521 161,5| 27.05] 193.8 | 36.G5| 258.4{ 45,05 {323.0| 54,05} 387.6 0.6 35,000 1140 19.37 § 57.1 | 29,05 | 85.6] 38.70; 114.2| 48.40 | 142.8] 58.05] 171.3} 77.45| 228.4| 96.80 [ 285.5| 116.2| 342.6 1240 18,05¢ 59.1 { 27.05 | 88.6| 36.10] 118.2 | 45.10 | 147.B| 54.15| 177.3 | 72.20 | 236.4| 90.2 [295.5| 108.3} 354.6 1340 17.09 1 59.7 | 25,65 | 89.6) 34,20 119.41 42,75 | ¥49.2} 51,36 179.1] 68.40| 238,88} 85.50 [298.5| 102.5! 358.,2 0.9 Sea level 1140 11.97 1 76,21 17.95 | 114,31 23,95} 152.4] 29.95{ 190.5] 35.90} 228.6% 47,90 | 304.8} 59,90 [381.0{ 71.80 457.2 1240 10,331 74.0 |} 15.50{ 111.0} 20.65] 148.0 25.801 185.0} 31.00{ 222,01 41.30} 298,01 51.70 {370.,0] 62.00} 444.0 1340 9.29 0 7151 13,951 107.2| 18.60] 143.0| 23.20 178.8| 27.90; 214.5 | 37.15| 286.0| 46.45 {357.5| 55.75 429.0 1540 8.09 ] 68.8 | 12,131 103.2| 16.20| 137.6| 20.20| 172.0| 24.30| 206.4 ! 32.35] 275.2} 40,40 }344,0] 48,50} 412.8 0.9 35,000 1140 18.05! 61.Y } 27.05 ] 9L.6] 36.10| 122.2] 45,10 152.8] 54.15] 183.,3{ 72.20| 244.4] 90,2 |305.5{108.3 366.6 1240 16,681 62.2 1 25,00 93.3] 33,35) 1244 41.70] 155,5] 50,00 186.6] 466,70 248.8] 83.40 ;311.04100.0 | 373.2 1340 15,681 62,61 23.50{ 93.9) 31.35| 125.2{ 39.20} 156.5| 47.00{ 187.81 62.70| 250.4| 78.30 }313.0! 94.00} 375.6 1540 14.3 62,9 | 21.05{ 94.3} 28.10| 125.8] 35.10{ 157.2| 42.20) 188.6| 56.20| 251.6{ 70.25 {314.4} B84.30} 377.2 1.5 35,000 1140 18.66 | 81.4 1 28,00 122,1% 37.30| 162.8| 46.70| 203.5| 56.00] 244.2| 74.70| 325.61 93.30 | 407.01104,5 | 488.4 1240 16,22 78.4 1 24,35 117.6} 32,50 156.8] 40.60| 196.0{ 48.70) 235.2%{ 64.90] 313.6] 81.20 {392.0{ 97.30| 470.4 1340 14,331 75.7 | 21.501 113,6{ 28.70] 151.4{ 35.85] 189.2| 43.00 227.1} 57.30] 302.8] 71.70 {378.5| 84.00] 454.2 1540 12,27 | 73,2 18,40 109.81 24.55{ 146.4{ 30.70] 183,0| 36.801 219.6] 49.10} 292.8| 61.30 | 366.0] 73.6¢| 439.2 1.5 45,000 11406 27.27 81,41 40,851 122,11 54,50 162.8] 68,201 203.5] 81.80) 244.2|109.0 | 325.6{1356.3 407.0f 163.5} 48B.4 1240 23,45| 78.4 ) 35.20 | 117,86} 46,25 156.8]| 58.70¢ 196.0} 70.45] 235.2| 93.90] 313.6(117.30 §392,0] 141.,0} 470.4 1340 20,491 75,7 1 30.75 ) 113.6| 41,00 ¥51.4) 51.30) 189.2| 61.50| 227.1| 82.10} 302.8]|102.5 1378.5{ 123.0| 454.2 1540 17.28 | 73.2 | 25.90 | 109.8| 34.55{ 146,4| 43.201 183.0| 51.80| 219.5| 69.15] 292.8| 86.40 | 366.0} 103.6| 439.2 W = Propulsion machinery weight, 10_3 ib. om **P = Reactor power, megowaits, cheMeeg T S ORNL-LR-DWG 69 370 NN STRUCTURAL WEIGHT =30%, OF GROSS WEIGHT POWER DENSITY =5.0 kw/em?® 320 | fi J‘ e | o b 570 te/hr, 1240: F - 1r/hr, 1340 F - 1r/hr, 1540° F \ o 10 r/hr, 12407 F - % 10 r/hr, 1340° F 2 100 r/hr, 1240 F =3 $ | \ S // 220 S/ e S ;S ,,,,,,,,,, // / / /. / / // //// / // ] L/ / / k o b N S 1600 v/hr A540°F ; S 100 r/hr, 1540°F ________ 4 % 1000 r/hr, 1340°F 1000 r/hr, 1240°F | / : /// N~ 400 r/hr, 434C°F --------- ‘ %7/ RO /b, 1540°F T /I/ 7 % V/ SO0 R B R /42/2/ . ////4 ] ‘_ NS NN 200 300 500 , (Ibx 1073) Fig. 1, Chart for Determining Aircraft Gross Weight at Mach 1.5 and 35,000 §t for Yarious Turbine Inlet Temperatures. 10 TABLE 4, L/D RATIOS FOR YARIQUS FLIGHT CONDITIONS MACH NUMBER ALTITUDE L/D (f1) {without nacelles) 0.6 Sea level 15 0.6 35,000 15 0.9 Sea level 10 0.9 35,000 12 1.5 35,000 6 1.5 45,000 6 the crew shield) to show the effects of various degrees of shield division in relation to the im- portant design conditions (Figs. 2 through 5), dose rate expressed here, for simplicity, in r/hrf‘! is actually the personnel exposure dose rate (rem/hr) from radiation made up of seven-enghfhs gamma rays and one-eighth neutrons, assuming u relative biological effectiveness of 10. A number of important conclusions can be de- duced from Figs. 2 through 5. Perhaps the most™ important is that the gross weight of the airplane: is very sensitive to reactor power density and the operating temperature, except under the subsonic design conditions with power densities greater | than 1 kw/em®. For a power density of about 1 kw/ecm® and a turbine air inlet temperature of aboit 1200°F, an increase in reactor temperature level of 100°F is more beneficial than a factor-of-2 in- crease in power density. The turbine cir inlet temperature will be lower than the peak fuel tem- perature by roughly 400°F, depending on the heat exchanger proportions, and thus a turbine air inlet temperature of 1140°F might correspond to a peak fuel temperature of about 1540°F., Since it is doubtful whether reactor structural materials will be available that will permit reactor operating temperatures of much above 1650°F, it is likely that, to achieve turbine air inlet temperatures of much above 1200°F, it will be necessary to provide for interburning of chemical fuel between the radiator and the turbine, For reactor power densities of more than 1 kw/ecm®, the aircraft gross weight is not very sensitive 1o the degree of division of the shield, except in the range of reactor shield design dose rates below 10 r/hr at 50 ft. This effect occurs The .;,; - partly because the incremental weight of a given radial thickness of shielding material increases at o progressively more rapid rate as a unit shield is approached and partly because, for the particular . » . series of shields used, the secondary gamma rays “Iproduced in the outer lead layer become of about the same importance as the prompt gamma rays from the core if the lead thickness is more than about & in, The secondary gamma rays make it necessary to add disproportionctely large amounts of lead to reduce the dose rote from the reactor shield to below about 10 r/hr at 50 f, Effect of Chemical Fuel Augmentation It is possible to use the same basic techniques for investigating coses in which chemical fuel is burned between the radiators and turbines to obtain extra thrust for take-off, landing, and high-speed flight. If the power required (in Mw) is multiplied by 3413 Btu/kw-hr and divided by the lower heating value of the fuel (about 18,000 Btu/Ib), the equiva- lent rate of chemical fuel consumption is obtained in Ib/hr, that is, (Fuel consumption, Ib/hr) 3413 103 = (Power, Mw) ~——1---—>-<-----~-—-—- 18,000 = 190 (Power, Mw) . The weight of the burners and the related equip- ment required for chemical augmentation of nuciear- powered turbojets should be roughly 25% of the installed weight of the basic engine without radi- ators. Thus the extra weight of the equipment for interburning may be readily calculaoted by multi- plying column ;i of Table 2 by 25% and by the total thrust required. It is assumed that the weight of il 450 oo w———I———?—-?__Tm?T"?_W_ ......... i DENSITY (kw/cm?) \ \ | \ i | , . s00 o Ao i | ‘ ; | | REACTOR POWER ORNL—-LR-DWG 195 " I I i : l [ ! ! : | i : P Q i | ] > : [ / J | ( ( 2 - | | ’ : o ‘ \ : L (1_,:} 300 : \ - - 4‘ ! o 4o ! 230 £ L | | ] | 0 TURBINE AIRINLET . [ ‘ & TEMPERATURE (OF ) ‘[ | & | ! ‘ 1 | ‘ 250 S } LA 1s0 g L=t : = @ | ‘ = g ‘ ‘ x - L) —" 4L L. P L L 1000 DOSE AT 50§t {r/hr) Fig. 2. EfHects of Shield Division, Power Density, and Tuibine Air Inlet Temperature on Aircraft Gross Weight at Sea Level and Mach 0.9. the fuel tanks and lines can be offset by savings in structural weight that can be effected by re- lieving the wing bending and torsional loads judicious location of the f{fuel storage system. Therefore, the fuel tonk system is treated through as if the entire weight were made up of fuel. If additional turbojet engines are required for use with chemical fuel exclusively, their approximate weight in pounds per pound of thiust can be ob- tained from column i of Table 2 as functions of altitude and Mach number by multiplying by a factor of 1.25 to account for the burner equipment. 12 The performance of an aircraft with chemically augmented nuclear power is illustrated in Fig. 6 to show the effect of sprint range on gross weight for various reactor design conditions for a sprint condition of Mach 1.5 and 45,000 ft. A comparison with Fig. 5 shows that, for sprint ranges of 1000 to 1500 miles, the chemically augmented nuclear- powered airplane is lighter than ths allenuclear- powered airplane and the reactor power is much lower, especially for the lower reactor operating temperafures. landing would be reduced, and the chemical fuel Furthermore, the gross weight for GAML-LR-DWG 693 500 250 450 DOSE IN 5-ft DIA, 10-ft LONG CREW COMPARTMENT: 4 /hr STRUCTURAL WEIGHT: 30%, OF GRDSS WEIG 30,000 1b \ 400 ] ‘ TR 200 350 r?h £ 300 150 ”» 0 l.._ I Q@ wl = »n 250 o3 Qo x O e < o 2 £ 200 400 ! 1 MATERIALS TESTED } 2 1 NATURAL RUBBER ({32a43) 1 2 GR-S30 (32847} ; 3 BUTYL RUBBER GR-150(32546) 108 4 NEQPRENE W (32n44) - 5 HYCAR OR-15 (32845) | 5 HYCAR Poa-24 (32848) i 5 7 SILASTIC 7170 ‘! 8 THIOKOL-ST (3000-5T) 9 VULCOLLAN _ 10 HYCAR OR-15 (PISIC2) S 2 11 HYCAR 0S5-10 (PISIC3) L’ 12 THIOKOL (Pisic 7 8 13 HYPALON-S2 (PISICAN 2 q0® | - ® . ! 1] : | = O i e 5 ‘ 7 ] | | I J ‘ 2 ‘ \ 7 216 i3 7 4 8 5 ‘ 10 S T e © G 2 . o - i : f 5 2 , | : 5 4 6 ‘ . *» 0 o7 e28l83e] 108 S e - i NO DAMAGE SLIGHT DAMAGE MODERATE DAMAGE ORNL-LR-DWG 920 — }_ _ I - F- | | ‘ R 9! 10 13 1 e 8 o e j e !.11 l . » - . e | .- o & — i T2 g 6] -5 43 ] I . ,,,013 i *® a2 » e — i 12 11 10| | 11012 13— f - ® . | ! | 2 - - * @& & - ® -8 @ e _— _ 2 6 8 5 9 4 3 LT _ T } I [ P . 1 - | | | | ! | j i o e o SEVERE DAMAGE COMPLETE FAILURE RELATIVE DAMAGE Fig. 7. Radiation Damage to Representative Elastomers lrradiated in the ORNL Graphite Reactor. neutrons and the engines are located 15 ft from the reactor, the shield may be divided to the point where the total dose at full power 50 ft from the reactor can be as high as 1000 rem/hr without introducing any serious maintenance problems re- sulting from neutron activation of equipment outside the reactor shield. Ground-Handling and Maintenance Preblems it is instructive to examine ground-handling and maintenance problems on the premise that, because of radiation damage to orgariic materials and acti- vation of structural materials, division of the shielding is limited so that the full-power dose at 50 ft from the reattor may not exceed 1000 rem/hr. Three major types of work that would require people to be within 50 ft of the reactor shield are involved. The first is the regular maintenance 18 work that could be scheduled for a period during which auxiliary ground shielding might be arranged. The second type of work is ground-handling or maintenance immediately prior to take-off or after landing in the course of which the use of auxiliary shielding would be very awkward and expensive. The third type of work includes unscheduled acti- vity required by an emergency such as a fire or a crash immediately preceding take-off or following a landing when auxiliary ground shielding would probably not be cvailable. The shielding require- ments for these three major types of ground-handling and maintenance work differ considerably because of differences in exposure time and dose. The first type of work, which might be carried out with auxiliary shielding in place, covers the bulk of the regular maintenance operations. It is of interest to"note that experience in the B-36 flight-test program indicates that over 2000 man- hours of work of this sort must be carried out per test flight and that there is an average of one flight per week. While the dose rate to be ex- pected in the vicinity of the reactor after shut- down will vary with the amount of gamma shielding around the reactor, within minutes of the shutdown it will generally drop by a factor of at least 20 from the dose rate for full-power operation. If much of the full-power dose is from secondary gammas generated in the shield, the reduction upon shutdown will be correspondingly greater. Decay of the short-lived fission products will effect a further reduction in dose rate by a factor of about Z in the first day and, again, by a factor of 3 in the next three days; after that the dose rate falls off very slowly. These effects are shown more explicitly in Fig. 8. [f the ground personnel worked 40 hr/wk, it should be possible, with little inconvenience, to arrange that they spend only one-third of their time in the vicinity of the airplane, while the rest of the time couvid be spent an appreciable distance awaoy. However, because of the character of the mainte- nance work fthat would have to be carried out, « disproportionally large amount of time would have to be spent in the vicinity of the airplone immedi- ately after shutdown, A fair assumption might be that ground personnel would have to take one-half their total weekly dose during the first 8 hr following a landing. Thus if a man is to receive not more than (.35 rem/wk and he receives 0.18 rem during the first 8 hr following shutdown and spends 2 hr of that time at an average distance of 15 ft from the center of the reactor, the permissible dose rate would be 0.09 rem/hr at 15 ft. This dose rate ot 15 fi would give a dose rate of about 0.01 rem/hr at 50 ft from the center of the reactor. If no auxiliary ground shielding were used, this weould require that the shield be designed to give 0.2 rem/hr at 50 ft for full-power operation; only a unit shield would meet this requirement. Auxiliary ground shielding for a divided shield could probably be arranged most conveniently by draining part or all of the fuel or water from the outer hydrogenous region of the shield and replacing it with zinc bromide, mercury, or an oil-metal shot mixture. The shield structure would, of course, have to be strong enough to carry the resulting loads. Since a load factor only « little greater than 1 is required when the aircrafi is af rest on the ground because of the absence of dynamic leads, it should not be difficult to handle the structural problem. The second type of activity for which tolerabie dose levels must be set involves much shorter periods of exposure to radiation. This category covers the ground-handling work that will be re- guired prior to the installation of auxiliary shieid- ing immediately after a landing or immediately prior to a take-off after the auxiliary shielding has been removed. This work would include towing the airplane into position, last-minute funeup, checking or repair operations, ond the installation or re- moval of the ouxiliary shielding. While this work might be carried ouwt with highly specialized equipment, the cost and time involved could be cut tremendously if the dose level in the vicinity of the airplane could be kept sufficiently low so that personnel could carry cut the necessary opera- tions without special protection. If the dose rate were | rem/hr at 50 ft ond if no appreciable amount of werk within a 50-ft radius were required, the personnel might be permitted to get the bulk of their weekly dose, say 0.25 rem, in ¢ 15-min period. This indicates that, to meet the requirements for this second type of work, it would be desirable to design the shield to give not more than | rem/hr at 50 ft after shutdown, which would meuan about 20 rem/hr at 50 ft at full power. The third type of ground-hondiing work that should be considered in establishing shield speci- fications is that associated with emergencies. While it is very difficult to predict the character of the work and the time required to cope with the emergencies that might arise in connection with u fire ot a crash, a total dose of 25 rem is permifted under such circumstances by AEC regulations. If the time of exposure to radiation were 15 min, a dose rate of 100 rem/hr could be telerated on this basis. Since emergency operations might have to be carried out up to 15 ft from the center of the reactor, it oppears that the shield should not be divided beyond the point where it would give 100 rem/hr at 15 ft immediately after shutdown, that is, 200 rem/hr at 50 ft at full power. If the crash were so violent aos to strew the surrounding area with fission products, the shield would be ineffective and the dose would not be a function of shield design; hence such a case does not pose o shield design problem,. 9 20 IN PER CENT OF FULL-POWER DOSE DOSE AFTER SHUT-DOWN ORNL-LR-DWG 1135 _ { - I o ] | T=OPERATION TIME (hr) Q . ,ENT':'L, ,,,,,,, _ | —— 0 100 200 300 400 500 SHUT DOWN TIME (hr) Fig. 8. Dose ot 50 ft from Reactor as a Function of Time After Shutdown, 600 The results of the above discussion have been summarized in Fig. 9, The left column gives the dose at full power 50 ft from the center of the reactor as an index of the degree of division of the shielding. The chart is applicable to any shield design. The first four columns simply give dose rates for representative conditions. (Note that the characteristics of any particular shield design fall along a horizontal line.) Typical radiation damage limits are given in the fifth column., The sixth shows the amount of auxiliary shielding required to reduce the dose after shutdown to a level that will permit @ man to work within 15 ft of the center of the reactor for 2 hr shortly after shutdown and for a total of an additional 10 hr during the suc- ceeding week (assuming one flight per week). The last two columns show the effects of neutron activation of turbojet-engine parts. In re-examining Fig. 9 and the preceding dis- cussion, it appears that the reactor shield may be divided to the point where at full power it would give 100 rem/hr at 50 ft from the center of the reactor without imposing exceptionally difficult limitations on ground-handling and maintenance operations. point where at full power it would give 1000 rep/hr 50 ft+ from the reactor, radiation damage to elas- tomers and greases would be serious. Also, ground operations would be severely restricted and time- consuming, and much expensive, specialized remote-handling equipment would be required. Shield Weight The weight of a carefully designed redctor shieid depends primarily on the reactor power, the power density, and the specified full-power radiation dose level at a given distance from the reactor, usually 50 feet. It is also heavily dependent on the dis- position of equipment such as pumps and heat ex- changers inside the shield and the presence of voids such as ducts and headers. Another factor is, of course, the kind of shielding material used. A diligent search for superior shielding materials has failed to disclose any that are markedly su- perior fo a combination of lead ond woter {the lead for gamma-ray attenuation and the water for neutron attenuation). While the investigation of materials is not complete, the most promising combination of shield materials found thus far is vranium, bismuth, and lithium hydride. A shield of these materials might moke possible a shield weight saving of as If the shield were divided beyond the much as 15% in comparison with the more conven- tional shield of lead and water. The shield weight also depends on the weights of the shield structure and of the cooling equipment required to dissipa}e the energy of the radiation absorbed in the shield. These items have been responsible for increases in shield weight of as much as 20% in some de- signs, and they may increase the weights for shields of special materials more than they in- crease the weights of lead-water shields. It should be mentioned that jet fuel is as effective as water as o neutron shield on a volumetric basis. The lower density of the jet fuel gives a small saving in neutron shield weight that is largely offset by the additional lead required for gamma shielding. A general idea of good shield design practice can be gained from a highly simplified approach. Roughly, one fast neutron and one hard gamma-ray (over 1.5 Mev in either case) escape from the reactor core per fission. This radiation can be attenuated by a factor of 2.72 by a thickness of approximately 3 in. of lead or water for the neutrons or thicknesses of 1 in. of lead or 10 in. of water for the gamma rays, The neutron flux from a 200-Mw reactor must be attenuated through the shielding material by a factor of about 100,000,000 if the resulting neutron radiation dose is to be reduced to 0.125 rem/hr (that is, one-eighth of the total dose), and the gamma flux must be attenuated by a factor of about 1,000,000 if the resulting gamma dose is to be reduced to 0.875 rem/hr (seven-eighths of the total dose) to give a total dose of 1 rem/hr, About 20 attenuation lengths will be required for the neutrons and about 15 for gamma rays, Since the fast-neutron attenuation lengths in lead and water are about the same, this means that about 60 in. of shielding material must be interposed between the reactor and personnel 50 ft away to cut their neutron dose rate to 0.125 rem/hr. Since .60 in. of water represents only six attenuation lengths for gammas, about 10 in. of the shielding material would have to be lead instead of water to cut the total dose rate to 1.0 rem/hr. The situation is complicated by the generation of hard, secondary gammas from inelastic scatter- ing of fast neutrons in lead or structural materials and from neutron captures in hydrogen, lead, or structural materials such as steel or aluminom, The production of secondary goamma rays can be inhibited by introducing boron or lithium to absorb the neutrons as soon as they are slowed down by 21 cC DOSE AT 50 ft FOR FULL -POWER OPERATION {rem/hr) ORML-LR-OWG 2009 6 7 5 RADIATION EFFECTS CAUSED BY — {0 — 10 — 1.8 sec — 5% 10 — [ o NEUTRON ACTIVATION OF AN ENGINE = 15 it FROM THE REACTOR CENTER ~ E AFTER 400 hr AT FULL POWER « = (/g OF DOSE iN NEUTRONS) A w % r ~ = RE DAMAGE T = - 2 L 40° L o° O | 18 sec = sx0f SEVERE DAMAGE TO =l 12in. OF LEAD (OR Er— ax10°f — o o T PETROLEUNM _UBRICANTS O FQUIVALENT —~ o 4 — 0w w O £ D O 2 £ o r o = T £ £ = i © o - . “’ < L - il 3 z g s < > SEVERE DAMAGE TO TIRES, o o ez = > K W L 10* = 10° S 3 min B sx10” _ [ RUSBER HOSE, O-RiNG, AND - K- 10in OF LEAD - 4xi0 W G0 Oy 5 = = = - W &0 . 