S A e - B - ORNL-2599 C-85 — Reactors-Aircraft Nuclear AEC RESEARCH AND DEVELOPMENT REPORT Propulsion Systems 7as I 3 445k 0251081 5 M-3679 (22nd ed.) L AIRCRAFT NUCLEAR PROPULSION PROJECT SEMIANNUAL PROGRESS REPORT A FOR PERIOD ENDING SEP TEMBER 30, 1958 : &é REpT Feetrrn CATION Crance: ' R R AEC 9-g-¢ < 7g OAK RIDGE NATIONAL LABORATORY operated by UNION CARBIDE CORPORATION for the U.S. ATOMIC ENERGY COMMISSION LEGAL NOTICE This report was prepared as an account of Government sponsored work. Neither the United States, nor the Commission, nor any person acting on behalf of the Commission: A. Mokes any warranty or representation, express or implied, with respect to the accuracy, ion contained in this report, or that the use of fringe completeness, or usefulness of the informat any information, apperatus, method, or process disclosed in this report may not privately owned rights; or B. Assumes cny licbilities with any information, apparatus, method, or process disclosed in this report. As used in the above, *‘person octing on behalf of the Commission® includes any employee or or contractor prepares, handl utes, or provides access to, any information pursuant to his employment or controct spect to the use of, or for damages resulting from the use of contractor of the Commission to the extent that such employe or dis with the Commission. ORNL-2599 C-85 —~ Reactors-Aircraft Nuclear Propulsion Systems M-3679 (22nd ed.) This document cansists of 232 pages. Copy ?/ of 234 copies. Series A, Contract No. W-7405-eng-26 AIRCRAFT NUCLEAR PROPULSION PROJECT SEMIANNUAL PROGRESS REPORT For Period Ending September 30, 1958 A. J. Miller, Project Coordinator DATE ISSUED JAN 22 1359 OAK RIDGE NATIONAL LABORATORY Oak Ridge, Tennessee operated by UNION CARBIDE CORPORATION for the U. S. ATOMIC ENERGY COMMISSION MARTIN MARIETTA ENERGY SYSTEMS LIBRA - i 3 445& 0251081 5 FOREWORD The ORNL-ANP program primarily provides research and development support in shielding, reactor materials, and reactor engineering to organizations engaged in the development of air-cooled and liquid-metal-cooled reactors for aircraft propulsion. Most of the work described here is basic to or in direct support of investigations under way at General Electric Company, Aircraft Nuclear Propulsion Department, and Pratt & Whitney Aircraft Division, United Aircraft Corporation. This report is divided into rour major parts: 1. Metallurgy, 2. Chemistry and Radiation Damage, 3. Engineering, and 4, Shielding. ' ANP PROJECT SEMIANNUAL PROGRESS REPORT SUMMARY PART 1. METALLURGY 1.1. Fabrication Studies of the effect of heat treatment on the mechanical properties at room temperature of arc- " cast columbium prepared by the electron-beam melting technique have indicated that specimens heated at intermediate temperatures in the range 700 to 1000°C are more ductile than those heated at the extremes of the range of treatment tempera- tures. There is evidence, however, that the in- creased ductility resulting from intermediate tem- perature anneals of welds would not be retained at higher temperatures and that stabilizing elements will be required to neutralize the effects of inter- stitial elements, Electron-beam-melted columbium billets were found to extrude readily if they were protected from atmospheric contamination during heating, Arc- melted material was found to be more difficult to extrude, possibly because of a higher impurity content, Cracks found in drawn duplex and unclad colum- bium tubing have indicated insufficient annealing before tube reduction, and therefore a high-tempera- ture, high-vacuum annealing furnace was designed and is being built, Since it appears that the production of satis- factory columbium bodies will require either high- purity metal or special alloys, purification and alloying investigations are under way. Resistance heating of columbium wire in vacuum at tempera- tures up to its melting point has shown significant reductions in nitrogen and oxygen content. Melting stock has been obtained from which special alloys will be prepared. Beryllium, zirconium, and cerium will be used to neutralize residual oxygen. High- purity columbium specimens contaminated with known quantities of oxygen or nitrogen are being prepared for corrosion testing in lithium, Methods for obtaining highly pure yttrium metal and hydrogen are being studied in order to de- termine the effects of impurities on the hydriding of yttrium. It is hoped that purification