3 5 § Sy 2 3 2 * 3 = o o © ‘5 —1 i E 3 L 10 O 0 = 5hr T 540 2 L_“I_J#%—Sin.OFLEAD Z1— 4x 10 W g 10 - Ll - W - = & = o <1 =~ o < - = or £ — O > l w Ll @ c r = - &g o wn o = = < T O Il Q Pt E o = 2 0 0 o O = < = T T w w ~ < x =3 = 2 = = -~ - D — o = © w? E | son 5 < T Ol 4inor Lean = L 4 r o & - in. L Z e — o w <1 READ HORIZONTLLLY 4CROSS i % - = - 5 [4p] & M a o n - ) 8 | SETS L 500 o L 5% 10 .~ REPLACE WATER al— - BY ZnBry, Fig, 9. Approximate Effects on Ground-Handling and Maintenance of Degree of Division of the Shielding. el A the water shielding. (B'® and Li® are the only materials that do not give off hard capture gamma rays.) The production of secondary gammas in the lead can be kept to an unimportant level by dis- tributing the lead through the shield in such a way that the neutron flux in the lead af any given point is low enough so that the secondary gamma flux produced there is below the local level of the This concept of the ‘‘matched’’ shield?® has proved invaluable. It is equally applicable to the disposition of primary gamma flux from the core. structural materials. The above concepts can be best illustrated by examining their application te a ftypical shield design. The arrangement of reactor, heat ex- changer, pressure shell, and shield assembly shown in Fig. 10 was evolved in an effort to get the lightest possible over-all assembly consistent with reactor physics, heat transfer, and other re- quirements for a circulating-fluoride-fuel reac- tor. 1821 The arrangement is such that, except for the pumps at the top, the varicus regions are enclosed by surfaces of revelution about the verti- cal axis of the reactor. It was found that the re- flector around the reactor core should be at least 12 in. thick and should be followed by a layer of about 0.13 in. of B0 if the Inconel pressure shell were to be kept from becoming a more important gamma source than the core insefuras gammas leck- ing from the shield surface are concerned. Similar reflector and boron-layer dimensions were found to minimize the activation of the secondary fluid in the heat exchanger by neutrons from the core. It was also found that decay gammas from the fuel in the heat exchanger would make the heat ex- changer a gamma source of about the same im- portance as the core, and therefore little would be gained by placing gamma shielding inside of the heat exchanger. It was found that attenvation of the fast-neutron flux by the 12-in.-thick reflector would be sufficiently great that o lead layer of up to & in. in thickness could be placed just outside the pressure shell without creating a seriously high level of secondory gamma-ray productien in the outer lead layers of the shield. The only fairly complete set of shield weights available to show the effects of reactor power, 281_“ Tonks and H. Hurwitz, The Economical Disitri- bution of Gamma-Ray Absorbing Maferial in a Spherical Pile Shield, KAPL-76 (June 8, 1948). power density, und shield division is a set com- puted for the basic arrangement shown in Fig. 10, This set is presented in Figs. 11 to 15, The shield weight increases with power at much less than a ltnear rate, and ot o given power, it is not very sensitive to reactor core diameter for core diameters of less than about 24 in.; however, it becomes progressively more sensitive for larger cores. The data for seven representative cases have been crossplotted in Figs, 16 and 17 to show that the total reactor und crew shield weight is not very sensitive fo the degree of division of the shield, except for the nearly unit shields in which the lead thickness exceeds & in. and hence dis- proportionally large amounts of lead must be added to take care of secondary gammas preduced in the outer lead layers, A considerable saving in weight might be effected by distributing part of the thick lead region throughout the water; however, such a step cannot be taken effectively until experimental test dota ore available. The lorger crew compart- ment of Fig. 17 gives less incentive for dividing the shield thon does the smaller compartment of Fig. 16. An effort was made to estimate the shield weight for a set of “‘ideal’” lead-water shields for the de- sign conditions used for the engineered shields described chove?® so that some insight into the effects of the heot exchuanger, pumps, struciure, etc. could be obtained. The enly daote available 1o indicate the proper distribution of the lead to give an ideal matched shield cre from experiments carried out by Clifford®® to establish the lead dis- position for minimum weight in a lead-water shield for a 36-in.-dia reactor. Perturbations were applied to Clifford’s data te obtain estimated fotal shield weights for a power density of 2.5 kw/cm® and a dose 50 ft from the reactor of 10 rem/hr with o crew shield designed to give 1 rem/hr inside of @ crew compartment 5 {4 in diameter and 10 1 in length, The results are plotted in Fig. 18. To facilitate comparison, similar data for enginsered shields (taken from the calculations made for Fig. 12) have been plotted on the same coordinafes, 29R, M. Spencer and H. J. Stumpf, The Effeci on the RMR Shield Weight of Varying Neutron and Gommoa Dosse Components Taken by the Crew and Comparison of the RMR Shield ‘Weight 1o That for an ldeafized Shield, ORNL CF-54-7-1 {to be published). 30¢ E. Clifford et al., ANP Quar. Prog. Rep. May 10, 1950, ORNI_-768, p. 36. 23 DWG. 18743 MAKEUF FUEL TANK-- L~ PUMP QUILIL SHART .~ CONTROL ROD ACTUATOR ’ »~ PUMP QUILL. SHAF'T THERMAL INSULATION -, THZRMAL Y / INSULATION Nok DUCT ™~ ... ot REACTOR PRESSURE SHELL - /. THERMAL NSULATICA #{f e / ] LEAD SHIELDING - il 4 4 COOLING PASSAGE — I BCRATED WATER ——- DRAIN VALVE —— FLEL DRAIN L NE -—" N - SEIF - SEALING RUBBER TANK VERTICA. SECTION LEAD SHIELDING - -~ REACTOR PRESSURE SHELI . -~ THRERMAL INSULATION THERMAL INSULATIGN~. “._WEB OF CANT.LEVER BEAM SUPPORTING REACTOR AND SHIELD TJBE BLNDLE - e , SELF—SFEALING RUBBER TANK — - BORATED WATER Q & 12 24 HOR!ZONTAL. SECTION e ] SCALE N INCHES Fig. 10. Sections Through a Lead-Water Type of Shield. 24 ORNL - LR-!QE 682 180 160 140 REACTOR AND SHIELD ASSEMBLY WEIGHT (ib X 40 ) 120 100 8C 60 40 l : O {00 200 300 400 500 600 700 80C REACTOR POWER (Mw) Fig. 11. Effects of Power Output and Core Bicmeter on the Weight of the Reactor and Reactor Shield Assembly (No Crew Shield) for Dose 50 ft from Recctor of 1 rem/hr (i/ Neutrons, / Gemmas). 9 : ORN -DWG 683 o . o - ,/ ~ ] 7 ,/ .:"":,/ ~ CORE DIAMETER, 45.3 in. P / e R AN L 7 2 POWER DENSITY, 1.25 kw/cm3—\ / e //:/ N f =~ 1 \ 4 m CORE DIAMETER, 36.0 in. ')/ -~ q/'/ /ff € o0 \ // P . ® / - AT }-Iu / \ (// / 7‘4\( \\ l x ’ //S'/ N \ PO\.«;ER DENSITY, 5.0 kw/om3 < : - : — , B cm > A7 P /))( N\ | | | g GO // ‘/ /; ~ A, N 1 ! . ! = /s /fi — POWER DENSITY, 2.5 kw/ecm3 o Z / \ | ! | 5 7 = \ < ] o /// ya //‘ \ — CORE DIAMETER, 28.5 in. L = s 5 | | T 80 ¥ ! t ! ; ,/ s — CORE DIAMETER, 22.7 in. < 2 M\ : 7/ \ 5 /) // = . L /é\ CORE DIAMETER, 8.0 in. N, 7 7 40 ¥ \— CORE DIAMETER, 14.3 in. | | | 20 ? : o 100 200 300 400 500 600 700 800 900 REACTOR POWER ( Mw) Fig. 12. Effects of Power Qutput and Core Diameter on the Weight of the Reactor und Reactor Shield Assembly (No Crew Shield) for Dose 50 ¢t from Reactor of 10 rem/hr (1’8 Neutrons, ?’7;3 Gammas). ORNL-LI!-I DWG 6B4 140 120 | 1CO 30 60 40 REACTOR AND SHIELD ASSEMBLY WEIGHT (Ib X10 %) 20 0 100 200 300 400 500 600 700 800 300 REACTOR POWER (Mw) ’ Fig. 13, Effects of Power Qutput and Core Di«zmeter or the Weight of the Reactor and Reactor Shield Assembly {No Crew Shneid) for Dose 50 f¢ from Reactor of 100 rem/he (/5 Neutrans, 7 % Gammas}, 8C ORNL - LR- DWG 685 12 O . l! E ; 5 ; i F a t i i * x 5 5\«;% , E : R’ OF’N % L 100 @0‘*“25 Wi 10%/ | T A REACTOR AND SHIELD ASSEMBLY WEIGHT (ib X 70“3) | J 0 100 200 300 400 500 600 700 800 900 REACTOR POWER {Mw) Fig. 14. Etfects of Power Output and Core Diameter on the Weight of the Reactor and Reactor Shield Assembly {No Crew Shield) for Dose 50 {1 from Reactor of 1000 rem/hr (3‘8 Mautrons, 1?8 Gammas). 60 W o B o TOTAL WEIGHT OF CREW AND SHADOW SHIELDS (b X 10 °) n O o o O Lo 6¢ ORNL-LR-DWG 6B L N I !;f | | | Pl HEa i A /f . K | | E. Q;l"\q, : i / / | } I 1 L p _ ! *\L///, 2O >d n Pt \% 5 Pl ' ! : L [ _t 20 50 100 200 500 GO0 2000 5GC0 10,000 50,000 DESIGN ATTENUATION OF DOSE FROM REACTOR SHIELD (r/hr) Fig. 15. Effects of Design Attenuation on Crew Shield and Shadow Shield Weights for Equal Attenuation of Neutrons and !G:o'mmu's. 0t ORNL—~LR—CWG 457 1.4 X 10 1.2 X10° 8.0 X 104 TOTAL WEIGHT (Ib) 6.0 x 10% 4.0 % 10? 2.0 x 10 N — |8} N O 50 100 200 500 1000 2000 DESIGN DOSE 50 ft FROM REACTOR SHIELD {r/hr) Fig. 16. Effect of Design Dose from Reactor Shield on Total Reactor and Crew Shield Weight for o Dose of 1 rem/hr in a Crew Com. pariment 5 ft in Diameter and 12 #t in Length, ' DWG. 21!'5’1 120 O 285'” dth = 1 ORE 200)\;1 ' £ 80 _22.7-in- c‘“GCORE 200 - M | = - ) 5 3 L % 18.0-in.-dig CORE 200 Mw P = 18.0-in-dia CORE; 100- MweJ‘ = . o — —48.0- in-dia CORE; 50 - Mw 1] 180m dmCORE 25 - Mw | - 1 B o 40 ; ( L - | | | | | ‘ < | ; | : Do ; ! - | ; i J | | : L | o ] AR ] ; : . : i i i : : i t | ! 5 by bbb ISR 1 2 5 {0 2 5 100 2 5 1000 2 CESIGM DOSE 50ft FROM REACTOR SHIELD {r/hr) Fig. 17, Effect of Design Dose from Reactor Shield on Total Reactor and Crew Shield Weight for a Dose of 1 rem/hr in a Crew Come partment § ft in Diameter and 12 ft in Length. ¢t 189 | | | [ | | | | : | | | | | | | | | | | ~ i | | SOURCE OF WEIGHT INCREMENT S I R I ] oo} —— g - ‘ TOTAL REACTOR, HEAT EXCHANGER, } DUC“‘fS, PUMPS, lflND STRUQTURE ‘ AND REACTOR SHIELD WEIGHT FOR R¥R ~ % ExTRA LEAD FOR DECAY 140 ; | 5 \ | GAMMAS FKOM HEAT EXCHANGER | o | i | INSERTION OF HEAT EXCHANGER, : : PRESSURE SHELL, AND LUMPING OF 120 ! : T LEAD OUTSIDE OF PRESSURE SHELL | : 1 ‘ o L | ~ 400 F— —— — — __(2 . g | hu ! 3 9 ] [’ - e [ -l ] z 80 \ | | | REACTOR AND SHIELD WEIGHT FOR IDEALIZED MATCHED LEAD-WATER SHIELD ! 1 ; o — ; e e e e : | | | | I 5 | 5 DOSE: 10 r/hr AT 50 ft FROM REACTOR = ‘ ‘ (/s gammas, g neutrons) ! : | 40 1 -t —_— - - i - - g [ POWER DENSITY: 2.5 kw/em? ‘ 7 | | i \ | i -3 | | | | | | | i 20 | = bt — WL—ffl—ffl S — ! 3 ‘ j ‘ ‘ : : : L N , 5 | | ! ‘ l } | 0 | | : . 1 | i | \ ‘; 0 100 200 300 400 500 600 700 800 900 1000 1100 1200 REACTOR POWER (Mw) - Fig. 18. Comparison of Reflector-Moderated Reactor Shield Weight with Weight of Corresponding ldealized Matched Lead-Water Shield. The major factors that cause the weight of the engineered shield to be greater than that of the idealized shield are the insertion of the heat ex- changer and the pressure shell and the fumping of the lead into a single region immediately outside the Because of the lumping of the lead, however, there is little increase in shield weight required to take care of fission-product decay gammas from the fuel in the heat exchanger. The ducts through the shield increase the weight, but this increase is not very serious because the ducts are filled with liquid and their cross- sectional area is small. The very much larger ducts required for an air- or helium-cooled reactor would be re- sponsible for far more serious increases in shield weight. ‘ In re-examining Fig. 18, it is interesting to note that, in the reactor power range of greatest current interest (200 Mw), the engineered shield is about one-third heavier thon the idealized matched shield. While this is a rough estimate since all the shield weight estimates are subject to errors totaling about 10%, the character of the calculations was such that the values for the various conditions should be comparable. Since the shield weight data given in Figs. 11 to 15 were used in the parametric aircraft perform- ance study presented earlier, it is pertinent tfo consider the applicability of this data to other types of reactor. Since most other reactor types make use of cores in the form of right circular cylinders, an allowance for their less favorable geometry must usually be made. 1t has been shown that the diameter of o sphere equivalent to a right circular cylinder from the shielding standpoint is about 25% greater than that of the cylinder.? Since the cube of 1.25 is about 2.0, it can also be said that the effective power density for the cylinder shouyld be considered as roughly twice its average actual power density to be equivalent to that for a spherical core of the same diameter. QOther items that might have substantially dif- ferent effects in other reactor types would be the heat exchanger and the ducts. From Fig. 18, it appears that elimination of the heat exchanger might make possible a weight saving of about 15%. Most of the other items considered in Fig. 18, ex- cept the ducts, seem to hove small effects for other reactor types. Insertion of the large voids required for the ducts and headers of a gas-cooled reactor would be responsible for major increases in weight, pressure shell, fead-water particularly for the more nearly unit shields. In fact, since the shield thickness increases little with reactor power, while the duct flow-possage areg must increase in direct proporfion to reactor power, and since the radiation Jeakage through a duct increases more rapidly than its cross-sectional area, many shielding experts feel that the extra weight required for the ducts for a near-unit shield for a high-power gas-cooled reactor represents an extremely large weight increment. The problem is clearly quite formidable, as is indicated by the fact that to date no shield weights for gas-cooled reactors have been published except those for highly divided shields that give 80,000 rem/hr or more at 50 ft from the reactor. While some weight savings might be effected by using shielding materials other than lead and water, no substantially lighter engineered design based on such materials has been prepared to date. In any event, a weight saving of more than 10% through the use of special materials seems uniikely. After reviewing all the above-mentioned factors, it ap- pears that there is little likelihood of getting an operational reactor and shield assembly that will weigh less than perhaps 85% of the values given in Figs. 11 to 15. NUCLEAR PROPERTIES Design proposals for high-powered reactors have ranged from those for the near-thermal water-moder- ated reactor of the supercritical-water cycle to those for fast reactors, as can be seen from the values given in Table 5 for median energy for fission. The important aircraft reactor design proposals are compared in Tobles 5 and 6 with other representative reactors. In general, it has appeared that the higher the median energy for fission, the greater is the crifical mass. This is particularly true for solid-fuel-element reactors, because relatively lorge core volumes of at least 2.0 #% are required to satisfy heat transfer and fluid flow requirements (to be discussed in a later section), Further, because of the shorter neutron lifetime inherent in the faster reactors, it has been felt that they would present markedly more serious control problems. Because of these factors, all the reactor designs that have looked promising enough to receive considerable attention for aircraft application have been thermal or epithermal, that is, have hod a medion energy for fission of between 0.025 and 1 ev. As is evident 33 ve TABLE 5. REACTOR PHYSICS DATA FOR REPRESENTATIVE REACTORS 1 | | . MTR STR sCw SIR EBR I ARE AC-100 | RMR Critical Mass i | | Clean, kg of U233 2.3 9.5 18 52 48.2 { 8to15 | 24.5 18 Poison allowance 500 g | 8.7 kg 5 kg/1000 hr | None (a) (@) at 62.5 Mw ' Burnup allowance 38 g/day 2.0 kg 2 kg/cycle | 87.2 g/day 1 g/day | 1.5 g/day 0.833 g/hr ' 2.1 g/hr ‘ at 62.5 Mw | ! Power, Mw 30 70 700 62.5 1.4 1.5 20 f 60 Hours at full power 600 600 144 900 > 3000 1000 | 100 1000 Per cent thermal fissions ~95 94 ~40 0.7 0 65 |76 35 | Mean neutron lifetime, ~15 8 ~1 ~2 1 14 4.214 P40 sec X 10—5 ‘ f | l | f Temperature cosfficient of | ~15x107% 3.9 10 15| 5 x 1075 | 1p-5 255x1075 | 2.8x 1075 | 3.6 10-40) | 7 10-5 reactivity, per °C i x 1079 E ! Peak-to-average power density \ 1.8 3.9 ~1.5 2,28 1.25 ; ~2.0 l 1.3 1.7 | {longitudinal) Fuel heat capacity, cal/°C 22,9 x10%(€) | 37 10% | ~5 x10¢ | 4.9 x 105 1.6 x 10° 3.8 x 194 26.3 x 10° 410 6 x 104 Median energy for fission Thermal Thermal | ~0.1 ev 52 ev Fast Thermal Thermai P V0.2 ev i | Free-flow ratio in core 0.63 F ~0,7 ~0,3 0.3 0.273 0.007 0.428 ! 0.75 Power density, kw/cm> ]! 0.3 .10 1.0 0.218 {avg.) 0.167 .03 0.05 ] 1.0 ( Preponderantly a slow ©ar 127°C. a)FueE to be added as required. temperature coefficient associated with the water moderator, TABLE 6. DIMENSIONS AND COMPOSITIONS OF REPRESENTATIVE REACTORS REFLECTOR | REFLECTOR I1ZE COR M 0 REACTOR CORE S E COMPOSITION {vol %) THICKNESS COMPOSITION MTR 73 % 23.4 X 60 cm 0.25, U3, 36.66, Al; 63.09, H,0 | 12 in. Be + 44 in. C | Be, C STR 36.4 in. dia x 43 in. high 0.15, U235; 57.8, Zr; 42.0, H,0 2 in. HZO SCW 2.5 ft long % 2.5 §t dia 0.3673, U233, 17.36, stainless 2Y in, H.,0 steel; 82.272, H20 SIR Hexogonal prism (eff.) 55.2, Be; 31.1, Na; 7.9, stainless | 8 in. Re 27.25 in. high x 27.25 steel; 2.2, MgO; 1.8, UO:2 (93% in., dia enriched}); 2.0 void EBR Hexagonal eylinder 7.5 in, | 53, uranium; 13.5, stainless 10 in, U238 i u?38 high, 7.5 in. across flats steel; 33.5, NaK breeding blanket ARE 33.0 in, dia x 35])3 in, high | 82.6, BeO; 7.6, fuel; 4.8, Na; 2.2, 7!‘{2 in. BeO inconel; 2.8, void for rods coolant: Na AC-100 Hexcgonal eylinder 30 in. | 40, HEO; 4.7, Al; 11, insulation; ~3 ft H20 high, 29.5 in, across 5.2, fuel, including UOZ; 39.1, flats, and 33.4 in. across void diagonal RMR Sphere, 23.75 cm radius 40.98, F; 39, Na; 17.47, Zr; 31 em Be 2.55, U3 in Table 6, this has necessitated that the concen- tration of iron-chrome-nicke! alloy structural ma- terial be kept to less than 18 vol % and that the microscopic neutron absorption cross section of the coolant be less than 1 barn. In fact, the only reactor listed in Tables 5 and 6 that does not meet these conditions is the EBR, a nonmobile reactor, Moderating and Reflecting Materials At first glance there appear to be several ma- terials to choose from for the moderator, Beryllium, beryllium oxide, graphite, water and heavy water, sodium hydroxide, sodium deuteroxide, lithium, copper, lead, and bismuth all might be used as either moderating or reflecting materials. Beryl- lium, beryllium oxide, D,0, and graphite have such low capture cross sections that they may be used in very thick sections without serious loss to the neutron economy. Thermal stress considerafions make beryllium oxide of doubtful value for reactors having core power densities of greater than 0.5 kw/cm?®, even though beryllium oxide is one of the best of the ceramics from the standpoint of thermal- shock resistance. Beryllium and grophite appear to be satisfactory from the thermal stress stand- point, although they present other problems, The cost of fabricated beryllium must be expected to be from $75 to $300 per pound, while the cost of reactor-grade graphite is only about $0.15 per pound, However, its much higher atomic density and its better high-energy scattering cross section make beryllium much superior fo graphite on a volumetric basis. The use of beryllium gives amuch more compact reactor and hence a much lighter shield, Normal water has such o short diffusion length that it may not be used in sections thicker than 1 in. without excessive loss of neutrons to captures in the water, Partly because of this and partly because of its predominantly forward scat- tering, normal water is much less effective as o reflector than beryllium, beryllium oxide, or graph- ite. For the same reasons, much the same can be said for NaOH, NaOD, and Li’OD. The properties of the principal moderating materials are shown in Table 7, 35 TABLE 7. PROPERTIES OF PRINCIPAL MODERATING MATERIALS H,0 BE BeO D,0 | NaOH | Li’OH C Density, g/cm> 1.0 1.84 2.84 1.1 1.8 | 1.4 1.6 Age-to-thermal, cm? 33.0 | 98.0 105.0 120.0 | 120.0 350.0 Thermal diffusion length, cm 2.88 | 23.6 28.5 100.0 5.0 | 5.9 50.0 Thermal conductivity, 0.35 | 48.0 15.0 0.35 0.7 72.0 Btu/hr-f?2-(oF/ft) Therma! expansion coefficient, 10.0 x TO“é 5.5 x ]0m6 1.1 % 10'—6 in./in.-°F Modulus of elasticity, psi 40.0 x 10° 42.0 x 108 1.5 % 108 Effect of Moderating Materiol on Design SECRET DWS. 14428 Any detail design is heavily dependent on the materials used, and many different materials combinations appear interesting at first glance. If moderating material is distributed throughout the core, it displaces fuel and coolant and makes the core larger for a given power than would be re- quired by heat transfer and fluid flow consider- ations. Normal water can constitute as little as 25 vol % of a recctor core for which the fuel in- vestment is kept to within tolerable limits., |f beryllium is used, at least 50 vol % of the core should be occupied by moderator, unless the principle of reflector moderation is employed, in which case a fairly uniform power distribution can be obtained with as little as 25 vol % of beryllium, Much the same relations hold for D, 0, NaOH, and Li’OH as for beryllium. The relationships on which these observations are based were discussed in earlier reports3'+32 from which Figs. 19 to 22 were taken to show these effects, [f allowances are made for the volume required for structure, control rods, etc,, the ratio of the flow passage area for the reactor coolant to the cross-sectional area of the reactor can hardly be better than indi. cated in the free-flow-ratio entry in Table 5 for representative reactors, Hydrogen is such an obvious choice as a moder- ator that some further remarks about its limitations must be made. The forms in which hydrogen could be used in a reactor are limited, namely, water, a 31w, K. Ergen, ANP Quar. Prog. Rep. Mar, 10, 1952, ORNL-1227, p. 48. 32c. B. Mills, ANP Quar. Prog. Rep. Dec. 10, 1951, ORNL-1170, p. 14. 36 j | CRITICAL MASS DETERMINED BY i r BARE REACTOR MAETHOD, USING A | | REFLECTOR SAVINGS VALUE OF £ g (THE DIFFUSION LENGTH) (10) MASS CRIT.CA: URANIUM CORE DIAMETER (ft) Fig. 19, Critical Mass vs Core Diameter for Hydroxide Moderated Reactors with Thick Hydrox- ide Reflectors. hydroxide, an organic compound, or a metal hy- dride. |f water were used, either it would have to be kept at a pressure of around 5000 psi, which would pose exceedingly difficult structural and pump seal problems, or it would have to be ther- mally insulated from the hot zone of the reactor, a measure that would be wasteful of core volume and would probably introduce poisons. An even KOH . NaOH N i URAMIUM REGUIRED FOR CRITICAL MASS {(wt %) CORE DIAMETER (ft) (o) Fig. 20. Digmeter for Yarious Hydroxides. Fuaction of Temperature. DM M I5658 - ORNL—-LR—DWG 11328 URANIUM SOLUBLE {wi %] 500 550 600 650 TEMPERATURE (°C) (5} (a) Weight Per Cent of Uranium Required to Achieve Criticality as a Function of Core (b) Weight Per Cent of Uranium Soluble in Varicus Hydroxides as a ORNL. - A== SOWG 1140 - [0 [ ] EXTERNAL SYSTEM VOLUME 100 | A IS ASSUMED TO BE 16 fi® 177 b L . J A 3-ft DIA CORE TOTAL 2 3Y,—ft DIA CORE IES‘F":'T%";Y ft 2 i 4= 11 DIA CORE £ 60 | bt i ------ SR . = | ; g Ll Ll o 2 L] = 40 e e oo 5 4-f1 DIA CORE | URANIUM : /j, 3%-ft DIA CORE N 50 | bl L] 3-f1 DIA CORE CORE - | =T ,,,,,, | s eedheee e O ] - 0O oM 020 030 040 030 060 070 VOLUME FRACTION FUEL~COOLANT IN CORE Fig. 21. Total Uranivm Inventory and Uranium in Core as a Function of the Volume Fraction of Fuel-Loolant in the Core for BeO-Moderafed Circu~ lating-Fuel Reactors with 3, 3/- and 4.-ft-dia Cores and NaF-UF , Fuel-Coolant. 240 Dt \ 177 ] ‘ T 1 _FUEL CODLANT Lif - ’\lal KF f‘Jr, T _____ [ B T VOLUME OF THE SYSTEM OUTSILE ] | | 1 ' 200 THE CORE: 150 fi¥ | ) )' T WATER OENSITY: 09567 arem® | L i' | i = | £ - 3%, ‘H UA CORE | ! 4 ¥ 4 22 4 (a2l ftDia coge > 5 5 2 = "n\ B i 5 3-fr-DIa CORE | T > - . ——eeers foooeom freneene L w 5/? fH-D1A CORE 5’# 7 - 3-ft~DIA CORE Z 50 L ke 2V Fr-DIA CORE | 2 p ..‘[__.__%_.___ b i - ol b oL | L 020 030 G40 030 B0 . Q70 030 030 040 VOLUME FRACTION FUEL ~COOLANT IN CORE Fig. 22, Total Uranium Inventory and Uranium in Caore as a Function of the Volume Fraction of Fuel- Coo!nn‘l in the Core for H,0- Moderuted CirculatingsFuel Reactors with 2/- 3-, and 3/’- ft-dia Cores and LiF-NaF-. KF«UF4 Fuei-'Cooiani'. more important factor if the water were thermally insulated would be that between & and 15% of the reactor output would go into heating the water; thus not only would heat be wasted, but the wasted heat would have to be dumped through o radiator at low temperature, and the radiator would impose a weight-and-drag penalty equivalent to a further loss in power-plant output of ot least 10%, An over-all performance penalty of 15 to 20% seems to be a stiff price to pay for the privilege of using water as the moderator, The design compromises that would be necessary to cope with problems of distortion and differential pansion would probably entail still further penal- ties, thermal thermal ex- Various organic compounds of hydrogen have been suggested as moderators, for example, di- phenyl oxide, cyanides, etc. However, radiation damage tests on organic compounds indicate that this is not a promising course because the gamma flux in the moderator would be about 10'° gam- mas/cm?.sec for a reactor core power density of 1 kw/em®. All orgonic liquids tested to date have shown severe radiation damage after an inte- grated gamma flux of, at most, 1078 gammas/cm?2. This would give an operating life of only 20 min for the moderator material, Not only would radi- ation decomposition of the moderator fluid present a problem, but it seems likely that deposits of carbon and sludge on heat transfer surfaces would tend to render them ineffective. Metal hydrides might prove sufficiently stable under radiation, but none with truly satisfactory physical properties hos been developed to date. Hydrogen gas is too diffuse for use as a moderator, and liquid hydrogen would present cooling problems inconsistent with high-temperature aircraft reactor design. The lowest estimated critical masses for the various configurations considered are for some of the hydrogen-moderated cores. However, there is less chance to get a low critical mass in o high- power reactor through the use of hydrogenous moderators than appears ot first glance, because the data of Figs., 19 through 22 do not include allowances for temperature effects, coatrol rods, burnup, ond fission-product poisons. The low- critical-mass reactors are highly sensitive to these poisons and require much larger allowances to take care of them. This can be deduced from the first two lines of Table 5. The same data show that the lower the critical mass for the clean, cold 38 condition the more sensitive is the reactor fo the accumulation of fission-product poisons and to fuel burnup. Reflector-Moderated Reactor The reflector-moderated reactor presents a num- ber of important advantages. By removing most of the moderator from the core to the reflector, the effective power density in the core can be nearly doubled for a given average power density in the fuel region. By heavily lumping the fuel, it is possible to eliminate much of the parasitic struc- tural material ordinarily required to separate the moderator and fue! regions, |f beryllium is em- ployed as the reflecter-moderotor, a substantial proportion of the neutrons are reflected back into the fuel region at epithermal energies so that they penetrate even fairly thick layers of fuel and keep the ratio of the peok-to-average fission density from exceeding something of the order of 1.5 to 2.0, Many factors influence the critical mass of the reflector-moderated reactor. Perhaps the most important is the poison concentration in the re- Other factors include the core radius and annulus thickness., Figure 23 shows critical mass plotted against these last two factors sl ORNL-LR-DWG., 2017 flector, the fuel EXTRAPOLATED REFLECTOR @ 49 THICKNESS = 30 CM NV ~ . | = | Y 30 | § 0 ’ g ] EQQ o - R; > 1@0\ FS R 0| EXTERNAL FUFL VOLUME =0 FT2 Fig. 23. Effect of Reactor Dimensions on Con- centration of Uranium in o NaZrF, Fuel for g Reflector-Moderated Circulating-Fuel Reactor, for a 30-cm-thick reflector containing an amount of poison representative of that which would be involved if canning of the beryllium with inconel should prove necessary. If the caonning is not required, the critical mass will be reduced by about 30%. A quite complete set of multigroup calculations (from which Fig. 23 was taken) is being made to determine these effects,33 REACTOR CONTROL3* The problem of reactor control is essentially one of matching the power of the reactor to the Joad.3? This uswally amounts to keeping the fuel elements at a prescribed temperature and moking absolutely certain that they do not go above a maximum temperature considered to be the threshold for damage., The ease with which this control can be accomplished is associated with the temperature coefficient of reactivity, It has been demonstrated that o reactor with a large negative temperature coefficient in the fuel does not even require a control rod; slow-acting shim rods for shufting down the reactor, for compensating for long-term drifts in reactivity, or for changing the operating temperature may be incorporated in some instances. A large negative femperature coefficient in the fuel has been achieved only in liquid-fuel reactors, such as the “water-boiler’’ and the Homogeneous Reactor Experiment. The Aircraft Reactor Experi- ment and the reflector-moderated reactor should also exhibit the demonstrated stability of other liquid-fuel reactors. Consequently, the control of the ARE oand the proposed reactor should be simple. |t moy well be that no nuclear instrumen- tation will be required for the circulating-fuel aircraft reactor; the proposed Homogeneous Test Reactor (HRT) is not to have mechanical control rods. ' ' | Reactors with solid fuel elements do not exhibit a large negative temperature coefficient in the fuel, although they may have o small over-all nega- tive temperature coefficient as a result of ex- pansion of the moderator or the coolant. |t is not 33(:. S5, Burtnette, M. E. LaVerne, and C. B. Mills, Reflector-ModeratedrReactor Design Parometer Study: Part [. Effects of Reacfor Proportions, ORNL CF-54-7-5 {(to be issued). #4This material was prepared with the assistance of W. H. Jordan, E. 5. Bettis, and E. R. Mann, 351nferim Repart of the ANFP Controi Board for the Aircroft Nuclear Propulsion Program, ANP-54 (Nov. 1950). ' implied that such recctors cannof be controlled or that they are even inordinately difficult to control; nevertheless, they do involve control problems that do not occur in the circulating-fuel reactors. The problems are summarized in the following statements, 1. A solid-fuel reactor must have a large number of shim rods to override xenon and to compensate for fuel depletion. This entails much mechanical gadgetry, as well as distortion of the flux pattern. Distortion of the flux pattern, in turn, makes the already difficult problem of hot spots much worse, By contrast, the liquid-fuel reactor does not have these problems because fuel can be added to give a uniformly higher fuel cencentration to take core of depletion, and xenon may be removed as it is formed. _ 2. While flux-sensing elements may be desirable in a liguid-fuel reactor, they are so vital in o solid-fuel reactor that they must be compounded, 3. Most of the proposed control systems for solid-fuel-efement reactors include a fast-acting servo-confrolled rod to compensate for quick changes in reactivity, Such o rod is a hozard in itself, since it might introduce a sharp increase in reactivity. Probably the only satisfactory sclution to this problem is to try to design the reactor so that abrupt increases in reactivity cannot occur, Even though step changes in reactivity are not anticipated, they afford a useful basis for analysis because a step change of the proper magnitude can be introduced into an analogous system to simulate most perturbations of practical interest. Thus, the controllability of a reactor can be deduced from the rise in fuel temperature that would result from a step change in reactivity, This is particulorly important in aircraft reactors where the operating temperature is made as close as possible to that likely to damage the reactor. When the response of a reactor with a negative temperature coefficient of reactivity (a) in the fuel is considered, it can be readily shown that the maximum temperature rise in the fuel (AT) as a result of a step change in reactivity 5&/k is given by Ok 2 — k AT = o Thus, if a =5 x 107%/°C and Sk/k = 3 x 1073, AT will be 120°C or 216°F. | 39 The temperature rise to be expected in a solid- fuel reactor can be approximated if it is ossumed (1) that the heat capacity of the fuel element is small, (2) that an increase in power produces a corresponding increase in the film drop between the fuel element and the cooclant, and (3) that the increase in power caused by a step change in reactivity is the transient term only, further in- creases being stopped by a servo control, In this case the power increase is given by 8k/&(3. Then , Bk (Qf ~ 0) = (1 +—16E> (9}, -~ 6) , where 0, = fuel temperature before the step change, 9;' = fuel temperature after the step change, G = coolant temperature, delayed neutron fraction. It can be seen that the increase in fuel tempera- ture over coolant temperature depends upon the original difference between fif and 0 _; for example, = I the ratio 07 ~ 0, Sk - 0= +—0= 1.4 6/ — 6.: kP for a Sk/k of 3 x 1073, The power and temperature perturbations for a sodium-cooled solid-fuel-element reactor were cal- culated on the ORNL reactor simulator according to the following conditions: 1. The volumetric heat capacity of the solid fue! elements was 1.0 cal/cm?® °C. 2. The fuel-region power density at design point was 5.7 kw/cm3, 3. The coolant was a liquid with thermal proper- ties comparable to those of liquid sodium. 4, The design-point power was 200 Mw and the coolant system was designed to extract power at this rate. 5. The step perturbation, Ak/k, was 0.305%. A servo system of reasonable proportions was simulated, and it was presumed that the power would be controlled from an error signal propoition- al to (p — p,), where p was the power at any time t and p, was the design-point power. The transient responses in power and in fuel temperature are shown in Fig, 24, where it is clear that the temper- ature rise depends on the original difference in temperature between fuel element and coolant, Thus in a loosely coupled system, such as an air- 40 UNCLASSIFIED ORNL-LR-DWG 2241 2600 : : , 2400 — 2200 e w [ve o % 2000 o Ly o = Lt l_ 1800 | w -l L = < 1600 + L = 1400 1200 b wmreeeeee e g z 300 ; | ! , o | wl i ; ! § 200 | — f | i | L ,| 0 5 10 15 20 25 TIME (sec) Fig. 24. Power and Temperature Overshocts for Step Changes in Reactivity of 0.305 Ak/k for a Sodivm-Cooled Solid-Fuel-Element Reactor with an Average Power Density in the Fuel of 5.7 kw/em>. cooled reactor in which the fuel element tempera- ture must be much higher than the coolant tempera- ture, the temperature rise would be much more severe than in the sodium-cooled reactor. MATERIALS The various moderoting materials that might be employed were discussed in the previous section because nuclear considerctions are dominant in their selection. This section covers structural, fuel element, and coolant materials, the selection of which is usually based mainly on engineering considerations. Structure . A key factor in the design of a reactor is the structural material of which it is to be built, An indication of the structural materials that might be employed in a high-temperature nuclear power plant may be gained from an examination of the program carried on during the past 15 vears for the development of superior materials for gas- turbine buckets, The most frequently used re- fractory alloys have been those of iron, chromium, and nickel, particularly the 18-8 stainless steels and Inconel. A group of alloys that give even better high-temperature performunce are cobalt- base alloys containing various amounts of iron, chromium, nickel, molybdenum, and tungsten. Un- fortunately, cobalt has o high neutron-absorption cress section and becomes an exceptionaily bad source of gammas it exposed fo thermal neutrons. if even trace amounts of cobalt were carried out- side the shield in a fluid circuit they would be serious sources of radiation. . Both ceramic ma- terials and cermets have also been employed, but their brittleness haos led to difficulties; they have yet to be developed to the point where they are capable of withstonding the severe thermal stresses imposed in turbojet engines. All the materials mentioned above were con- sidered because of their oxidation resistance, However, in certgin types of reactor it would be possible to employ refractory materials such as molybdenum, columbium, and graphite in an ombient completely free of oxygen, for example, a molten metal. Further, it is conceivable that a completely new refractory alloy might be developed from such high-melting-point materials as molybdenum, tung- sten, columbium, zirconium, chromium, and vana- dium. The recent development of iron-aluminum- molybdenum alloys, such as Theromatfor, lends credence fo this possibility. A particular system must be examined in order to evaluate the relative merits of the various structural materials, but, in general, the structural metal should have both high creep strength at high temperatures and ductility throughout the operating temperature range of at least 2 or 3% so that high focal thermal stresses will be relieved by plastic flow without cracking. It further seems necessary in most instances that the structural metal be highly impermeable and weldable, with ductility in the weld zone of at least 2 or 3% throughout the temperature range from the melting point to room temperature. Some of the materials that have been considered for use in aircraft nuclear power plants are listed in Table 8, together with their significant properties for this application., The availability of the material is o most important consideration in the conduct of o development program, because a good assortment of bar stock, tubing, and sheet is essenfial to the fabrication of test rigs, It has been primarily the availability consideration that has led to the use of iron-chrome-nicke! alloys in most of the development work to date. It is hoped, however, that better materials will be available for future, more advanced reactors. The effects of temperature on the stress-rupture properties and the creep rates of some typical metals and alloys are shown in Figs. 25 through UNCLASSIFIED ORNL-LR~DWS 2010 100 e Sty T . f ‘ T 8% af 1800°F 90 o o ] | f . i 80 . e e ol e INCONEL -~ ] o ¢ ELONGATION - HASTELLOY © (AS CAST) f HASTELLOY B J N ' { \ * | P ! i R ) ;/{@LJBD_ENUM RO e 1200 1200 1400 TEMPERATURE (°F) Fig, 25. Effect of Temperature on Elongation of Yorious Materials. 28.3% Unfortunately, these curves do not tell the whole story., Al the iron-chrome-nicke! alloys over-age at temperatures above 1650°F, because the hardening constituents, such as the carbides, tend to migrate to the grain boundaries. Annealing and grain growth inevitably accompany over-aging. Intergranular corrosion would be likely to follow and would probably cause trouble in thin sections where a grain might extend all the way through a 0.010- to 0.20-in.