will alle- viate the tendency of yttrium metal to crack during hydriding and during thermal cycling of the hydrided metal, The electron-beam melting process was shown to be useful for removing fluorides and traces of magnesium from yttrium metal, but it was ineffective in removing traces of oxygen or nitrogen or farge quantities of magnesium. Equipment for the production of yttrium metal from YF 5 has been expanded to pilot-plant scale, 1.2. Corrosion Studies of the various methods of purifying lithium of oxygen and nitrogen have continved. As part of these investigations, samples were taken of the lithium shipments received from vendors and ana- lyzed for oxygen by several methods in order to compare various analytical techniques. The analy- ses showed that it will probably be necessary to purify all lithium received. Experiments are being conducted to determine the solubility of lithium nitride and lithium oxide in tithium, An average value of approximately 1300 pPpm oxygen was obtained recently that is consider- ably lower than the value obtained previously. The ability of yttrium to getter impurities from lithium is being investigated. Hardness increases and vacuum-fusion analysis results indicate that yttrium getters both oxygen and nitrogen from lithium, but no reductions in the oxygen or nitrogen contents of the lithium bath have been found. Static tests of several grades of arc-cast colum- bium tubing have been conducted. Preliminary results indicate that surface contamination and subsequent diffusion of the impurities into the cofumbium during annealing may account for some of the previously observed attack of columbium by lithium, Two joining techniques for columbium, braze welding and fusion welding, and three types of gaseous environment were investigated to determine their effects on the hardness of columbium and its corrosion resistance to lithium. A marked reduc- tion in depth and severity of attack was noted as the purity of the gaseous welding environment was improved. Appreciable decreases in the hardness of the as-welded columbium specimens as a result of exposure to lithium again indicated gettering of oxygen from the columbium by the lithium. A series of seesaw-furnace tests of molybdenum in contact with lithium was conducted. The results indicate, in general, that molybdenum has excellent resistance to attack and mass transfer at hot-zone temperatures as high as 1900°F, Dissimilar-metal mass transfer of Inconel to the molybdenum was observed in the Inconel-sodium-molybdenum section of the test capsule. No molybdenum failures either in the base material or in weld zones have been encountered in the tests conducted thus far, Similar seesaw-furnace tests of columbium-lithium systems have been initiated. In the one test (hot zone, 1600°F; cold zone, 1100°F; test duration, 300 hr) completed to date, columbium exhibited excellent resistance to attack and mass transfer, Static corrosion tests of beryllium specimens in lithium in iron capsules and in beryllium capsules at 1500 and at 1830°F showed beryllium to be quite resistant to attack by static lithium in an all-beryltium system. It was very heavily attacked, however, when tested in the iron container, Refractory-metal-base brazing alloys are being corrosion tested in high-temperature lithium, since commercially available brazing alloys have very limited corrosion resistance to lithium, A method was developed (under subcontract at NDA) for measuring the rates at which container metals dissolve in liquid lithium, The solution rate measurements confirmed the deleterious effects of nitrogen and oxygen and the beneficial effect of aluminum on mass transfer in lithium as noted in thermal-convection loop tests. 1.3. Welding and Brazing A survey of phase diagrams of refractory-metal- base binary systems has revealed several promising brazing alloys. Vacuum-brazing equipment has been constructed for studying these systems, and flowability and corrosion tests are under way. Studies of the diffusion of the brazing alloy into columbium are also being initiated. If diffusion or alloying occurs it may increase the remelt tempera- ture of the brazing alloy and thus increase the maximum service temperature, A small,glass dry box containing a stationary torch and a movable table was assembled for welding studies on columbium and molybdenum. This equipment will be used in investigating the weldability of these materials under highly con- trolled