-thick sheet or tube wall. Solid Fue! Eiements While the bulk of the ORNL-ANP effort since the fall of 1951 has been directed toward the develop- ment of a circulating-fluoride-fuel reactor, the 3‘SJ. M. Woods, Mechanical Properties of Metals and Alloys at High Temperatures, ORNL-1754 (to be issued). UNCLASSIFIED ORNL-LR-DWG 2011 20 w RUPTURE STRENGTH {psi x 10"3) FOR 1000 nr 3 L e g oo b e e e e e o e e o o T WIRE-SPAGED PLATE-TYPE FUEL ELEMENT N T S L 1200 1400 1600 1800 2000 TEMPERATURE (°F) Fig. 26. Effect of Temperature on Rupture Strength of Yaricus Materials. 42 major effort prior to that time was on the develop- ment of reactors utilizing stationary fuel elements, The work on solid fuel elements is continuing, but on a very limited basis, so that another avenue of approach to the high-temperature aircraft reactor may be kept open. The fissionable material for a high-power reactor with stationary fuel elements may, in general, take the form of uranium metal, uranium metal alloy, uo,, UC,, or, possibly, other uranium compounds. However, at high temperatures serious difficulties are encountered because of the low melting point of uronium metal and most of its alloys.?7 The melting point of pure uronium is about 2066°F, that is, not much above the proposed fue! surface temperature for the reactor, so that the metal would be so weak at operating temperature as to require additional support, The support might be provided 37R. W. Bussard and H. E. Cleaves, Journal of Metal- lurgy and Ceramics, Yol. 1, No, 1 (1948), UNCLASSIFIED ORNL-LR -DWG 2012 25 :_"'"""“—h”’ ““‘“*T—_"""" "*—_N——‘”'I : | ! i | ! | I \ v 66 AT 1200°F | i i i il < MOLYBDENUM (10Q hr) | | | N ‘ ~+-TYPE 316 ‘ N ! 10 | - \-— STAINLESS STEEL | N i {1000 hr) ! *X«“ STRESS TO PRODUCE 1.0% CREEP {psi X 107%) WIRE ~SPACED LEZTE— TYPE FUEL ELEMENT | 0 Lb——d | | { 200 1300 1400 1500 1600 1700 1800 1300 TEMPERATURE (°F) Fig, 27. EHect of Temperature on Siress to Produce 1.0% Creep for Yarious Materiols, TABLE 8, PROPERTIES OF REFRACTORY METALS AND ALLOYS THERMAL.NEUTRON COEFFICIENT OF MELTING | DENSITY ABSORPTION MODULUS OF 1 e CMAL EXPANSION | SPECIFIC HEAT THERMAL cosT POINT NEAR 20°C CROSS SECTION ELASTICITY bER OF (cal/a/%0) CONDUCTIVITY AT 70°F {WELDABILITY | AVAILABILITY 5716} 2 0 ; o} (g/ cm> i callg Btu/hr 12 (°F /1) ( ¢/cm) {barns/ atom) (psi) (in./in.) o Tungsten 6170 19.3 19.2 52 x 108 2.4 x 10~9 0.032 96 Poor Poor 10.43 (ingot) Tantalum 5425 16.6 21.3 27 %108 3.6 x 1079 0.036 31 Good Fair 39 (sheet) Molybdenum 4760 i0.2 2.4 48 % 10° 2.7 x 10~ 0.061 85 Poor Poor 4 {pressed ingot) Niobium 4380 8.57 1.1 18 x 108 4.0 x 10~°% 0.065 Good Fair 75 {powder) Vanadium 3150 5.1 4.7 21.5 x 106 4.3 % 10_6 0.15 17 No data Poor 30 Zirconjum 3200 6.5 G.18 11 x ]06 3.0 x 10‘-6 0.08 14 Fair Fair 35 Titanium 3300 4.54 5.6 15 x10° 4.7 x 108 0.13 100 Fair Fair 15 (sheet) Chromium 3430 7.19 2.9 3.4 x 1076 011 39 Bad Difficult 3.60 (electrolytic) Iron 2802 7.87 2.43 29 108 6.5 % 10~9 0.11 36 Good Good 0.11 to 1.48 Cobalt 2723 8.9 34.8 30 x 10° 6.8 x10~% 0.099 40 Poor Difficult 2.60 Nicke! 2650 3.90 4.5 30 x10° 7.4 x10~% 0.105 34 Good Good 0.865 (sheet) Nichrome V 2550 8.4 30 x 108 9.8 x 10~ 0.107 7.8 Good Fair 1.00 Inconel 2600 8.51 4.0 31 x 108 6.4 % 108 0.11 8.7 Good Good 0.925 (sheet) Incone! X 2600 8.3 4.0 31 x 108 16.0 x 10~% 0.13 8.5 Fair Good 2.75 Type 316 stainless steel 2550 8.02 2.9 28 x 106 9.7 x 10~ 0 G.12 9 Good Good 0.645 (sheet) Hastelloy B 2900 9.24 3.9 30.7 % 10° 5.6 x 1079 0.091 6.5 Fair Fair 2.50 Zircaioy-2 (1.44% Sn; 6.55 0.25 13.8 x 108 6.5 x10~% 0.08 8.2 Foir Fair 35 0.05 Ni; 0.12 Fe; 0.11 Cr} 43 UNCLASSIFIED ORNL -LR-DWG 2C13 T e FATIGUE STREMGTH {psi x 1073 10 Jrmeremesmnen — o 1200 1400 1600 1300 TEMPERATURE (°F) Fig. 28. Effect of Temperature on Fatigue Strength of Various Materials, by using rods of solid moderator (such as BeQ) cooted with uranium metal, but the uranium coating would be very thin and would be very likely to break up and spall off because of thermal stresses. Even worse, .the metallic uranium would migrate by diffusion and mass transfer in the coolant to the walls of the pressure shell, heat exchanger tubes, ete., where it would tend to diffuse into the base metal and form o low-melting-point eutectic in the grain boundaries. The eutectic of iron and vranium melts at 1337°F, and the eutectics of uranium with nickel and chromium melt at tempera- atures well below 1800°F. alloy mefts at 2345°F over most of the composition Uranium-molybdenum range, and thus the presence of uranium metal would seriously molybdenum, About the only metal that does not form a low- melting-point eutectic with uranium is columbium, but this material is expensive and difficult to reduce the strength of procure and fabricate. Cladding or canning metal- lic uranium or its alloys would serve to reduce the diffusion rate but would not reduce it suffi- ciently ot the operating temperatures involved, The difficulties associated with the low melting points of uranium metal and uranium alloys, can be avoided by introducing the uranium as uao,, which is a chemically stable material with a very high melting point, 3949°F, Uranium carbide might also be used, but it is less stable chemicaolly, and it would react with most moderators, coclonts, or canning materials at the temperatures considered here. Therefore, UC, seems quite inferior to UQ,, except, perhaps, on the basis of thermal conduc- tivity and resistance to thermal shock. Other uranium compounds have been considered, but none appears to be superior to UG, ; Uranium oxide can be fabricated into fuel eles ments in a number of ways. For an air- or helium- cooled reoctor it might be contoined in o matrix of chromium and Al;0,, in the form of o ceramel, or in a matrix of silicon carbide, in the form of a ceramic. Since o large surface area is essential, the fuel elements could be in the form of thin flat plates or tubes or a pebble bed, The support of such fuel elements would be difficult if they were to be used at high temperatures. [ thin plates or tubes of o ceramic or a ceramel were held rigidly, they would be virtually certain to crack under thermal stress; if they were supported loosely, they would flutter in the high-velocity gas stream and fail as a result of abrasion of the contact surfaces, If o pebble bed were used, the same difficulties would be encountered, In any case the support structure would have to be metai, While the metal could be cooled, it is hard to see how hot spots could be prevented if the ceramic or the ceramel were af operating temperatures much above the temperature of the metal, The UO, might be poured Ioosely into long siender metal tubes or pins, but at high power densities the temperature at the cenfer of even 0.080-in.-1D pins would exceed the melting point 45 of the UO, and cause fusion. Hence, there would probably be objectionable concentrations of UO, that would create hot spots at indetferminate regions in the pin. A better arrangement would appear to be to place a thin layer of UO, on the inside of the tube wall, as proposed in the KAPL-SIR de- sign.>® The problem of supporting and accurately spacing these pins or tubes is a most serious one, however, as has been clearly shown by experience at KAPL. Probably the most promising way to fabricate U0, into a fuel element is to clad with stainless steel o sintered compact of UO, and stainless steel3? in which the U0, may constitute as much as one-third of the volume. Sandwiches of this type can be rolled to give plates with minimum thicknesses of 0.006 in. of cladding and 0.008 in. of UQ. compact in the core. One obvious way to use such fuel plates would be to stack alternate flat and corrugated plates to give the grrgngement shown in Fig. 29. The coolant would flow between the corrugations. This arrangement has the dis- advantage, particularly when vsed with low-thermal- conductivity coolants, of giving hot spots in the low-velocity regions in the vicinity of the points of contact between the flat and corrugated plates, Short spacers containing no fuel can be placed between the corrugated and flot shests to avoid this, as in the arrangement shown in Fig. 30. In a third arrangement, shown in Fig. 31, wire spacers are passed perpendicularly through flat plates at intervals sufficiently close to maintain good spacing in spite of tendencies toward thermal distortion. A fourth arrangement, shown in Fig. 32, is based on the demonstrated procticality of fabricating the UO, stainless compact in the form of tubes. This arrangement gives o fue! element that is very resistont to warping and thermal distortion, Yet another arrangement, shown in Fig. 33, depends on the use of UQ, packed into small-diameter tubes which can be drawn or swaged to give wires as small as 0.020 in. in diameter, Ancther arrangement, shown in Fig. 34, employs 38K nolls Atomic Power Laboratory, Reactor Engi- neering Progress Report July, August, September, 1951, KAPL-614, p. 13, 39G. M. Adamson, ANP Quar. Prog. Rep. June 10, 1951, ANP-65, p. 181; E. 5. Bomar and J. H. Coabs, ANP Quar, Prog. Rep. Sepf. 10, 7957, ORNL-1154, p. 147; E. S. Bomar and J. H. Coobs, ANP Quar,.Prog. Rep, Dec, 10, 1951, ORNL.-1170, p. 128; E. 5. Bomar, J. H. Coobs, and H. Inouye, Met. Div. Semiann. Apr. 10, 1953, ORNL-1551, p. 58. 46 e ORN_-LR-CW6 1441 COOLANT FLOW Fig, 29. Corrugated Plate Type of Fuel Eiement. NON FUEL - BEARING SPACER BARS N Tl = ‘Jr; T /'4,;,\\’\__/;?_ S o o B T e, ,,". B At j ] B g gy COOLANT FLOW Fig. 30. Corrugated Plate Type of Fuel Element with Nonfuel-Bearing Spacer Bars, sintered blocks of UQ, and stainless steel compact in which a closely spaced hole pattern would provide coolont flow possages and heat transfer surface area, If erosion or spalling should prove a problem with this arrangement, the holes might be lined with thin-walled tubes and the gap be- tween the blocks and the tube walls might be filled with a molten metal, such as sodium, to provide @ good thermal bond. Some idea of the amount of core wvolume that must be devoted to the fuel elements can be gained from an illustrative example. If the critical mass for a reactor were 50 Ib of UZ3° and a sintered stainless steel matrix containing 33 vol % UO, were employed, the ceramel matrix volume would have to be about 0.25 ft3. The volume of cladding HPNL—I!'!WG a3 . m.’ ORNL -LR- Biv5 1145 —7 WIRE SPACERS e . 71\ STAINLERS SYEFL CLADDING — - \SINTEHED LG, AND STAINLESS STEEL CORE Fig. 33. Wire Type of Fuel Element, SECRET ~ Ol ~[ R -[CWG {148 e i : AT S . P o . COOLANT FLOW \:\\W P o “ Fig. 31. Wire-Spaced Plate Type of Fuel Element, ~. SELRET ORNI - LR~DW5 1144 < STAINLESS STEEL % %.CLADDING G e Fig. 34. Sintered UO, and Stainless Steel Block Type of Fuel Element, 3 fn summary, a good detoil design for a solid-fuei- element system for an aircraft reactor should pro- vide the following: < SINTERED UG, AND STAINLESS STEEL CORE _ . ' : 1. an adequate volume of UD, to insure criticality, Fig. 32. Sandwich-Tube Type of Fuel Element. 2. adequate surface area to meet heat transfer requirements, material required to provide adequate surface arec 3. a surface that would not give trouble with cor- to meet heat trensfer requirements usually proves rosion, mass transfer, erosion, or spalling, or to be about the same as the matrix volume, and have a tendency to pick up surface films that therefore the volume of material in the fuel ele- would impede heat transfer, ments would be about 0.5 ft3, This would consti- 4. a geometry that would give a fairly uniform tute 12% of ‘the volume of a 2-ft-dia spherical temperature distribution throughout the fuel reactor core. . element and avoid both excessive temperatures 47 in the interior and thermal stresses that would induce cracking or warping under power- and temperature-cycling conditions, 5. adequate strength and stiffness to insure struc- tural integrity and the surface spacing required by heat transfer considerations so that hot spots could be avoided, 6. a fuel element that could be consistently fabri- cated with the requisite quality at reasonable cost, 7. structural material in an amount consistent with a reasonable critical mass requirement, High-Temperature Liquid Coclants and Fuel-Carriers A thorough survey of materials that appear prom- ising as heat transfer fluids for high-temperature aircroft reactors was presented in ORNL-360.40 The first requirement is that the fluid must be liquid and thermally stable over the temperature range from 1000 to 1800°F. A melting point con- siderably below 1000°F would be preferable for ease in handling, while a substance that would be liquid at room temperature would be even better, Other desirable characteristics are low neutron absorption, high volumetric specific heat, and high thermal conductivity. Above all, it must be pos- sible to contain the liquid in a good structural material at high temperatures without serious corrosion or mass transfer of the structural ma- terial, The principal substances so far suggested that show much promise of satisfying these re- quirements are listed in Table 9, together with Of these ma- terials, sodium hydroxide, lead, and bismuth are some of their physical properties. considered to be only marginally useful because of their corrosion and mass transfer characteristics, The promising liquid metals can be separated into a light group, lithium and sodium, and a heavy group, lead and bismuth, As will be discussed in the next section, the light metals are highly preferred because of their superior heat transfer properties and corrosion and mass transfer charac- teristics. Liquids intended to serve as vehicles for uranium in circulating-fuel reactors that operate at high temperatures are subject to the same criteria as those that serve as coolants. In addition, the 4OA. S. Kitzes, A Discussion of lLiquid Metals as Pile Coolants, ORNL=-360 (Aug. 10, 1949). 48 solubility of uranium and its effect on the physical properties of the fluid must be considered. All the fluids in Table 9, except the fluorides, can be shown to be unsuitable, for one reason or another, as vehicles for uranium. Fortunately there are many different fluorides that can be used, 4! The NafF-ZrF, melt (NaZrf;) was chosen for the ARE because the materials were readily available, nontoxic, and not too expensive. Unfortunately, the physical properties of this fluoride mixture, particularly the melting point, vapor pressure, and viscosity, leave much to be desired. Both BeF, and LiF can be used to reduce the melting point, but BeF, is toxic and LiF would require Li7, a material that is not avoilable, although it could be obtained at a price that should not be unreason- able. Other promising components are KF and RbLF; however, KF has a neutron absorption cross section that is higher than is desirable (Tabie 7), while RbF commercial demand for it, is expencive because there is no It has been determined that ample stocks of rubidium-contcining ore are available, and the price of Rbi should not be un- reascnable if substantial amounts are ordered. The terms corrosion and mass transfer need some clurification. Corrosion implies the remaval of surface material from the container by a chemical reaction with the liquid or by simple solution in the liquid, as is the case with liquid metals, Cor- rosion damage to a solid material caused by contact with a fluid results in a loss in strength of the solid material, Mass transfer in liquid metals is a phenomenon that involves removal of container material from the hotter portion and deposition in the cooler zane of a closed circuit with a tempera- ture gradient in which the liquid is being circu- lated. variations in solubility as a function of tempera- The removal and deposition result from However, when the circulating fluid is a fused salt, the container material is transported ture, from the hotter to the cooler zone of the circuit because of variations in the equilibrium constanis of the chemical reactions as functions of tempera- ture. tor example, in an Inconel system circu- lating a fused-fluoride-salt fuel, the differences in chemical equilibria at the two temperature zones may cause mass transfer of chromium according to 41w, R, Grimes and D. G. Hitl, High-Temperoture Fuel Systems, a Literuiure Survey, Y-657 (July 20, 1950); The Reactor Handbook, Vol. 2, Sec. 6, p. 9215 (1953). TABLE 9. PROPERTIES OF REPRESENTATIVE REACTOR COOLANTS MACROSCOPIC BOILING MELTING POINT, THERMAL VISCOSITY SPECIFIC HEAT DENSITY VOLUMETRIC THERMAL-NEUTRCN PREFERRED COOLANT POINT 7 CONDUCTIVITY o /o3 HEAT CAPACITY ABSORPTION CONTAINING REMARKS 5 60 mm 2 o (cp) {cal/g/"C) (g/em™) o 3 (°F) (°F) [Btu/hr-#t°-(OF /)] (cal/°Crem®) CROSS SECTION MATERIAL (em™ ) Li’ 354.0 2403 25.0 0.4 1.0 0.46 0.46 £.00131 Type 430 stainless Severe mass transfer above 1150°F steel Na 208.0 1621 34.5 0.2 0.30 0.78 0.23 0.0092 Type 316 stoinless Virtuatly no corrosien or mass transfer up to 1600°F steel or Incone! NaK (56% Na, 66.2 1518 16.7 0.161 0.253 0.742 0.188 0.0183 Type 316 stainless Virtually no corrosion or mass transfer up to 1600°F 44% K) steel or Inconel Pb 621.0 3159 8.6 1.2 0.037 10.0 0.37 0.00592 Type 430 stainless Severe corrosion* and mass transfer above 1150°F ‘ steel Bi 520.0 2691 9.0 1.0 0.039 9.4 0.367 6.000406 Type 430 stainless Severe corrosion and mass transfer above 1150°F steel NaQOH 0.7 1.0 0.49 1.7 0.83 0.021 Nickel Severe corrosion and mass transfer above 1150°F NerF5 950.0 2.0 7.5 0.29 3.0 0.87 0.00367 Inconel No corrosion or mass transfer up to 1500°F when used as a vehicle for UF3 NaoF-KF-LiF 851.0 2.5 2.5 0.40 1.9 0.76 tnconel Prabably no corrosion or mass transfer up to 1500°F when used as a vehicle for UF3 HZO {100°F) 32.0 212 0.35 0.7 1.0 1.0 1.0 0.048 Type 347 stainless Severe corrosion above 1650°F** steel Air (sea level) 0.04 0.04 0.26 0.00032 0.00008 0.00002 Type 310 stainless Severe corrosion above 1800°F steel or Nichrome V *The term '‘severe corrosion’’ is used where the attack exceeds a depth of 0.010 in. after 500 hr of testing, because in most reactors the thickness of many structurcl elements must be less than 0.025 in, **G, H, Hawkins et al., Trans. Am. Soc. Mech., Engrs. 65, 301 {1943). 49 the reaction Cr + 2UF, =2UF, + CrF, . Another set of reactions, known as dissimilar metal transfer, makes it desirable that complex plumbing systems be fdbricated entirely from one metal or alloy. Dissimilar metal transfer involves the re- moval of one or more of the constituents of one alloy and its transport through the liquid to another alloy where the deposited material diffuses into the base metal. The transport driving force in this case is a difference in chemical potential; the chemical potential of a constituent of a complex alloy is lower than that of a pure metal or of a simple-solution alloy. Examples of dissimilar metal transfer have included the plugging of nickel heat exchanger tubing by iron which was trans- ported to the nickel surface through the liguid medium from a stainless steel pump chamber and a stainless steel expansion tank. HEAT REMOVAL The power density in the reactor core is limited by the rate at which heat can be removed by the fluid passing through the core. rate depends, in turn, on the permissible fluid The heat removal velocity and the temperature rise through the re- actor and on the density and specific heat of the heat transfer fluid, The optimum coolant tempera- ture rise depends upon the characteristics and proportions of the over-all power plant, While the relations -are quite complex and depend in large measure upon the characteristics of the various components of the system, for most aircraft reactor types the optimum temperature rise for the fluid passing through the reactor core appears to be of the order of 400°F.42 In fact, a temperature rise greater than 600°F has been proposed for only one of the detailed major cycle proposals made to date — the air cycle. For the air cycle the allow- able temperature rise will be the difference be- tween the maximum reactor air outlet temperature obtainable and the turbojet compressor outlet tem- perature. The resulting temperature rise is likely to be of the order of 600°F, depending on the com- pression ratio. Once a permissible femperature rise is estab- lished and a coolant is chosen, a major limiting 425 M. Walley, W. K. Moran, and W. Graff, Off Design Turbojet Engine Performance of a Nuclear Powered Aircraft, ORNL CF.53-9-80 (Aug. 1953). factor for solid-fuel-element reactors is the per- missible pressure drop across the reactor core, While there is some variation in the pressure drop associated with different types of fuel element and different reactor core arrangements, it appears, in general, that the pressure drop across the core should be kept to something of the order of 30 to 50 psi because of limitations imposed by pumping power and fuel-element stress considerations. A third important facter associated with heat removal from the core of a solid-fuel-element re- actor is the difference in temperature between the fuel element and the coolant. The higher the heat transfer coefficient obtainable, the lower this temperature difference becomes. In attempting the detailed design of any particular reactor, it soon becomes evident that, regardless of how desirable an increased amount of heat transfer surface area may be, the problems associated with the fabri- cation of the fuel elements become progressively greater as the amount of heat transfer surface area per unit of volume is increased and the structure becomes progressively more delicate and ““lacey.”’ In almost every instance the inclination is to de- crease the hydraviic radius of the coolant passage to a value as low as possible consistent with problems of fabricating the fuel-eiement surfaces and with stress considerations associated both with the fluid pressure drop and the thermal stresses that would produce thermal distortion. A third factor affecting the power density ob- tainable from a reactor core is the free-flow ratio, that is, the ratio of the effective flow-passage area to the total cross-sectional arec of the reactor core. For most reactors the moximum practical value for this parameter appears tc be about 0.40, but for the reflector-moderated high-temperature- liquid type it appears to be closer to 0,60, and for the circulating-moderator type it appears to be of the corder of 0.85. Water-moderated reactors in which water is not the prime heat transfer medium can be designed for free-flow ratios as high as 0.50 because water is so potent a moderator, {f an attempt is made to get the maximum power density from a given core matrix geometry with a given coolant, it soon becomes evident thot any actual reactor must be expected to differ consider- ably from the commonly assumed ideal reactor in which there is ¢ perfectly uniform fiuid fiow distri- bution and the core matrix is stressed by a perfectly vniferm loading. Careful consideration of the 51 usual perversities of velocity distribution under turbulent flow conditions disclosed marked devi- ations from ideal conditions. Also, the ideal con- ditions are clearly unreasonable if allowances are made for ordinary amounts of thermal distortion. Experience in brazing radiator core matrices, for excmple, has shown that even the relatively slow rates of temperature change associated with the furnace brazing operations produce variations in passage thickness of as much as 30%. Thus it is felt that even if great care is taken in the design to minimize the cumulative effects of thermal distortion and fabrication tolerances, variations in effective thickness of the coolant passage of the order of at least 20% will occur. Heat transfer analyses show that variations in coolant passage thickness would lead to the formation of hot spots, and thus in some regions the local temperature difference between the fuel-element surface and the coolant would be greater than the design value, Substantial variations in power density through- out the core matrix can be expected to result from nonuniform fission densities, since even in the ideal reactor, there would be variations in fission density because of the effects of geometry on the neutron flux. Allowances must be made in any actual reactor for additional irregularities caused by the presence of control rods and by the non- uniform distribution of the fission-product poisons that will accumulate, |t therefore appears that local power densities at least 50% greater than the mean power density must be expected. |If further allowance is made for vagaries in flow distribution and for irregularities in channel shape as a result of thermal distortion, it would seem that in a realistic design, local temperature dif- ferences hetween the fuel element and the coolant of at least twice the mean should be anticipated. A careful examination of the siress anclysis problem for any core matrix shows that the same basic reasoning must be applied as that applied to thermal distortion. Heat removal reguirements for the maximum available flow passage area and the maximum possible heat transfer area, coupled with nuclear requirements to minimize neutron absorption in structural material, have led to u relatively complex, finely divided structure in every design proposed to date. Since a complex structure is inherent in o solid-fuel-element reactor, the stresses induced in the fuel elements and their supports by the pressure drop acioss thes core 52 matrix always constitute o problem. The most probable cause of failure would be the fotigue stresses arising from the pressure fluctuations associated with the turbulent flow of the coolant through the core matrix, Just as in the blades in turhojet engines, the stresses induced would probably be two or three times the direct stresses indicated by the average pressure drop across the core matrix, As was pointed out in the previous section, the iron-chrome-nicks! alloys are the only structural materials from which it seems reasonable at this time to expect to fabricate fuel elements. The data available on the high-temperature strength of these alloys indicate that even if the siresses are kept low, the permissible operating temperature can scarcely exceed 1800°F. The strength proper- ties of Inconel and other possible structural ma- terials are presented in Figs. 25 through 28 as functions of temperature, Dotted lines on Figs. 26 and 27 show the stresses unticipated in one of the most favorable fuel element matrices devised to date, that is, the wire-spaced plate-type fue!l ele- ment shown in Fig. 31. A consideration of the safety factor that would be acceptable for so vital a structure as a reactor core indicates that the maximum allowable operating temperature of a solid fuel element would probably be between 1600 and 1800°F, Since the wire-spaced plate-type fuel element (Fig. 31) is representative of the possible solid fue! elements, it was used as a basis for com- paring the characteristics of various potential re- actor coolaonts. For this study it was assumed that the plates were 0.020 in. thick and spaced on 0.120-in. centers and that the wire spacers ob- structed 5% of the effective flow passage area, The operating conditions assumed were a fluid temperature rise of 400°F, a fluid pressure drop of 50 psi, and a limiting fuel-element-metal tempera- ture of 1700°F. The limiting fluid outlet tempera- ture was to be determined by hot-spot consider- ations; that is, the maximum permissikble tempera- ture differential between the fluid and the fuel If a higher limiting fuel element temperature were wused, a lower fluid pressure drop would sesm to be neces- sary. Table 10 shows the results of a set of calculations based on these assumptions. element was to be twice the mean. Air, as a coolant, was ireated as a special case. The limiting flow velocity for the air was de- TABLE 10. LIMITING POWER DENSITIES FOR VARIOUS REACTOR COOLANTS HEAT-TRANSFER-LIMITED ANT P VELOCITY HEAT REMOVED PER COOLANT-FLOW-LIMITED | HEAT TRANSFE EAN TEMP E | POWER DENSITY FOR A REACTOR COOLL:L A:SAGE FLUID DENSITY FOR SPECIFIC HEAT | UNIT OF PASSAGE | FREE-FLOW PowTE-; DEN';;YT D CO-II:Z:FICTE;TR M DllFEREiéEUR ocar TEMP;RAFTURE COOLANT G7 (g/cms) 50-psi Ap (col/g/OC) FLOW AREA RATIO 3 2 o o (in) 2 o (kw/ em3) (Btu/hr-#12:F) (°F) DIFFERENCE OF 100°F {ft/sec) {Btu/sec-ft°.°F) 3 {kw/em”) L7 20 0.46 68 1.0 1960 0.6 10.4 55,000 128 8.1 Na 20 0.78 53.5 0.30 782 0.6 4.17 39,300 71.6 5.7 NaK 20 0.742 55 0.25 636 0.6 3.4 24,100 95 3.6 Py 20 10.0 14.9 0.037 343 0.6 1.83 12,800 96.5 1.9 Bi 20 9.4 15.2 0.039 347 0.6 1.85 13,000 95.5 1.9 NaOH 30 1.7 29.5 0.49 1530 0.85 1.6 6,100 601 1.9 NaZrFg 30 3.0 22.2 0.29 1206 0.6 4.27 8,000 362 1.18 NaF-KF-LiF 30 1.87 28.7 0.4 1340 0.6 4.76 16,000 201 2.4 H,0 (100°F, no boiling) 30 1 38.8 1 2420 0.85 18.3 1,200 4840 0.38 Supercritical water™ 30 0.4 0.70 200* 0.70 Air (sea level) 40 0.0059 350+ 0.26 33.4 0.5 0.088 360 198 0.088 Air (45,000 f) 40 0.00134 322% 0.26 7.0 0.5 0.027 18 187 0.027 *Supercritical water calculations were based on heat transfer and pressure-drop data given in Pratt & Whitney Aircraft Div., Nuclear Propulsion Program Engineering Progress Reports, No, 9, PWAC-75 and No. 10, PWAC-83, The limiting temperaturs difference of 200°F was the temperature difference in the outlet region. **Mach .20 at inlet. 53 termined by a compressibility loss consideration; that is, the Mach number was controlling. In all other cases the fluid velocity was computed to give an ideal pressure drop across the fuel element of 30 psi, and an additional 20 psi was assigned to the spacers and supports for the fuel plates to give an over=all pressure drop of 50 psi. The temperature rise in the air also required special trectment. |t was taken as being equal to the difference between the compressor outlet tempera- ture given in APEX-943 and 1700°F minus twice the mean temperature difference between the fuel element surface and the air, It is evident from columns 9 and 10 of Table 10 that, except for air and the liquid metals, the principal limitation on reactor power density is the temperature differential between the fuel ele- menf and the coolant rather than the rate at which coolant can be forced through the fuel element matrix, Therefore the mean local temperature differential between the metal surface and the coolant was specified as 100°F so that the peak fuel-element-surface temperature would be 1700°F, the average fuel-element femperature would be 1600°F at the coolant-outlet face, and the average coolant outlet temperature would be 1500°F. The resulting heat-transfer-limited power densities are given in the last column, It can be seen that on a heat removal basis for a consistent set of con- ditions lithium is clearly the best reactor coolant and that sodium is a close second, while air is the poorest in that it requires a reactor core volume 50 times greater than that required by sodium for a given power output. A remarkable point is that the molten salts are actually superior to the heavy liquid metals as heat transfer me- diums. H. F. Poppendiek and M. W. Rosenthal are preparing a report covering o more sophisticated and complete analysis than that given in Table 10. In their work they also varied the hydraulic radius of the heat transfer passages; however, their work leads to essentially the same conclusions as those presented here, An important point for which no allowance was made in the above analysis is that the local heat transfer coefficient is much less sensitive to vagaries in the local fluid velocity for molten metals than for the other coolonts. This makes 435 eneral Electric Co., Aircraft Nuclear Propulsion, Department of Engineering, Progress Report No. 9, APE X9 (Sept. 1953). the molten metals definitely more desirable be- cause with them the likelihood of hot spots and thermal distortion would be reduced. Unfortunately, however, of the good heat fransfer mediums only sodium, NaK, and the molten fluorides can be used for periods of 100 hr or more at temperatures of around 1500°F in any structural material currently available and fabricable, TEMPERATURE GRADIENTS AND THERMAL STRESSES Thermal stresses have been referred to o number of times in previous sections. These stresses may be induced by a temperature difference between two fluid streams, as in the tube walls of a heat exchanger, or by a temperature difference between the surface and the core of a solid in which heat is being generated.?? FExamples of the latter are fission heating in solid fuel elements and gamma and neutron heating in solid moderator materials, A calculated thermal stress gives a good indication of the behavior of a brittle material; that is, cracking is likely to occur if the calculated thermal stress exceeds the normal tensile or shear strength of the matericl. Only a small amount of yielding is necessary in a ductile material, however, to relieve the thermal stress. Therefore the calcu- lated thermal stresses for ductile materials are significant only in that they indicote that if the elastic limit of the material is exceeded, plostic flow and, possibly, distortion will result. Progres- sively greater distortion may result from thermal cycling, This might lead, for example, to partial blocking of a flow passage between adjacent plates in a solid-fuel-element assembly. Thus thermal stresses in o fuel plate might lead to a hot spot and hence to burn-out of a fuel element, For most purposes, thermal stresses can be approximated by considering one of two ideal configurations, namely, flat slabs and thick-walled cylinders with uniformly distributed volume heat sources, Charts for the simpler flat-slab configu- ration are presented in Figs. 35 and 36 to show both the temperature ditference and the thermal stress between the surface and the core for the materials of greatest interest. The values given are for a uniformly distributed heat source giving ‘MF. A. Field, Temperature Gradients and Thermai Stresses in Heat-Generating Bodies, ORNL CF-54-5-196 (May 21, 1954). 55 ORNL~LR--DWG 1147 10 0020 0050 010 0.20 050 08 PLATE THICKNESS (in.) Fig. 35, Moximum Temperature Ditfereatial in Flat Plates with Uniformly Distributed Volume Heot Scurces. 56 STRESS {psi) Fig. 36. < ORNL —LR-DWG H4B 005 04 0.2 05 08 ' PLATE THICKNESS (in) Thermal Stress in Flat Plotes with Uniformly Distributed Yolume Heat Sources. 57 a power density of 1 watt/em®, Both the tempera- ture difference and the thermal stress are directly proportional to the power density. Many other geometries can eusily be reduced to the flateslab configuration. A flat plate with heat generated in a plane at the center will have twice the tempera- ture differential and one and one-half times the stress of a plate of equal thickness with uniform heat generation, A cylindrical rod with uniform heat generation will have one-half the temperature differential and three-eighths the stress of a flat plate with uniform heat generation and a thickness equal to the diameter of the rod. It is instructive to apply Figs. 35 and 36 to some typical structures, for example, the fuel plates for a sodium-cooled reactor with stainless-steel-clad UQ, and stainless steel fuel elements. By taking the solid fuel element design shown in Fig. 31, on which Table 10 was bosed, and a reactor core power density of 4,2 kw/cm3, the power density in the fuel element will be 35 kw/cm® because it constitutes only 12% of the total core volume., It can be seen from the curve for Inconel (Fig. 36), which has properties about the some as those for stainless steel, that the thermal stress for 0.020- in.-thick fue! plates would be 0.4 psi for 1 watt/cm?3, or 14,000 psi for 35 kw/cm®, if the fuel is uniformly distributed throughout the plates, I, instead, the cladding constitutes one-half the total thickness, it can be shown that the temperature differential and the thermal stresses are approxi- mately half again as great, and thus there would be a 120°F temperature difference between the center and the surface and a thermal stress of about 20,000 psi. By referring to Fig. 27, it can be seen that this therma! stress is mony times the creep strength of the stainless steel at a temperature of 1700°F ; therefore severe thermal distortion would be likely to result. Thus the actual power density might have to be substantially less than the 4.2 kw/cm® permitted by heat transfer considerations, The more complex geometry of the thick-walled cylinder requires a more complex representation.4? For the purposes of this report, the typical set of curves shown in Fig. 37 will suffice to show the basic relotionships, These curves apply to an important particular case; namely, a reactor moder- ator region cooled by equilaterally spaced circular passages, It can be shown that the temperature and the thermal stress distribution in a rigid block cooled by equilaterally spaced parallel circular 58 5 ORNL-LR-DWG 49 Z ~ T ] T —- VOL. % iN HOLES J e — CCOLING FOLE DIA = 0125in 108 A c _ 3 "’}#‘T:_:_;;;‘: POWER DENSITY = 200 watts/cm T m.:lll__:' Sl T I R ] | {»— e o T o — ,J T o T . ! : i ) - Be YELD POINT AT 1200°F ' T L] oTH o HOLE SPACING (in} Fig. 37. Effects of Heole Spacing and Diameter on Temperature Distribution and Thermal Stress in a Rigid Block Having a Usiformly Distributed Volume Heat Source and Equilaterally Spaced Circular Cecling Possages., passages can be closely approximated by con- sidering the block to be a stack of thick-walled cylinders having hole diameters the same as those in the block and an outside diameter equal to 105% of the hole spacing in the block. For the case shown, the power density in the reactor core was taken as approximately 4 kw/cm3, which gives gamma- and neutron-heating density in the moder- ator of about 200 watt/em® (that is, 5% of the power density in the core). The chart is equally applicable to a reactor core geometry similar to that of the ARE or to the regions in the reflector or the island immediately adjocent to the fuel region of the reflector-moderated reactor. It is quite evident that BeO, because of its brittleness, could be considered for use only if pierced with many closely-spaced cooling passages; however, such o structure would be flimsy and easily damaged. TEMPERATURE DISTRIBUTION IN CIRCULATING-FUEL REACTORS The circulating-fuel regctor poses some special temperature distribution problems that have not demanded attention in other fields of technology. These problems arise beccuse the temperature of any given element of fluid in the reactor core at any given instant is a complex function of the time that it has spent in the fissioning region, the power density, the amount of heat that it has gained from or lost to the rest of the fluid through conduction or turbulent mixing, and its own heat capacity. As a consequence, there are two major sets of problems that may arise in any circulating-fuel reactor, The first of these is the formation of severe local hot spots as a result of flow sepa- ration. If the hot spots caused local boiling in a reactor having a high power density, there would be erratic fluctuations in power and, possibly, instability of the reactor. Therefore it seems essential that the fuel flow passages be carefully proporfioned o ovoid flow separation. The second problem, boundary-layer heating, arises because fissioning in the nearly stagnant fluid at the fuel- ° channel surface makes the temperature there tend to be much higher than that of the free stream. A rigorous and comprehensive study of the boundary- layer phenomenon has been 'in process since 1952.4%:46 A few curves based on that study are presented here to show some of the more important relationships. ' A good insight into the problem can be gained from examination of an important typical case ~ that presented by an ARE type of right-circular- cylinder reactor core containing parallel circular passages proportioned so that 50 vol % of the core is filled with fuel while the remainder of the core is moderator and structural material, If the pas- sage wall between the moderator and the fuel were not cooled, the wall temperature would exceed the local amount, as can be seen in Fig. 38. At a given fuel velocity the temperature difference between the wall and the fuel is directly proportional to power mean fuel temperature by a substantial density, but, if the reasoning of the previous section is followed and the fluid temperature rise is kept constant at 400°F for a given reactor core, the fuel velocity becomes directly propottional to the power density. Surprisingly encugh, the re- duction in boundary-layer thickness associated 45H. F. Poppendiek and L. D. Palmer, Forced Cone vecfion Heat Transfer Between Porallel Plates and in Annuli with Volume Heat Sources Within the Fluids, ORML-T701 {May 11, 1954). ' 464, F. Poppendiek and L. D. Palmer, Forced Cons vection Heat Transfer in Pipes with Volume Heof Sources Within the Fluids, ORNL-1395 (Dec. 2, 1952). ORNL-LR-DWG 2014 ‘ { i ‘ i 500 x’ PASSAGE DIAMETER, 8in. — 1 / ‘ | 400 |- T N 200 100 UNCOOLED WwaALL TEMPERATURE LESS MIXED-MEAN FUEL TEMPERATURE (°F) 0.5 1.0 1.5 2.0 25 30 35 40 FOWER DENSITY IN FUEL (kw/ecm®) Fig. 38. Effect of Passage Diameter and Power Density on the Difference Between the Uncooled Wall Temperature and the Mixed-Mean Fuel Tem- perature for an ARE Type of Reactor Core, A 21- in.