degrees of atmospheric purity. Several butt welds made on arc-cast columbium sheet in a puri- fied argon atmosphere were found to be ductile, vi 1.4. Mechanical Properties Equipment has been developed for tensile and creep tests of materials in controlled environments at temperatures up to 2500°F. Several successful tests have been completed on molybdenum at 2000°F. Creep and rupture properties of Inconel have been investigated in tubes with both internal pressure and axial stress, The creep data obtained compare favorably with the theoretical curves derived from Soderberg’s analysis and the von Mises and Tresca flow rules. Fracture cannot be predicted by the maximum principal stress criterion, Dynamic load tests have been conducted on Inconel rods and tubes that were programed on the basis of plastic strain rather than stress. A simple relationship, in the form N = K, exists between the number of cycles to failure and the plastic strain absorbed in each cycle, The constants a and K are found to vary as a function of the ma- terial and test conditions. On the other hand, geometry, temperature variations in the range 1300 to 1600°F, and small variations in chemical compo- sition do not significantly alter the cycle life of a material, Strains induced thermally show excellent correlation on the basis of equivalent amounts of plastic strain per cycle with mechanically induced strains in temperature ranges where the material is metallurgically stable. The frequency at which the material is strained is observed to infiuence the number of cycles that can be survived. The slow frequencies, particularly at low amplitudes, appear to produce the most deleterious results. 1.5. Ceramics Small batches of zirconium boride were synthe- sized for incorporation into BeO bodies. The ma- terial prepared by an improved process was found to be >97% pure. Several additives were tested as means for increasing the density of BeO. Of the various combinations tested, it was found that a body with 94 wt % BeQ, 5 wt % MgO, and 1 wt % B4C had the greatest density. Screening tests were run to determine the oxida- tion resistance of TaB,, TiB,, CrB2, HfBz, BN, and ZrB, when added to high-density BeO. The ZrBz-BeO and HfBz-BeO combinations were found to be superior to the other mixtures. Attempts to fabricate large hot-pressed blocks of the ZrB,-BeO mixture for which a commercial grade of ZrB_ was used revealed the need for pure ZrB.,. The commer- cial grade of ZrB2 was found to contain about 10% Fe. Uniform blocks were obtained when pure ZrB, was used, 1.6, Nondestructive Testing Studies of the use of x-ray sensitive Vidicon in a closed-circuit television system for remote viewing of x-ray images were continued. It was found that the sensitivity of the selenium photoconductor to x-rays could be improved by admitting visible light to the photoconductor surface, Contrast sensitivity was increased by as much as a factor of 3. Thick- ness changes of about 2,5% were detected in E}B-in.- thick aluminum, Investigations are being made of the applicability of existing ultrasonic techniques to the inspection of duplex tubing, It appears that the resonance ultrasonic technique can be developed as a means for evaluating the quality of the bond between the materials, 1.7. Metallography A 20-tube semicircular heat exchanger fabricated of Inconel in which a fused salt flowing around the tubes exchanged heat to NaK flowing in the tubes was examined metallographically to determine the location and effects of the leak that caused termi- nation of operation. The heat exchanger, which had operated 1139 hr at 1200°F and above, had experienced 194 thermal cycles between isothermal operation at 1200°F and operation at design condi- tions (fused salt inlet and outlet temperatures of 1600 and 1250°F and NaK inlet and outlet tempera- tures of 1070 and 1500°F.) The leak of NaK to fused salt was found to have been the result of fracture of three tubes in the high-temperature header area. The cracks were radial and propagated from the tension side of the tubes; they were in the tubes with the shortest tube bend-to-tube sheet length (]'/8 to 3/4in.). The maximum depth of corrosion attack on the fused- salt side of the tubes was 0,006 in., and, on the NaK side, it was 0.002 in. PART 2. CHEMISTRY AND RADIATION DAMAGE 2.1. Materials Chemistry Sufficient fluoride mixture for the first charge of the equipment in the yttrium-metal pilot plant was prepared. The