~dia right circular-cylinder reactor core con- taining 50 vol % of NaF-ZrF4—UF4 as the fuel was dssumed., 59 with the increase in the Reynolds number causes the temperature difference between the fuel and the wall to diop somewhat with an increase in power density, This effect is shown in Fig. 38, together with the effects of variations in passage diameter. The smaller passoges give markedly reduced wall temperatures. Unfortunately, reducing the diameter of the passage also increases the amount of structural material in the reactor and, hence, increases the critical mass. A similar set of data is presented in Fig. 39 for a 2l-in.-dia fuel onnulus that is typical of reflector-moderated recctors. The effects of variations in fuel physi- cal properties are indicated by curves for two different fluoride welts having respectively about as good and as poor sets of heat transfer proper- ties as are likely to prove of practical interest, It is clear from Figs. 38 and 39 that, since the temperature of the structural metal wall is the limiting temperature in the system, there is a sirong incentive to cool the walls, The temperature distribution through the fuel stream, the wall, and the wa!l coolant for several conditions is shown in Fig. 40 for a 4-in.~-dia fuel tube, a ]/é-in.-'rhick Incone! wall, ond a 1/3-in.«dic wall coolant channel, Sodivm was assumed as the wall coolant, and the fuel assumed had physical properties similar to those of NaK-KF-LiF-UF, (10,9-43,5-44.5-1.1 mole %). Similar curves are given in Fig. 41 for a fuel 1800 —— - ———————— LESS TEMPERATURE TEMPERATURE (°F) UNCOOLED WALL MIXED-MEAN FUEL 05 1.0 15 20 25 30 35 4.0 POWER DENSITY IN THE FUEL fkw/cm?) Fig. 39. Effects of Power Density in Twe Dif- ferent Fuels on the Difference Between the Un- cooled Wall Tempercture and the Mixed-Mecan Fuel 21.in,~dia Core Temperature for o Reflector- Moderatad Reactor. 60 having the same physical properties except vis- cosity, which was assumed to be ten times greater than that of the fuel assumed for Fig. 40. The difference in temperaiure betwesen the center of the stream and the peak fuel temperature is nearly twice as great as that for the lower viscosity fuel, The sensitivity of the system to velocity distri- bution in either the fue! or the wall coolant fluid streams is shown in Figs. 42 and 43. The curves in Fig. 42 are for sodium-cccled walls, while those presented in [ig, 43 are for the same system except that NaOH is used as the wall coolant, An examination of these two sets of curves shows that moderate variations in fuel velocity have relatively little effect on the temperature distri- bution. Further, from the temperature distribution standpoint, the NaQH coolant is inferior to the sodium because it gives fairly wide variations in wall temperature for variations in the wall-coolant ORNL-LR-DWG 2046 400 e o - UNCOOLED WALL N s w <= = g —COCLANT TAKING 3.3% & <200 OF FUEL HEAT GENERATION o = u L O o s o 4ol z -~ COOLANT TAKING 4.9% W OF FUEL HEAT GENERATION a g g = SODILM VELGCTITY = 10 1ps = W FUEL—POWER DENSITY = { kw/cm> o ; Re = 100,000 4 Pr =10 z @ 3 M 5 ul _d r lad E o o [T i L =S0DIIV o R, e = INCONEL WALL \ 2 r—FUEL 30UNCARY LAYER } O /’ S , 4—in. DlA Fig. 40. Effect of Wall Cooling on the Temper- ature Distribution Through the Fuel Stream, the Wall, and the Wall Coonlant (Sodium). CRNL~-LR-DWG 2017 80C - e, - . W 800 - UNCOOLED WALL - ‘q‘: T © -COOLANT TAKING 34% OF o Py ! FUEL HEAT GENERATION T - 7 / I ) x 200 - - N PN e L Ll L = 12 2 O DOLANT TAKING 45% = L x 2 b g OF FUEL HEAT GENERATION w T -200 s e e 2 3 4 <= ot SODIUM VELOCITY = 10 fps o » FUEL-POWER DENSITY = 1 kw/cm? z o400 e T - fj Re = 10,000 ! w Pr= 10.0 v o . E L < -500 Ay ULy SRSy NPy Ay - j._. e L SODIUM - £, S = a1 A e INCONEL wWALL S TBO0 g e — o FUEL BOUNDARY LAYER ~——w - lpinflyin] 4-in. CIA : ~1000 Lo arent Fig. 41, Ef#fect of Fuel Viscosity on Temper- ature Distribution Through the Fuel Stream, the Wall, and the Wall Coolant (Sedium). Fuel vis- cosity assumed to be ten times greater than that of fuel assumed for Fig, 40. velocity. The wall temperature variations would be likely to lead to thermal distortion and warping or buckling of the wall. The use of a hydroxide as both moderator and wall coolant for an ARE type of core has some attractive possibilities. As can be deduced from Figs. 38 and 39, the use of perhaps 50 fuel tubes about 2,0 in. in diameter instead of g thick annulus of fuel would give lower uncooled wall tempera- tures, Figure 43 gives some idea of the possi- bilities of such a design. Closely fitted baffles would be required fo direct the hydroxide flow over the tube walls at a uniformly high velocity, Ir- regularities in wall temperature would tend to give progressive thermal distortion and deterioration in the hydroxide velocity distribution. The cumulative effects of this process might lead to a hot spot and severe corrosion of the tube wall. _ While quantitative data are not now available, some comments on the fuel boundary-layer heating ORNL-LA-DWS 2018 {eF} 200 FUEL CHANNEL O o= e — e I rrn i - - - - - - Lo o o w—— FUEL e = 100,000 2 o0 f oA —== FUELL Re=87000 1 [ Pr=10 = POWER DENSITY =1 kw/em3 w o & -2R0 T OSOGIUM FLOWING AT 10 fps axnp 15 T REMOVING 4.9% OF FUEL 10 >10 1.9 0.70 0.027 4,2 0.66 density, kw/cmB Limiting reactor fluid outlet 1550 1700 1200 1200 1300 1280 1730 3100 temperature, °F Limiting air temperature, °F 1240 1350 990 990 455 1280 1160 1230 Specific thrust {less nacelie drag), 22.4 29.0 14 14 19.4 26.0 28.8 38.5 1b/1b of air/sec Weight of propulsion machinery at 2.5 2.0 4.0 4.0 3.3 1.9 4.3 2.7 Mach 1.5 and 45,000 §1, 1b/1b of thrust Probable temperature coefficient +1073 -5 % 1073 ~5% 10"% ~10~3 +10™2 +10~7 +10-% +10~5 for fast transients, Ak/&-°C Remarks on controllability Difficult and Simple and Simple and Startup Difficuit and Difficult and Difficult and Diffieult and complex inherently inherently procedure complex complex complex complex reliable reliable difficult Total fuel investment (U2%) in 60 120 60 25 30 100 60 100 reactor, Ib Cost of fabrication and 16.00 1.50 1.50 16.00 16.00 16.00 16,00 16.00 reprocessing per gram of U235 dollars'® Efficacy of chemical fuel Very good Very good Very good Very good Poor Good Poor Poor dugmentation Major hozards Reactor runaway and Fuel spill; Fuel spill; NaOH spill; Burst of some Reactor runaway and Reactor runaway Reactor runaway and melt-down; sodium NaK fire NaK fire NaK fire part of 5000-psi melt-down; burst of melt-down; burst of fire system 300-psi system 3000-psi system (O)qu calculated from Pratt & Whitney Aircraft Div., Nuclear Propuision Program Engineering Progress Reporf, No, 9, PWAC-75, p. 28. (b)General Electric Co., Aircraft Nuclear Propulsion, Department of Engineering, Progress Report No. 9, APEX-9 {Sept. 195‘3). (C)Dm‘q calculated from work of A. Dean and . Nakazato, Investigation of Mercury Vapor Power Plant for Nuclear Propulsion of Aircraft, NAA-5R-110 (Mar. 21, 1951). (d) Data caleulated from work of H. Schwartz, An Analysis of Inert Gas Cooled Reactors for Application to Supersonic Nuclear Aircraft, NAA-SR-111 (Sept. 8, 1952}, (e)Dcm caleulated from memorandum from C, E., Larson to G. Beardsley, Preliminary Comporison for Reprocessing Fuels from an SCWR and a CFR, ORNL CF-53-12-11 (Dee, 1, 1953). 65 temperatures required of an aircraft reactor coupled with leaktightness requirements and shield weight and residual radiation considerations make it seem unlikely that o sodiumecooled solid-fuel-element reactor could be reloaded readily. On the other hand, littie information was available in 1950 on fluids that might serve as vehicles for fuel. Hy- droxides had been considered by NEPA,4? but it was felt that serious corrosion problems would be inherent in their use because the oxygen in them is not bound tightly enough to give assurance of their remaining inactive at high temperatures. R, C. Briant, in April 1950, pointed out that, on the basis of chemical thermodynamics, the alkali fluo- rides should be inherently stable relative to the iron~-chrome-nickel alloys even at the high temper- atures required and advocated their use as circu- lating fuels. It was recognized, however, that the use of the fused fluorides as o circulating fuel would mean the opening of o whole new field of reactor technology that would be filled with un- knowns and that there would be no guarantee of success, |herefore it was decided in September 1950 that the major emphasis should be placed on the sodium-cosled solid-fuel-element reactor, while a substantial research program should be directed toward the solution of the corrosion problems associated with the hydroxides and the fluorides. Shield and Heat Exchonger Designs To implement the design effort on the sodium- cooled solid-fuel-element reactor, an intensive study was made of the shieilding problem by o joint ORNL-NEPA committee in the fall of 1950. A number of the reactor and shield designs included in the committee report®® are of interest. Figure 44 shows the first design prepared, which followed the NEPA practice of using conventional tube- and-shell heat exchangers disposed relotive to the reactor and to the pumps in a quite conventional fashion with the shield simply wrapped around the resulting assembly. The estimated shield weight for this assembly was over 230,000 1b, The lay- outs shown in Figs. 45 and 46 make use of an ONEPA Project Quarterly Progress Report for Pericd April T—June 30, 1950, NEPA-1484. SOReport of the Shielding Board for the Aircraft Nuclear Propulsion Program, ANP-33 {Oct. 16, 1950). unconventional heat exchanger in which the reactor coolant flows axially through the interstices be- tween small-diameter, closely spaced tubes, while the secondory circuit fluid passes through the tubes to give a virtually pure counterflow system, The shield weight for the tandem heat exchanger arrangement of Fig., 45 was estimated to be about 160,000 1b, while that for the annular heat ex- changer arrangement of Fig. 46 was 122,000 ib. This was close to the weight of the ideal maiched lead-water shield, the weight of which was esti- mated to be 116,000 Ib. A fourth configuration, which made use of lead as the reactor coelant, is shown in Fig. 47, This arrangement, the weight of which was estimated to be about 120,00C Ib, was designed to employ the lead reactor coolant as shielding material by placing the heat exchanger at the same region in the shield as would normally be occupied by the gamma-ray shielding material. Differential thermal exponsion appeared to pose some rather difficult structural problems in con- nection with the fairly large volume, low-temper- ature shield region inside the high-temperature heat exchanger shell, The design also had the disadvantage that the lead-corrosion showed no promise of solution. Concurrently with the 1950 Shielding Beard investigation, a second joint ORNL-NEPA group carried out an intensive study of the reactor and engine control problem.®® This group reluctantly reached the conclusion that a solid-fuel-element, high~temperature, high-power-density reactor might be unstable ond that if at all possible an effort should be made to obtain a reactor with a negative temperature coefficient, even if it meant compro- mising the reactor design, While there was liftle doubt that the sodium-cooled solid-fuei-element reactor cowld be controlled, it appeared that on problem unusually complex control system would be re- quired which, when coupled to the very complex contrel system required for the turbojet engines, would probably seriously impair the reliability of the power plant, In view of this serious develop- ment, the situation wos reappraised. Mt was de- cided that the matertials research work was still not sufficiently far along to permit shifting the major emphasis to a circulating-fuel type of reactor and therefore development would have to continue on a stationary-fuel-element reactor, It was felf, however, that it would be possible to use a design 67 89 ORNL—-LR - DW5. 1150 ORNL—L.R~DWG 1151 HEO AND ©h ADDED TO SHIELD ~=-—— e BOTTOM COMPOSITE SECTION Fig. 44. Shield Design for o Sodium-Cooled Solid-Fuel-Element Reactor with Coaventional Tube-and-Shell Heat Exchengers, ORNL ~LR-DWG 4152 -—-—- PRESSURE SHELL e e SODIUM ‘\'\ AN il e / - SECONDARY SODIUM OUTLET _\fl\\‘ —- fy N . JlENy 7 e PRIMARY COOLANT PUMPS L SECONDARY SODIUM INLET ———" 3 COOLANT EXPANSION TANK - Fig. 45. Design of a Sodium-Cooled Solid-Fuel-Element Reactor with o Tandem Heat Exchanger Arrangement, 69 PRIMARY COOLANT PUMP -, j SECONDARY COOLANT INLET -+ LEAD —==— . _ SPACE FOR THERMAL EXPANSION _,,// ey T B,C POWDER -— — o — INNER STEEL SHELL OF SHIELD — — % ORNL-LR-DWG 1153 T STEEL REFLECTOR __——-COOLLANT PASSAGE | CONTROL ROD ,//ff L 4 __——REACTOR CORE DIAPHRAGM TO PERM!IT /—‘ o THERMAL. EXPANSION OF CORE COQLANT TUBES —~—B,C POWDER ---SECONDARY COOLANT T QUTLET WC + B,C = Fig. 46. Design of a Solid-Fuel-Element Reacter with an Annular Heot Exchanger Arrangement. 70 ORNL-LR~- DWG {154 /_—-—"" A“__“-‘"‘"‘—‘ e T / \\\\ THIS OUTER SPHERE TO BE FILLED & ., ™~ ¥ o * 1 { A . . —_— y - . ? + =t : LiIRY . v X iy ; - . : } / \ e A ————— ——— / /s e 7N \\ LA : " N . 7 ) : ¥ X . 7 e e =, ) il 5 L | A / » S R PR i 2 O e e r LN a7 7 \ e N < ; e W “AREA RESERVED FOR CONTROLS WITH GASOLINE IN SELF-SEALING CELLS AN N H\{:O fiPUMP \ - 040-in. Fe, 0.25-in. INSULATION, 0.02-in. Fe _~0.02-in. Fe, 0.25-in. INSULATION, 0.10-in. Fe \Q SAME FOR ALL DUCTS AND INTERMEDIATE HEAT EX- CHANGER WALLS - BOTH SIDES OF THI5 PLATE TO BE SPRAYED WITH SORON Fig. 47. Design of a Lead-Cooled Solid-Fuel-Eiement Reactor with Heat Exchangers Aranged to Act as Gamma-Ray Shielding. similar to that of the KAPL Submarine Intermediate Reactor but with a molten fluoride salt containing uranium in solution instead of solid UO, in the fuel pins. Thermal expansion would push some of this fluid fuel out of the reactor core into an expansion tank to give the desired negative tem- perature coefficient of reactivity., Work on this design proceeded for nearly o year until, in the fall of 1951, it developed that the GE-ANP project had dropped its plan to base development on high- temperature liquid-cooled reactors and had instead returned to the air cycle. Since this change eased the pressure for ORNL to get an experimental re- actor into operation at the earliest possible date, the entire nuclear-powered aircraft situation was reappraised. No structurally satisfactory design had been evolved for a high-power-density reactor core em- ploying the molten-saltfilled fuel pins, and two problems associated with the fuel pins seemed well-nigh insuperable. First, there was, inherently, a temperature drop of over 1000°F between the center and the outside of the fuel pin. While this would not have been serious in a low power redactor of the type to be used for the Air- craft Reactor Experiment, it probably would have been quite serious in a full-scale aircraft reactor, major Second, as in any solid-fuei-eiement reactor, ade- quate spaucing of the fuel elements would have been exceedingly difficult tc arrange. Thus, on the basis of structural and heat transfer considerations, support and satisfactory maintenance of 71 attention was turned to the circuloting-fluoride- fuel reactor with its then more difficult materials problems. Considerable progress had been made in the investigations of the chemistry of fluoride fuels, and it was believed that in the long run the problems associated with the use of circulating fuels would prove easier to solve than the less obvious, but nonetheless vital, problems inherent in fixed-fuel-element reactors, initiated in October 1951 for examining the problems associ- ated with a full-scale aircraft nuclear power plant employing a circulating fluoride fuel. In one of the first studies, an examination was made of the possibility of piping the fluoride fuel directly to An intensive design effort was heat exchangers in the turbojet engines and thus eliminating the complications associated with an intermediate heat transfer circuit., The design study of this proposal is covered in ORNL-1287.31 Even by going to the exceptionally large reactor- crew separation distance of 120 ft and a crew- engine separation distance of 135 ft, which badly compromised the airplone design, the arrangement led to a shield weight actually greater than that obtainable intermediate heat transfer circuit, Also, the radiation dose level of about 6 x 10% r/hr at 50 ft from the reactor that would result from this arrangement would be completely intolerable. with an Shielding of the engine radiators ap- peared to be out of the question because of their large size. Ground-handling and maintenance prob- lems seemed to many people to be insuperable, Thus this arrangement was dismissed and attention was directed to reactor systems employing an intermediate heat transfer fluid. A careful study of arrangements of the reactor, intermediate heat exchanger, and shield was made in an effort to determine the effect of configuration on shield weight,>? Activation of the secondary coolant threatened to be a much more severe prob- lem in these arrangements than in the arrangements for use with solid-fuel-element reactors (Figs. 44 through 47) because delayed neutrons from the circulating fuel would be released in the heat exchanger. It was found that this key problem 3TR. W. Schroeder and B. Lubarsky, A Design Study of a Nuclear-Powered Airplane in Which Circulating Fuel is Piped Directly to the Engine Air Radiators, ORNL- 1287 (Apr. 16, 1953). 52A. P. Fraas, Three Recctor-Heat Exchanger-Shield Arrangements for Use with Fused Fluoride Circulating Fuel, Y-F15-10 (June 30, 1952). 72 could be handled by keeping moderating material out of the heat exchanger so that most of the neutrons would escape before they would slow down. By filling 5 to 10% of the heat exchanger volume with boron carbide, most of the neutrons that did not escape would be captured in boron rather than in the secondary coolant. This solution to the problem of neutron activation of the second- ary coolant makes the circulating-fuel reactor superior to o homogeneous reactor. In homogeneous reactors, moderation of the deloyed neutrons by moderator-fue! would greatly aggravate the prob- lem. The first circulating-fluoride-fuel reactor, inter- mediate heat exchanger, and shield arrangement studied — an arrangement in which the reactor and heat exchanger were placed in tandem — is shown in Fig., 48. To keep the octivation of the sodium in the seconduary circuit to a tolerable level, it was found that it would be necessary to separate the heat exchanger from the reactor core by at least 12 in. of good moderating material followed by a 1 in. thick layer of boron carbide (or, if B'® were used instead of natural boron, a thickness of 0.2 in.). A carefu!l analysis of this arrangement dis- closed also that the pressure shell should be separated from the reactor core by a layer of boron carbide of similar thickness to keep the pressure shell from becoming ¢ more important source of gammas than the core. |t also became clear that activation of the secondary coolant by delayed neutrons emitted from the fue! in the heat ex- changer could be markedly reduced by spreading the heat exchanger out in a thin layer and thereby increasing the neutron escape probability, Further, it was observed that with the tandem arrangement (Fig. 48), the lead shielding required just for the heat exchanger constituted ¢ major portion of the total shield weight, The annular heat exchonger arrangement shown in Fig. 49 was evolved to place the heat ex- changers around the reactor within the primary reactor shield and thus eliminate the extra lead shizlding of the heat exchangers. ment gave an estimuted shield weight of 128,000 Ib, as compared with 156,000 Ib for the tandem arrarigement, led to the conclusion that an additional weight This arrange- Careful examination of this design saving could be realized by changing the gecmetry of the design to make it more nearly spherical, The spherical arrangement shown in Fig. 50 was €L — s LT e — i T 5 A #7 H,0 MODERATOR 7 INLET PIPE ~ DWE E-Y-F 7-1T7{R!