YF, needed was prepared by direct fluorination of dry Y,0,. A mean conversion efficiency of 99% and suitably low oxygen contents were obtained. Magnesium fluoride for the process was prepared by dry hydrofluorination of MgO; the mean conversion efficiency was 97.7%. Sufficient information with which to construct a portion of the phase diagram of the LiF-YF3 system was obtained, A single eutectic occurs in the system at 19 mole % YF,; mp, 682°C., Further experimental information was obtained regarding a tentative procedure for purifying lithium metal by extracting oxide, nitride, and carbide impurities with a eutectic mixture of lithium halide salts. The second extraction of each series of extractions showed considerable reduction in the concentration of impurities. 2.2, Analytical Chemistry The method for the determination of oxygen in fluoride salts by fluorination with KBrF , was successfully applied, The coefficient of variation of the method, as based on duplicate analyses of seven samples, was 6% in measurements of the salt LiF-Mng-YFB. The recovery of oxygen added as MgO and Y203 exceeded 98%, with a precision of 2%. A rapid method was adapted for the determination of magnesium and yttrium in LiF-Mng-YF3 by titration with a standard solution of ethylene- diaminetetraacetic acid (EDTA). Amalgamated gold serves as the indicator electrode. Both ele- ments are titrated in the same solution by adjust- ment of the pH of the solution. No interferences were encountered in the application of the method to LiF-MgF ,-YF,. The precision of the method is about 1%. 2,3. Radiation Damage Experimental apparatus is being prepared with which to test the stability of yttrium hydride and beryllium oxide thermally cycled at high tempera- tures in a high gamma-ray heating region of the Engineering Test Reactor. Mockup tests have been run on the heat removal system and on the thermocouple design. The heat removal system will provide a temperature control method in that it will be possible to vary the heat conductivity of the cooling gas by varying the concentration of argon and helium in the gas mixture. Stainless steel was selected as the test capsule material as a re- sult of compatibility tests. Examinations and meas- urements of the samples will be made before irradi- ation and repeated after irradiation. Stress-rupture experiments in air on tube-burst specimens of Inconel were completed in the MTR. vii Preparations are being made for similar tube-burst tests and creep tests in the fast flux of the ORR, In conjunction with studies of the effects of radiation on semiconductor barriers, a cryostat was constructed that has a temperature range of —200°C to +350°C, Requirements, other than the tempera- ture range, were long term reliability and “‘fail _safe’’ operation. A cooling water failure or an electrical failure will shut the cryostat down with- out harmful effects to any of the system components, Temperatures in the sample chamber can be main- tained to within ]/4°C, with a gradient along the length of the copper section of the sample chamber of ]/4°C. A grown-junction silicon diode has been irradiated in a fast-neutron flux facility in the ORNL Graphite Reactor to an integrated dosage of 1.8 x 101> neutrons/cm?. Changes in forward and reverse current at l-volt bias were recorded. A change in reverse current on the order of a factor of 100 in- crease was observed that is to be compared with an increase of 200% for a comparable irradiation of an alloy-junction diode and an increase of 30% in a point-contact diode. Attempts were made to alloy indium and germanium at low temperatures, and crystal growths were pro- duced that had not been noted before. Penetration of indium into the germanium conformed to precise geometric configurations, and it is thought that such penetration may be indicative of the early steps of the alloying process. PART 3. ENGINEERING 3.1, Component Development and Testing The oil-lubricated pump rotary element that is being operated in a gamma-radiation field in the MTR canal will have accumulated a total dosage of 1 x 10'? by September 30. Operation of the seal continues to be comparable to that of a seal operated without irradiation for a similar period of time. Analyses of the bulk and leakage oil have indicated little change. A Fulton-Sylphon seal on a sump-type centrifugal pump that is circulating NaK at 1200°F, 1200 rpm, and 1200 gpm has operated satisfactorily during more than 7200 hr. A similar pump that had oper- ated in the cavitation region with NaK as the pumped fluid for 1096 hr was stopped because of a drain valve leak. No damage was evident upon disassembly of the pump, The data are being viii analyzed. A similar pump that is operating with a fused salt has logged over 10,300 hr of operation. - A third thin-shell mode| that was tested under high-temperature thermal cycling conditions is being disassembled for examination. A test loop failure caused termination of the test after 602 thermal cycles; that is, 302 more than the scheduled 300 thermal cycles. 