¢ p \\\\ : \\ :‘APPROX, 12 1n. HZO AROUND HEAT ) \ | ; \ EXCHANSER AND PUMP HOUSINGS ElL PLATE FOR REACTOR AND UPPCRT, 2 REQUIRED o 2, A 2, 2 r’?’ \\ / Y, i \ g \"'\ § H,0 MODERATOR Y /! SIS QUTLES ; SECONDARY i M PHIPE QUTLET. ! : ook U AR \ % ) MEITEEEN L i U7 COULANT i ; S LINES il ’ ¥ —] N ) f b gl‘ P e 7 : : \ i W & . —- ¥ % oplh L —IFC e ———— / Y pefh g PRE ' i e 4 \w:‘\ /’/pz', VW ji, — y‘(—'\’ / 3 R Y YIS SRS e e ——— - & é 5/ “ Yioin, THICK SHELL | A "\ Y VCIRCULATING AR ANG . X THERNMAL EXPANSION GAP Vs ' L EXPANSION /f/ Ya-in THICK NK, 4 REQ'D N TANK, 4 REQ %\ 3 P E o 7 A o LS A —in. THICK ALUMINUM B, I T / Jl e A 2 ENTIRE REACTOR AND HEAT EXC Fig. 48, Water-Mederated Circulating-Fuel Reactor with a Tandem Heat Exchanger. Z NEUTL PUNP HOUSING VL DWG. E-Y-F 7-213aR; i 7 o b 7 Narx INLET NagK OUTLET /Yo BaC CANNED WITH Q060 SIEEL / - Yo NGOH /"{/2 STEEL THEAT EXCHANGER OUTLET MANIFOLD NoOH - NaK HEAT EXCHANGER by OD TUBES ON J/3d -1 5Q P CP e L TTL 4 T2 :_"- - HELICAL COl_COOLANT TUBES / 1250 00 TUBING i % STEEL ooééiggoE fii?i—l by o I %G Tk fl T f /8 STEEL 0050 STEEL - 2 PRESSURE SEELL NaOH PUMP *-‘ufi, fl@@@@@@@@@@@fidc \ CA@GCOO@QQGQde“ O \\M“QQQOQQQuOQVQOQdOO Y B,C CANNED IN 0.060 STEEL NaOH OUTLET TO [ NaOF-Nek HEAT EXCHANGER -~ /) 0% CIRCULATING NaOH REFLECTOR HEAT FXCHANGER INLET MANIFOLD - = ' R EXCHANGER QUTLET 8 ANNULAR TYPE HEAT LINE, 6 REQ'D EXCHANGER % 00 TUSFS ON @32 CoG SQ PITCH q “ & o ,.,/ & g 7 ra A / P Wi A /,;/ A P & NaK INLET T ET # r A NOTE: DIMENSIONS ARE IN INCHES Fig. 49. Schematic Longitudinal Cross Section Through Hydroxide-Meoderated Circulating-Fuel Reactor with Annular Heat Exchengers, \ e Nax INLET*-‘Q%J{\////,,/ S // \\ ~ 4 T I S0% BeO._ 7 ] 109 Na ™ f f PRESSURE SHELL— { INNER PRESSURE VSHELL OF SHIELD - \\\ " BORATED H,0 N\ AN ., ™ ~ R SELF SEALING RUBBER-—™N. SN \Q Narx OUTLET . _._____fi —t 7 ; _PUMP VE S - , / PUMP DRIVE SHAFT DWG. 15646 / \ / / \ | / N | Mo / \ e Y . ~ \\ \ / . E‘ e \\\ \\ \\ [ e~ NN \ | O, ™0 NN 5 ’ }‘[} ~ NN \\\ i ' ", \\\\ \ " A \ . NN \ 5 4 { ‘\\ \\ \\ Y y Y } \ i | \ L95% B0 || \ | . [ 5% Na b ------ HEAT EXCHANGER TUBE BUNDLE PUWMP DRIVE SHAFT Fig. 50. Arrangement of Circulating-Fuel Reflector-Moderated Reactor with Spherical Heat Exchanger. then designed, and it was found to give an esti- mated shield weight of 120,000 b, The shield weights given here were estimated for o quasi unit- shield design condition, namely, 7 r/hr at 50 fi from the reactor, and 1 r/hr inside the crew com- partment. Various degrees of shield division were also considered!” in an effort to reduce shield weights for the various designs, and in each instance the spherical-shell heat exchanger ar- rangement shown in Fig. 50 was found to be superior. In all cases, the reactor output design condition was 400 Mw, and the reactor core di- ameter was about 30 in, 75 Reactor Core Configurations The reactor cores used in the shield design studies of Figs. 48 and 49 were conventional in that the moderator was distributed throughout the active fuel region with relatively little lumping. However, the reoctor core design shown in Fig. 50 was evolved on a quite different basis. A brief account of the reasoning that led to this reflector- moderated reactor design may be of interest. A number of people had felt that a small (perhaps 18-in.-dia) fast reactor might be built to utilize one of the uranium-bearing fluoride-salt fuels. Rough calculations made by T, A, Welton indicated that the high concentration of uranium atoms re- quired to oachieve criticality with this type of reactor would be difficult to obtain in any fluoride salt ‘melt likely to have desirable physical proper- ties. Others felt that the concentrations of Li” and Be in the fluoride melt might be increased to the point where their moderating effect would be sufficient to make possible a homogeneous fused- fluocride reactor., A minimum critical mass of the order of 150 kg was indicated by one- and two- group calculations for such a reactor. Since the shield design studies had clearly shown the desirability of a thick reflector, it was felt that it might be possible to capitalize on this thick reflector and effect a major reduction in critical mass with a quasi-homogeneous flucride fuel. (At the time, a fairly high uvranivm concentration in the fluoride mixture was not considered to be too serious,) Multigroup caleulations indicated that a beryl- lium reflector could be made so effective that the critical mass could be cut to something of the order of 15 kg.>® This prediction was later con- 18 |t has also been firmed by critical experiments, found that the good high-energy neutron-scattering cross section of the fluorine in the fuel is more important for o reactor of this type than the moder- ating effects of Be or Li’. In fact, the neutron- scattering cross section is so much more important that heavier elements such as No and Rb may be used in the fluorides instead of Be or Li’7 with fitile effect on critical mass. The heavily lumped fuel region of the reflector- moderated reactor has a number of major advan- tages. The removal of all structural material from 33C. B. Mills, The Fireball, A Reflector-Moderated Circulating=Fuel Reactor, Y-F10-104 {June 20, 1952), 76 the core except the core shells reduces parasitic neutron capture in structural material to a minimum and hence reduces the critical mass. The place- ment of most of the moderating moterial in the reflector gives a smaller diameter core for a given power density in the fuel and hence a lighter shield. Many circulating-fuel reflector-moderated reactor arrangements have been proposed to take advantage of the spherical-shell heat exchanger and shield arrangement shown in Fig. 50, In general, it appears that there are eight basic types of con- struction that might be employed. The simplest type, shown in Fig. 51, comprises a thick, spheri- cal shell of moderator surrounding o spherical chamber containing liquid fuel. Ducts at the top and bottom of the shell direct cold fuel into the reactor core and carry off high-temperature fuel. Such an arrangement has two major disadvantages. First, the well-moderated neutrons reflected to the fuel region from the reflector tend to be absorbed near the fuel-reflector interface so that the power density falls off rapidly from that interface to a relatively low value at the center. Second, the flow pattern through such a core is indeterminate, and large regions of flow separation and probable stagnation would be likely to occur in a highly irregular, unpredictable fashion, although vanes or screens at the inlet might be effective in slowing down and distributing the flow, The arrangement shown in Fig. 52, which makes use of a central *island,”’ appeared to be more promising. The central] islond has the advantage that it reduces the critical mass and yields a more uniform power distribution. Thus the extra complexity of a cooling system for the island appears to be more than offset by the reduced critical mass, improved power distribution, and much superior hydrodynamic characteristics. The most serious problem associated with the arrangement of Fig. 52 appeors to be that of cooling the moderator,®4 a problem common to all high-power density reactors. If beryllium is used as the reflector-moderator material, closely spaced cooling passages must be employed in those portions close to the fuel region to remove the heat generated by gamma-absorption and the S4R. W. Bussard ot al., The Moderator Cooling System for the Reflector-Moderated Reactor, ORNL-1517 {Sept. 1953). ORNL-LR-DWG 283% e REFLECTOR- MODERATOR Fig. 51. Simple Two-Region Reactor Core with Thick, Spherical Shell of Moderator Surrounding a Spherical Chamber Containing Liquid Fuel. ' neutron-slowing-down processes. Other arrange- ments have been considered; for example, the high femperature gradients and thermal stresses induced in the beryllium in this fashion might be avoided if a layer of a liquid such as lead or bismuth could be interposed between the fuel and the refiector-moderator regions, as indicated in Fig. 53. This liquid could be circulated to carry off the heat and the beryllium coeling problem would be markedly relieved. | It appears feasible to use graphite in direct contact with fluoride fuels without damage to the graphite or contamination of the fuel. Therefore a possible design (Fig. 54) comprises a block of graphite drilled to give a large number of parallel passages through which the fuel might flow, This design, in effect, gives a very nearly homogeneous mixture of fuel and graphite in the reactor core. A variation of this design is shown in Fig. 55. Several concentric shells of grophite might be placed in such a way that they would serve to guide the fuel flow and at the same time oct as moderating material, From the hydrodynamic stand- point, either of these arrangements oppears to be preferable to the screens or vanes placed in the " ORMNL-[.R-DWG 2836 L SO 2 OO 5. A -MODERATOR REFLECTOR -MODERATOR Fig, 52. Three-Region Reactor Core with Central Island of Moderating Materials. fuel inlet mentioned in connection with Fig, 51. While these arrangements appear to have the advantage of simplifying the core design and dispersing moderator through the fuel region, calculations indicate that the fluoride fuel com- pares favorably with graphite as o moderating material, and therefore the arrangements of Figs. 54 and 55 are little better from the nuclear stand- point than that shown in Fig. 51, A number of different types of fluid-moderated reactors has been considered. One variant is shown in Fig. 56. A set of coiled tubes through which sodium hydroxide could be pumped might be placed in the reactor core, These could be made to serve both to improve the fuel velocity distribution and to moderate fast neutrons in the reactor core, The principal disadvantage associ- ated with such an arrangement is that it would be difficult to avoid local hot spots in the liquid fuel in zones where flow separation and stagnation might occur. Alse, the relatively large amount of structural material in the tube walls would capture 77 ORNL-LR-DWG 2837 X ORNL -LR-DWS 2528 LEAD QR BISMUTH GRAPHITE BLOCK — \REFLECTOR—MODERATOR MODERATOR REFLECTOR-MODERATOR Fig. 54. Reactor Core with Fuel Channels in Fig. 53. Five-Region Reactor Core with Pro- Graphite Block. vision for Ceoling Reflector-Moderator Regions. ORNL-LR-DWG 2334 ORNL-LR-DW6 2833 . - HYOROXIDE GRAPHITE SHELLS - Fig. 55. Reactor Core with Grophite Shells in Fig. 56. Fluid-Moderated Reactor Cere with Fuel Channel. Coiled Tubes for Circulating the Moderator. 78 CHNL-LR-BWE 284G / . HYDROXIOE - T REFLECTOR-MOBERATOR Fig. 57. Fluid-Moderated Reactor Core with Straight-Tube Fuel Possages and Provision for Circulating Moderator Around Fuel Tubes, ORNL--L'!'DWE} 2841 / REFLECTOR-MODERATOR - HYDROXIDE Fig. 58, Fluid-Moderated Reoactor Core with Spheriodized Fuel Passages and Provision for Circuloting Moderator Around Fuel Passages. ' a substantial percentage of the neutrons, and therefore the critical mass would be increased. The arrangement of Fig. 57 also presumes the use of a fused hydroxide as a fluid moderater, The fluoride fuel would circulate through the circular passages and pass down through the reactor core, while the hydroxide moderator would circulate through the spaces between the fuel passoges. Because of the fuel boundary-layer heating problem (cf., section on “‘Temperature Distribution in Circulating-Fuel Reactors’’), boffles would have to be provided arocund the fuel tubes so that the hydroxide could be circulated at a high velocity over the tube wail with good velocity distribution to prevent hot spots. The arrangement of Fig, 58 is similar to that of Fig. 57, except that the tubes are specially shaped to reduce the volume of the header regions and fo give a more nearly spherical core and hence a lower shield weight. ' DETAILED DESIGHNS OF REACTORS Sodium=-Cooled Solid-Fuel-Element Reactor The first detailed design studies of reactors were based on sodium-cooled solid-fuel-zlement reactor cores, ond several types of fuel element were examined. The pin type used in the SIR core?? appeared to be attractive, but the problems of supporting the pins and maintaining wniform spacing between them were exasperatingly difficult, particularly . for high-power-density cores. An arrangement that promised to give a much higher power density potential incorporated stainless- steel-clad sandwich fuel plates having a sintered UQ, and stainless steel core, os described in the previcus section on "*Materials.”” The most highly developed design of this character is that shown in Fig. 59, which was prepared in the summer of 1952 by A. S. Thompson, This core was designed to employ a fue!l element of the type shown in Fig. 60, but it is equally well adopted to the use of sandwich tube fuel elements of the type shown in Fig. 61. The core design was based upon the flow of sodium downward through the annular fuel element matrix and the reflector and then radially outward and upward through the heat exchanger in the annulus between the reflector and the pressure shell. Pumps at the top of the pressure sheil were designed to toke the sodium os it left the heat exchanger and deliver it back to the core inlet passage. 79 CRNL-LR-DWG 528 7 |~ FLANGED JOINT Mak IN SHELL ISLAND o HEAT EXCHANGER --PRESSURE SHELL - COR - No-TC-NaK ~-—FUEL ELEMENTS __ ——-Be REFLECTOR ———Be S ha PASSAGE TO REFLECTOR. ium-Cooled Solid-Fuel-Element 100-Mw Reacter. 59. Sod Fig. 80 L8 km«—---———--—la (FUEL BUNDLE HEIGHT) - e oo AT T )//f AR 5 ey Y ‘:‘——__;f}_ 7; 4 ! FUEL CORE ASSEMBLY L% - A t—SEE DETAIL B ORNL-LR-DWC 526 P N .’/ “ /,// /,r\ . ‘5‘500 ‘\ ol . i \\:‘ :“ ‘;\:\m’“‘ e R Fu" bi i BERE .gh_gk\ufi#‘flt\r JERpE wh n R ;,H‘ \J S L - il T j “fHfiH‘-”#jfi: s — - fiw:j i fp!_"_!' it j s et W‘ i T D i” [r/ld;‘!'l ’3!‘ '/“L;g' PLATE SPACER rey RODS 0.040 DIa~ SEE DETAIL A~ ' FUEL PLATE BUNDLE PLAN ASSEMBLY BOTTON HEADER - THZRMAL SHIELD CAP N o DweG. 18336 e THERMAL SHIZILD CAP D TOP MANIFOLD 1 ____TOP HEADER -—-BeC MODERATOR AND RZFLECTOR i ‘I/’THERMAL SHIELD ASSEMBLY TICM TUBE SHEET —— STUD -———SUPPORT ASSEMBLY ~~| - FUEL OUTLET MAN:FOLD "~ THERMAL. SHIZLD BOTTOM - 80% in. REF. - 2395 in REF —— —»— Y \ " “SUPPORT ASSEMBLY HELIUM MANIFOLD Fig. 62. Vertical Section of the ARE. 84 02 4686 8 12 18 SCAILLE IN INCHES i ) OWG. 15847 —INSTRUMENT TUBE - (INGCMEL) TUBE GOl (INCONEL ) - REFLECTOR DGE BLOCK - (BERYLLIUM DXICE) (NGONEL )~/ —-— REFLECTOR BLOCK (BERYL_IUM OXICE) e REFLECTOR COOLANT TUBES {INCOMNEL) /. e / T ; ) , o A 0 . el o3 jpem=e e AN ) , , A A ONGONEL) N i N (XD W \‘ e GORE SLEEVE INCONEL TTTSGARETY ROD GUIDE SLEEVE (INCONEL) SCALE N INCHES Fig. 63. Cross Section of the ARE. 85 DWG Y-F7-17'4 - FUEL TUBE SPACER AND CONTRCL ROD TRACK SUPPORT BAFFLE STEEL ROLLER — o CADMIUM CYLINDER . | |7 f i) STEEL SHAFT - ENDLESS TRACK FOR SPROCKET- DRIVEN CONTRCL RODS . Tlem— 10 N (RCTATED 20°) - TCP TUBE SHEET L FUEL EXPANSION fb Py ) TANK ASSEMSLY - — -~ //\ T COOLANT “ COOLANT QUTLET ,,_J He= - COOLANT INLET e = ; 4 < oy ‘ l 2 e : | i . (] . [ : i T gt C i |_ il | [ e i r fii T- i < [ |7~ HEAT EXCHANGER CENTRAL = “; ISLAND SPACERS O " ; o Il Jlj " HEAT EXCHANGER RING SPACER “~FUEL PUMP VOLUTE FUEL PUMP.. |/ i CASING Fig. 64. Reactor Design for Use with Liquid Moderator (Water or Hydroxide) ond Circulating Flvoride Fuel, 86 freezing the fuel at zero or low power if water were used, it was planned that a thin layer of insulation would be placed between the walls of the double-walled fuel tubes. It would thus be possible to operate the reactor with the water at a substantially lower temperature than that of the fluoride fuel. Unfortunately, the thermal insulation would preclude cooling of the fuel-tube walls during high-power operation, and hence the allow- able flucride fuel temperature would be perhaps 300°F lower thaon might otherwise be possible. The temperature differential between the water and the fuel could be reduced, of course, by pro- viding o heavy pressure shell and operating the reactor with the water at high temperature and pressure. However, a major disadvantage of this arrangement would be that to keep the stresses in the fuel tube walls to within reasonable values it would be necessary for the fuel system to operate at high pressures. In tum, there would be dif- ficulty with the pump-shaft seals, and the pressure shell would be excessively heavy, Two variants of this design were prepared. In the first, the reactor was designed fo generate steam for a supercritical-water cycle in which the moderator region of the reactor would serve as the feedwater heater. In the alternate arrangement, the heat added to a hydroxide moderator could be dumped at a high temperature, while the bulk of the heat would be transmitted from the fuel to NaK in the heat exchanger and the NaK would, in turn, be directed to turbojet engines. A major innovation in heat exchanger design was introduced which involved the use of a fairly large number of smali tube bundles with the tubes terminating in small, circular-disk headers. This arrangement had the advantage that the heat exchariger could be fabri- cated in elements, and each element could be carefully inspected and pressure tested. The efements or tube bundles could then be welded into the pressure shell with a relatively simple, rugged joint. By breaking the heat exchanger up in this fashion it was believed that the ultimate cost could be markediy reduced and the reliability substantially increased. The principal uncertainty associated with this alternate arrangement was that it was difficult to see how a sufficiently uni- form hydroxide flow distribution could be main- tained over the outside of the fuel tubes through the core. If the flow were not uniform hot spots might form ond rapid corrosion of the tube wall by the hydroxide would result. As discussed previously, both designs gave a high shield weight because of the unfavorable geometric effects associated with the tandem reactor—heat exchanger arrangement, The problems associated with the reactor core arrangement designed for use with annular heat exchangers (Fig. 49) are in direct contrast to thase of the tandem heat exchanger arrangement. Al- though the hydroxide flow through the moderator tubes in the core could prebably be kept at a uni- formly high velocity and hence the hydroxide tube wall would be cooled effectively, the turbulence pattern in the fluoride fuel flowing across the moderator tube coils would probably be erratic and unpredictable and local hot spots in the fuel would be likely to occur. The hot spots in the fuel might cause local boeiling and, possibly, instability from the reactor control standpoint. Reflector-Moderated Circulating«Fuel Reactor The design shown in Fig. 65 is representative of a series of reflector- moderated reactors employing sodium-cooled beryl- lium as the moderator and reflector material, A fairly complete set of dota for these reactors is given in Tables 12 and 13. The designs for these have been the most carefully worked out of any full scale ORNL-ANP reactor designs prepared to date and hence merit special attention, particularly since the problems dealt with are common to most high-temperature liquid-cooled reactors. The cross section (Fig. 65) through the reactor core, moderator, and heat exchanger shows a series of concentric shells, each of which is a surface of revolution about the vertical axis. The two inner shells surround the fuel region at the center (that is, the core of the reactor) and separate it from the beryllium island and the outer beryllium re- flector. The fuel circulates downward through the bulbous region where the fissioning takes place and then downward and outward to the entrance of the spherical-shell heat exchanger that lies he- tween the moderator outer shell and the main pressure shell, The fuel flows upward between the tubes in the heat exchanger into two mixed-~ flow fuel pumps at the top. circulating-fluoride-fuel From the pumps it is discharged inward to the top of the annular passage leading back to the reactor core. The fuel pumps are sump-type pumps located in the expansion tank at the top, A horizontal section through this B7 '/ 7/ 7 HEAT EXCHANGER - // HEADER FOR Na TO ' (STRES$ =400 psi) _ . INNER GORE SHELL — Ly« (STRESS =200 psi) TUBE WALL SN ' REFLECTOR Na { / REFLECTOR--— - OUTER CORE SHELL — | CUTLET HEADER-- | Be ISLAND 17 1 (STRESS =50 psi) - . L ~ e LT . < S i <, ¥ - b < DWG. 22342 ~—IMPELLER (STRESS=500 psi) * - UPPER PUMP DECK .