3.2. Heat Transfer Studies Experimental studies of the effect of thermal- stress cycling on structural materials were con- tinued with the pulse-pump system, The dependence of surface failure (cracking of Inconel) on cyclic frequency was investigated in tests at 0.1, 0.4, and 1.0 cps. Maximum damage occurred at 0.1 cps, with cracks as deep as 0.207 in. being observed. At the highest frequency, no cracks were found at the end of a 100-hr (360,000 cycles) exposure. An attempt was made to correlate the results with mechanical fatigue data, and, in general, it appears that, for a given strain, failure occurs earlier (fewer cycles) with surface thermal cycling than with mechanical cycling. v Analyses of data obtained from a mercury system in the study of heat transfer in a liquid metal with internal heat generation were completed. On the average, the data indicate surface-to-fluid mean temperature differences 49% higher than those predicted by the analogy. Since the experiment was designed so that gaseous films or scale at the wall could not influence the experimental results, it is concluded that the diffusivity ratio is less than unity for mercury. Studies of heat transfer between a heated tube surface and air in high-velocity vortex motion were continued. Data with both forced- and free-vortex flow indicate values of the heat-transfer coefficient significantly above those obtained with linear flow and equal flow power dissipation in identical geometries. 3.3. Instrumentation and Controls An initial check of the data acquisition system purchased from the G. M. Giannini Company was made and minor troubles and design inadequacies were noted, The system is being modified and - should be back in operation soon. Life tests of three level probes operating in NaK were continued. Similar tests in sodium were initiated. Overrange tests designed to accelerate - the effects of aging on pressure transmitters were continued. Slight bulging of the Inconel bodies was observed after 5100 hr at 1400°F. Life tests of a series of Inconel-sheathed thermo- couples exposed to a fused salt were terminated after about 11,000 hr at 1500°F. The thermocouples were performing well when the test was stopped, but the errors were increasing. Similar thermo- couples have been under test in a sodium environ- ment for about 16,000 hr. Many of these thermo- couples now show considerable error. Studies of the fundamental behavior of thermo- couples in the range 300 to 1100°C are under way under subcontract at the University of Tennessee. Work on the improvement of the Oracle program for thermocouple-data handling is nearing completion, 3.4, Applied Mechanics Basic studies are under way of the elastic be- havior of five structural configurations commonly used in reactor design. The configurations being analyzed are a flat circular plate, a tapered circular plate, a cylindrical shell, a conical shell, and a spherical shell. The stress and displacement formulas and tables of numerical values for all the functions involved are being obtained. The study of the conical shetl, which has been completed, provides for analyses of a conical shell with a combination of axisymmetrical loads. The loads may include membrane forces and uniform pressure, in addition to edge moments and shear forces. 3.5. Advanced Power Plant Design As part of a program for examining the feasi- bility of the vortex reactor concept, experimental studies were made of the vortex strengths ob- tainable in a gas under a variety of conditions. A survey was made of various types of auxiliary power units suitable for use in satellites, A pre- liminary evaluation was made of reactor cycles employing either aluminum chloride or rubidium vapor as the heat transfer medium. PART 4. SHIELDING 4.1, Shielding Theory A calculation was performed to obtain predicted pulse-height spectra of capture and inelastic- scattering gamma rays which could be compared with the experimental spectra obtained in an ex- periment at the Tower Shielding Facility. While experimental cross sections were available for the calculation of the nitrogen capture gamma-ray spectrum, many of the cross sections used to pre- dict the spectrum of inelastic-scattering gamma rays were theoretical, In spite of this, the shapes of the calculated and experimental pulse-height spectra were in agreement, It has previously been reported that an Oracle Monte Carlo calculation of the penetration of mono- energetic, monodirectional gamma rays in a lead and water shield has included a total of 512 prob- lems. Some typical plots of the heating results from these calculations are presented as the per- centage of the total energy incident upon the slab absorbed in a specified region. The results are compared with an empirical formula, The Oracle Monte Carlo code was used to calcu- late the primary gamma-ray heating in a stab shield similar to that being used in the GE-BSF experi- ment. |t was assumed that the slab thicknesses were the same as those in the experiment and that they were infinite in the other two directions. Most of the assumptions made in the calculation caused the results to be too high, possibly as much as a factor of 1.5. This is corroborated by GE calcu- lations of the heating which were based on an analysis of the thermocouple measurements and which were a factor of 2 lower than the results of this Monte Carlo calculation, A Monte Carlo code is being developed for the IBM-704 machine to calculate the angular and energy distributions, at a point detector, of gamma rays emitted from a monoenergetic, point isotropic, or point monodirectional source embedded in an infinite homogeneous isotropic medium, The code is now in the ‘‘debugging’’ stage. Neutron dose-rate distributions beyond water slabs 1, 3, 4, and 6 mfp thick were calculated for plane monodirectional, monoenergetic sources incident on the slabs at angles of 0, 30, 60, and 75deg. The source energies considered were 0.55, 1.2, 2, 4, 6, and 8 Mev. Dose-rate buildup factors calculated for the various sources are pre- sented, as well as the dose rates at the rear of the slabs resulting from neutrons multiply scattered within the slabs, 4.2. Lid Tank Shielding Facility The final measurements taken in the L TSF ex- periment designed by GE-ANP for the study of the production of secondary gamma rays in configura- tions containing advanced shielding materials are presented. |n most of these last configurations, all of which were tested in oil, the effectiveness of various thickness of stainless steel with and without boral was investigated, One configuration that consisted only of berylium and lithium hydride in oil was also tested, A study of the production of secondary gamma rays in lead was made with lead thicknesses that varied from 1 to 9 in. The lead was followed by either an oil medium or a borated-water medium in which gamma-ray dose-rate measurements were made. From these tests it appears that the first 3 in. of lead attenvated most of the primary gamma rays from the LTSF source plate; further, the total dose rutes at fixed distances from the source were not affected when the lead thickness was increased beyond 3 in, The measurements in the borated water were a factor of 3 lower than the dose rates in the oii in the region close to the lead and were a factor of 13 lower approximately 120 cm beyond the lead. A corrected curve of thermal-neutron flux meas- urements in oil at the LTSF is presented. It is 16% higher than the curve presented previously. Shields consisting of randomly distributed ab- sorbing chunks in a relatively transparent matrix: must be from a few per cent to several hundred per cent greater in mass than homogeneous shields which give the same amount of attenuation, This is the result of radiation ‘‘channeling’’ in the spaces between the absorbing chunks. A method for calculating the transmission of radiation through heterogeneous shields has been developed, and a typical calculation for boral has been performed. The computed results are compared with experi- mental results. 4.3. Bulk Shielding Facility The planned series of measurements for the GE- BSF study of the production of heat in radiation shields was completed, but the results are of questionable value since it has been discovered that all of the heating samples leaked and oil entered the air region surrounding the samples. The series, which consists of configurations of beryllium and lithium hydride separated by a gamma-ray shielding section (iron, lead, or Mallory 1000), is being repeated with new samples. 