~ LOWER PUMP DECK TN ——TUBE HEADER SHEET N (STRESS =1250 psi) . TUBE WALL A (STRESS = 45C psi) - - REFLECTOR SHELL 8'C LAYER INCONEL JACKET 71 COOLING TUBE CONNEGTER v wdj g 4 TUBE BUNDLE vy SPACERS JF ghl/ " TUBE BUNDLE i /7 JACKET FOR 8'° / / LAYER - B0 LAYER - PRESSURE SHELL NER T Na PASSAGE / " PRESSURE SHELL < {STRESS = 625 psi) T ({STRESS = 1250 psi) - TUBE HEADER SHEET (STRESS = 600 psi ) Fig. 65. Reflector-Moderated 100-Mw Reactor, Stresses for key structural elements are given, TABLE 12. PRINCIPAL DIMENSIONS OF A SERIES OF REFLECTOR-MODERATED CIRCULATING-FUEL REACTORS Power, megawatts Core diameter, .in. Power density in fuel, kw/emS Pressure shell outside diameter, in, Fuel System Fuel volume in core, f3 Core inlet cutside diameter, in. Core inlet inside diameter, in. Core inlet areq, in.2 Fuel volume in inlet and outlet ducts, f3 Fuel volume in heat exchanger, ;3 Fuel volume in pump and plenum, 3 Total fuel volume circulating, 13 Fuel expansion tank volume, ftz (8% of system volume) Fuel Pumps Fuel pump impeller diameter, in.. Fuel pump impeller inlet diameter, in. Fuel pump impeller discharge height, in, Fuel pump shaft center line to center line spacing, in. Plenum chamber width, in. Plenum and volute chamber length, in. Plenum end volute chamber height, in. Impeller rpm ‘ E stimated impeller weight, b Impeller shatt diameter, in, Impeller overhang, in. Critical speed, rpm Sedium Pump Na pump impeller diameter, in. Na pump impeller inlet diameter, in. Na pump impeller discharge height, in. Ha exponsion: tank volume, 13 (]D% of system volume) Na in Be paséoges, 3 Na in pressure shell, fi Ma in pump and heat exchanger, f13 Fuel-te-NaK Heat Exchanger Heat exchanger thickness, in. Heat exchonger inside diameter, in. Heat exchanger outside diameter, in. Heat exchanger volume, £ Angle between tubes and equatorial plane, deg Number of tubes Tube diameter, in. Tube spacing, in. Number of tube bundles Tube arrangement in each bundle 50 18 1.35 48.5 1.3 10 40 0.4 1.25 0.3 3.25 0.26 075 3.5 1.1 20 14.5 30 2.0 2700 1.5 12 6000 3.4 2.4 0.75 0.08 0.43 0.15 0.20 1.7 42 45,4 27 2304 0.1875 0.2097 12 B x 24 100 18 2.7 50.6 1.3 10 40 0.4 2.5 0.3 4.5 0.36 4.5 1.5 21 15 31 2.0 2700 12 1.75 13 6000 4.1 2.9 0.9 0.09 0.47 0.15 0.25 2.75 42 47.5 10 27 3744 0.1875 0.2097 12 13x 24 200 20 3.9 56.4 11 7.7 49 0.5 0.5 7.8 0.62 8.5 5.5 1.8 22,5 15.5 33 2.4 2500 17 14 5200 5.0 3.5 1A 4.65 44 53.3 20 27 6600 0.1875 0.2119 12 22 % 25 300 23 3.9 62.0 2.7 12.8 67 0.7 7.5 1.0 11.9 0.95 10 6.75 3.2 27 17.5 37 3.0 2300 24 2.25 15 5000 5.9 4,2 ]l2 5.9 .47 58.8 30 27 9072 0.1375 0.2094 12 28 x 27 89 TABLE 12 {continued) Moderator Region Volume of Be plus fuel, £13 Volume of Be only, 3 No. of coolant holes in reflector No. of coolant hales in island Na coolant tube inside diameter, in. Na coolant tube wall thickness, in. Na pressure shell annulus thickness, in. region is shown in Fig. 66. A pump of the type proposed recently completed 1600 hr of operation in a fluoride system with pump inlet temperatures ranging from 1000 to 1500°F. The moderator is cooled by sodium which flows downward through passages in the beryllium and back upward through the annular space between the beryllium and the enclosing shells, Two centrifugal pumps at the top circulate the sodium first through the moderator and then through the small toroidal sodium-to-NaK heat exchangers around the outer periphery of the pump-expansion- tank region. Two sodium pumps and two sodium- to-NaK heat exchangers are provided so that failure of one pump will not completely disable the re- actor. Two fuel pumps were provided for the same reason. The design of Fig. 65 presumes that canning of the beryllium will be required to protect it from the sodium, but that trace leaks of sodium through the Inconel can connections can be tolerated. As a result, the Incone!l canning tubes that would be fitted into the rifle-drilled holes in the reflector were designed to be driven into tapered hores in the fittings shown at the equator, while the outer ends of these same tubes would be rolled into their respective header sheets at the top and bottom, The tube-connecting fittings at the equator would also serve as dowels to locate the two beryllium hemispheres relative to each other. Corrosion tests on the berylliumesodium-lnconel system are under way, and preliminary tests indi- cate that there is good reason to hope that it will be possible to allow the sodium to flow directly over the berylliuvm; if so, the rather complex canning operation would be unnecessary. Of even more importance, however, elimination of the Inconel canning would remove poison from the reflector and reduce both the critical mass and the production of capture gammas {and hence the shield weight). 90 22.4 22.4 25.8 31.5 21.1 21.1 24,0 28.8 208 208 554 554 86 86 210 210 0.155 0.187 0.187 0.218 0.016 0.016 0.016 0.016 0.125 0.125 0.187 0.200 The spherical-shell heat exchanger that makes possible the compact layout of the reactor—heat exchanger assembly is based on the use of tube bundles curved in such a way that the tube spacing is uniform irrespective of “latitude.”>® The indi- vidual in headers that resemble shower heads before the tubes are welded tube bundles terminate This arrangement facilitates assembly because it is much easier to get a large number of small tube-to-header assemblies leaktight than one large unit, in place. frurther, these tube bundles give o rugged, flexible construction {resembling steel cable) that is admirably adapted to service in which large amounts of differential thermal expan- sion must be expected. This basic tube bundle and spacer construction was used in a small NaK- to-NaK heat exchanger that operated for 3000 hr with a NaK inlet temperature of 1500°F3% and in a fluoride-to-NaK heat exchanger that successfully for over 1600 he,37 The allowable power density in the fuel region may be limited by radiation~-damage, control, moder- ator-cooling, or operated considerations, While the experimental results obtained to date are difficult to interpret, no clearly defined radi- ation-damage limit to the power density has been established, and it is entirely conceivable that radiation-damage considerations will prove to be less important than other factors in establishing a limit on power density, The kinetics of reactor control are very complex, Work carried out to date indicates that conirol considerations are likely hydrodynamic SSA. P. Fraas and M, E. LaVerns, Heat Exchanger Design Charts, ORNL-1330 (Dec. 7, 1952). 6G. H. Cohen, A, P, Fraas, and M. E. LaVerne, Heat Transfer and Pressure Loss in Tube Bundles for High- Performance Heat Exchangers and Fuel Elements, ORNL-1215 (Aug. 12, 1952). 578.. Wilner and H. Stumpf, Intermediate Heat Ex- changer Test Results, ORNL C¥F-54-1-155 (Jan. 29, 1954), TABLE 13. HEAT TRANSFER SYSTEM CHARACTERISTICS FOR A SERIES OF REFLECTOR-MODERATED CIRCULATING-FUEL REACTORS REACTOR POWER, megawatts Fuel-to-Nak Heat Exchanger and Related Systems Fuel temperature drop, °F NaK temperature tise, °F Fuel AP in heat exchanger, psi NaK AP in heat exchanger, psi Fuel flow rate, |b/sec NaK flow rate, Ib/sec Fuel flow rate, cfs NaK flow rate, cfs Fuel velocity in heat exchanger, fps Fuel flow Reynolds number in heat exchanger NaK velocity in heat exchanger, fps Over-all heat transfer coefficient, Bfu/hr'ffz'OF Fuel-NaK temperature difference, °F Sodiumeto-Nak Heat Exchanger and Related Systems Na temperature drop in heat exchanger, °F NoK temperature rise in heat exchanger, °F Na AP in heat exchanger, psi NoK AP in heat exchanger, psi Power generated in islond, kw Power generated in reflector, kw Power generated in pressure shell, kw Na flow rate in reflector, |b/sec Na flow rate in island and pressure shell, Ib/sec Total Na flow rate, |b/sec No temperatfure rise in pressure shell, °F Na AP in pressure shell, psi Na tempercture rise in island, °F Ne AP inisland, psi Na temperature rise in reflector, °F Na AP in reflector, psi Na-Nak temperature difference, °F Shield Cooling System Power generated in 6-in, lead layer, kw Power generated in 24-in. HZO layer, kw to limit the power density in the reactor to a value such that the temperature rise in the circulating Hluoride fuel will not exceed something like 1000 to 2000°F /sec. A 2000°F /sec temperature rise in the fuel would imply a power density of approxi- mately 4 kw/cm®. The difficulties associated with 50 100 200 300 400 400 400 400 400 400 400 400 35 51 61 75 40 58 69 85 263 527 1,053 1,580 474 948 1,896 2,844 2.1 4.2 8.4 12,6 10.5 21.0 42,0 63.0 8.0 9.9 1.2 12,2 4,600 5,700 6,700 7,000 36.4 44.9 50.9 55.4 3,150 3,500 3,700 3,850 95 100 110 115 100 150 150 150 100 150 150 150 7 7 7 7 7 7 7 7 500 1,000 2,000 3,000 1,700 3,400 7,500 11,200 190 350 500 620 53 72 154 234 22 28 51 76 75 100 205 310 28 39 30 26 4 6 6 6 72 m 120 124 32 21 12 12 100 150 150 150 36 27 18 18 43 110 210 300 350 <3 <6 <12 <18 cooling the moderator and with the hydrodynamics of fuel flow through the core increase with power density, as shown in the section on ““Temperature Distribution in Circulating-Fuel Reactors.”” The results of that work alse indicate that it would be desirable to keep the average power density in the 91 ¢6 DWS E-179864 ~—FUEL PASSAGE TO CORE _—Na PASSAGE IN ISLAND ——- —Na T0 REFLECTOR TUheoNoK INLET - Na- TO - NaX HEAT EXCHANGER -———— NoK INLET Na RETURN FROM ISLAND AND PRESSURE SHELL —— - Fig, 66. Horizontal Section Through Pump Region of Reflector-Moderated Reactor. fuel to about 4 kw/em®, This power density would mecn that a core diameter of 21 in, would be re- quired for a 200-Mw reactor, : Two major tenets of the design philosophy have been that the pressures throughout the systems should be kept fow, particularly in the hot zones, and that all structure should be cooled to a temper- ature approximately equal to or below that of the secondary coolant leaving the heat exchanger. Great care was exercised in establishing the pro- portions of the designs presented in Table 12 to satisfy these conditions. The temperature, pres- sure, and stress values calculated for the various stations in a typical design are indicated in Fig. 65. The stresses in the structural parts have been kept to @ minimum and the ability of the structure to withstand these stresses has been made as great as practicable. Thermal stresses are not indicated on Fig. 65, because it is felt that they will anneal out at operating temperatures and, at worst, will cause a little distortion which should not be serious. Examination of Figs. 35, 36, and 37 discloses that the pressure stresses in the major structural elements of Fig. 65 are quite modest. SECONDARY FLUID SYSTEM The ORNL effort has been devoted almost wholly to the reactor and shield, but a small amount of preliminary design and developmental work has been done on the rest of the system. This has been necessary portly because the feasibility of the power plant as a whole depends to o consider- able degree on the components outside the shield and partly because only by doing work of this character has it been possible to evaluate the incentives toward higher temperatures and power densities and such factors as the penalties at» tached to low T - \\ . A - BLEED AIR CONTROL VALVE f\ 3 NaK -TO-AIR — g RADIATOR \ /{ COMPRESSOR TURBINE NOZZILE Fig. 67. Diogram of Reflector-Moderated Reactor Power System, TABLE 14, EFFECTS OF SECONDARY CIRCUIT FLUID ON THE WEIGHT OF MAJOR SYSTEM COMPONENTS WEIGHT WEIGHT WEIGHT WEIGHT OF SHIELD WEIGHT TOTAL WEIGHT OF OF OF : PUMPS AND INCREMENT INCREMENT LIQUID PIPES RADIATORS PUMP DRIVES RELLATIVE TO RELATIVE TO {ib) {ib) (Ib) {ib} MakK (Ib) NaK (Ib) Nak 2,600 3,300 6,000 1,600 ¢ 0 Lithium 900 2,300 2,700 1,600 ~1,500 ~5,100 Sodium 2,400 3,100 5,850 1,400 =400 ~1,350 Potassium 3,300 3,900 6,500 2,100 z,000 4,100 lead 57,700 5,500 2,000 1,400 2,000 60,300 NaF-KF-LiF 1,150 1,45C 3,900 650 2,000 -1,5950 These comparisons were made on the basis of a 200-megawatt system. In order to keep stresses in high-temperature metal walls to conservative values, the peak pressure in the system was limited to 100 psi. Some attempts to optimize line size have been made, which indicate that the 100-psi value gives close to a minimum system weight if allowances are made for the extra weight of pumps and pump drive equipment and the thicker pipe walls required for the higher pressures. A number of methods of system conirol have been considered, If the circulating-fuel reactor performs as expected it will serve as an essentially con- stant-temperature heot source., Hf the pumps are operated at a constant rpm, the temperature rise in the NoK passing through the intermediate heat exchanger will be directly proportional to the power output, but the mean temperature of the fuel system wiil remain consfant.- Unfortunately, o substantial amount of power is required to drive the pumps both for the reactor and for the secondary system, and it probably will not prove practicable to keep the pumps running ot full speed if the turbojet engines are idling, This will probably be frue irrespective or whether the pumps are driven by air turbines or by electric or hydraulic motors. Actually, the turbojet engine characteristics are such that it would be more desirable 1o allow the pump speed to vary with engine speed; in fact, it seems likely that after the turbojet sngines have been started, the speed of the pumps could be allowed to reach equilibrium under any conditions from idling to full power and still give a reason- able set of flow rotes ond temperoture rises. Since the torque output of o turbine wheel falls off with rpm and since the torque required to drive the pump impeller increases as the sguare of the rpm, it is evident that the turbine-pump system would be exceedingly stable, Preliminary estimates indicate that the pumpedrive turbine speed would follow the turbojetengine speed very closely during an acceleration of the turbojet, It is clear that a full-scale power plant will be a quite com- plex system and that possibilities of instability and oscillation exist, but it also appears that the components con be proportioned so that o stable system will result. 95 Priority A-1 96 MAJOR DEVELOPMENT PROBLEMS The following outline of the key design problems of the circulating-fuel reflector- moderated reactor and the status of these problems at this time serves as a summary of the work that has been covered by this report and of the work that remains to be done to provide a sound basis for the design of a full-scale aircraft power plant, The principal reports that cover the work that has been done and an indication of the current priority of the remaining problems is included to give some idea of the progress that has been made and of the magnitude of the task that remains, OUTLINE GF MAJOR RMR DEVELOPMENT PROBLEMS Development Problem Status May 1254 Reports Fuel Chemistry and Corrosion Corrosion Harp tests and simple thermal-convection loops High-temperature-differential, high-velocity loops Radiation Damage and Corrosion In-pile capsule tests In-pile loop tests Physical Properties NaF-KF-LiF, NaF-BeF,, NaZrFyg, ete. NaF-RbF-LiF Other fuels and fuel carriers Selubility of UF4 and UF3 Methods of Preparation Xenon Removal Reprocessing Techniques High-Performance High-Temperature Heat Exchangers NaK-to-NakK Pressure losses for flattened-wire tube- spacer arrangement Heat transfer and endurance test NaK-to-Air Fabricability, perfoermance, and endurance tests (including study of character of failure) Fluoride-to-NaK Tube-to-hecder welding, endurance and performance tests Effects of trace leaks, and fabricability of spherical shell type Much favorable data No data Some favorable data Neo data, equipment being assembled Adequate data Data expected soon Data expected by Dec. 30, 1954 Some data Considerable experience Little data Seme favorable data Adequate data More tests needed Mare tests needed More tests needed Little data available ORNL-1515, -1609, -1649, -1692 ORNL-1649, -1692 ORNL-1692 ORNL. CF-53-3-261 ORNL-1215 ORNL-1330 ORNL-1509, -1692 ORNL. CF-54-1-155 OUTLINE OF MAJOR RMR DEVELOPMENT PROBLEMS (continued) Priority Development Problem Status May 1954 Reports Shielding Preliminary Designs Lid Tank Tests of Basiec Configurations Effects of thickness of reflector, pressure shell, lead, and boron layers Estimated Full«Scale Shield Weights Effects of power, power density, degree of division Activation of Secondary Coolant Estimated Measurements for neutrons from core Measurements for neutrons from heat exchanger Measurement of Short-Half-Lived Decay Gammas Refined Lid Tank Tests Experiments on Air Scattering Static Physics Multigroup Calculation—~Effects of Moderator Materials Effects of core diameter, fuel-region thickness, reflector thickness, reflector poisons, and special materials Critical Experiments Critical maoss with various fuel regions—Na, fluoride, fluoride- graphite Control red effects (rough) End duct leakage Danger coefficients for Pb, Bi, Rb, Li’, Na, Ni, etc. . Check on Multigroup Calculation Twao-region Three-region Core shell effects EHects of end ducts Many designs available ANP-53, Y-F15-10, ORNL-1575 Adeguate dota for ORNL-1615 preliminary design Adequate data for ORNL-1575 preliminary design Data adequate , ORNL. 1575 Data adequate ORNL-1616 Data needed Tests in progress Tests planned for late 1954 Tests in progress Adequate data expected ORNL-1515 by Sept. 1, 1954 Preliminary tests ORNL-1515 promising Some test data available Some test data available Some test data available Some data expected by Sept. 1, 1954 Some data expected by Sept. 1, 1954 Some data expected by Sept. 1, 1954 Some data expected by Oct, 1954 Some data expected by Dec. 1954 97 OUTLINE OF MAJOR RMR DEVELOPMENT PROBLEMS (continued) Priority Development Problem Danger coefficient Control rod effects Moderator Cooling Estimation of Heat Source Distribution Be-Na-lncone!l Corrosion Tests Static capsule tests Harp tests Al High-~temperature-differential, high-velocity loop A-1 Thermal Stress and Distortion Test with High Power Density Effects of Temperature, AT, Surface Volume Ratio, etc, 1 Creep-Rupture Properties of Inconel Under Severe Thermal Cycling Pumps Shakedown of Pumps with Face-Type Gas Seals 1 Model Tests of Full-Scale Pump 1 Endurance Tests of Full-Scale Pump 1 Fabricability of Full-Scale Pump Impeller Power Plant System Preliminary Designs Performance and Weight Estimates Effects of Temperature, Power Density, Shield Division, etc, Reactor Kinetics 1 Theoretical Analyses ARE Temperature Coefficient Measurements 1 Xenon Effects 98 Status May 1954 Some data expected by Dec. 1954 Some data expected by Dec. 1954 Good estimates made Some favorable data Some data Some data Test nearly ready to run Tests planned, data badly needed Tests planned, data badly needed Adequate data for design Tests being run Tests planned Tests planned Adequate data Adequate data for preliminary design Adequate data for preliminary design Preliminary analysis completed Tests planned Data badly needed Reports ORNL.-1517 ORNL.-1692 ORNL.-16%92 ANP-57, ORNL. - 1255, -1215, -1330, -1509, -1515, -1609, -1648 ANP-57, ORNL- 1255, -1215, -1330, -1509, -1515, -1609, -1648 ORNL CF-54-2-185 ORNL CF.53-3-231 OUTLINE OF MAJOR RMR DEVELOPMENT PROBLEMS (continued) Priority Development Problem Hydrodynamic Tests 1 Flow Separation at Core Inlet Effects of Heat Genaration in the Boundary Lavyer Fill and Drain System Preliminary Design Water and High-Temperature Tests High-{emperature Test with Radiocactive Material Full-Scale Reactor Tests Control 1 Temperature coefficient Xenon effects Stability Performance Heat exchanger, pumps, etc. Temperature distribution Endurance Tests Status Moy 1954 Reports Y-F15-11, ORNL. 1692 ORNL-1701 Some data available Theoretical analyses com- pleted for ideal case Design looks promising Tests planned for fall, 1954 Tests might be run in 1955 Some information expected from ARE Information badly needed Information badly needed Tests being planned Tests being planned Tests being planned 99