4.4. Tower Shielding Focility The experiment performed at the Tower Shielding Facility in cooperation with Convair, Fort Worth, to obtain information complementary to that obtained from the Nuclear Test Airplane (NTA) program has been completed. During the NTA program of ex- periments, gamma-ray and fast-neutron dose rates and thermal-neutron fluxes were measured both inside and outside the airframe containing the reactor while in flight and on the ground. Measure- ments were also made on the ground in the absence of the airplane structure. This last set of measure- ments was duplicated in the TSF experiments, and, in addition, measurements were made as a function of altitude in the absence of the airplane structure. With these additional measurements the influence of the air, the ground, and the aircraft structure on the various measurements can be determined. The TSF experiment also included angular mappings of the radiation around the various reactor shield configurations, near the ground and at altitude, to obtain data which will yield more accurate source terms than were previously available. Some gamma- ray and neutron energy spectra were also determined. Predictions of the gamma-ray and fast-neutron dose rates at large distances from the Tower Shielding Reactor |l (TSR-Il) have been made for two operating conditions. For the first case the reactor was assumed to be bare and to be operating at a 5-Mw power at an altitude of 200 ft, a condition which would result in the highest dose rates that could be achieved with the reactor, The gamma-ray dose-rate calculations for the bare reactor were based on measurements taken at large distances from the TSR-|, while the fast-neutron dose-rate calculations were based both on the TSR-| data and on other measurements made at large distances from the Aircraft Shield Test Reactor (ASTR). For the second case the reactor was assumed to be encased in a “‘beam’’ shield which would allow a highly collimated beam of radiation to be emitted from the reactor. This calculation, which was based on measurements made at large distances from the TSR-l in a simulated beam shield, is con- sidered to be more representative of a typical operation. The results for both cases are plotted as a function of distance from the reactor. As part of this investigation, a calculation was made of the uncollided neutron flux 4200 ft from the TSR-I to obtain a fast-neutron buildup factor for the case Fad in which an air-filled collimator extending through the reactor shield is pointed at the detector, 4.5, Tower Shielding Reactor Il The latest design of the TSR-Il, with its asso- ciated controls and 5-Mw water cooling system, and several studies supporting the design are pre- sented. The critical mass has been set at 8.1 kg with 1,6% excess reactivity. Control will be achieved by moving six umbrella-shaped grids of Inconel-clad cadmium in the internal water re- flector, A description of the first shield is also presented, CONTENTS SUMMARY ...ttt ettt e et et ee e e e e e e e es o2 s ee ettt ee s e et e PART 1. METALLURGY LTe FABRICATION ..ottt et e et ee et seee e e e es e ses e eee oo, Columbium INVeSH GAHONS ..cvee ettt ettt ee e e et e e s e s s ee s et sese e Effect of Heat Treatment on Mechanical Properties w.oo.uooooioooeeeeeeeee oo EXTPUSION STUAIS ettt et e oo et e e e s e s ne e o PUrification EXPErimMeNts ....ooiuiuieiiiiiiieceeeeeeeee e ee et e e e e e e e oot SPECIal Alloy Preparations ...t e eeee et s et e eeseese s e s s e e reseseeseanes Effects of Oxygen and Nitrogen Contamination on Corrosion by Lithium.....ccocvueeeuemrervernnn, Yttrium and Yttrium Hydride Investigation S ... iieeciieoe oo eeeee e e e v s es s eses s see s ereras e e Effect of Metal Purity on Hydriding .cooovvvoee oot s se e Dissociation Pressure STUdies ...t ettt ettt er e eee e aeeeesaeas Deformation Mechanisms of YHIIUM wo.ociiiiccc ettt et es et seet e e reas Electron-Beam Melting of Yttrium and Yttrium-Magnesiom Alloy......ocoocoieeieeieeeeeeee s Expansion of Ytrium Preparation ProCess ..o et e e ee e eseeeeve e e e e, 1.2, CORROSION L.ttt ee et et e et ese s ees eeeseasesetessasasen s ssnsessesessseasas Lithium PuUFHiCation ..ot et ne st ee ettt s s essaeeeeemessssemeesraenas Columbium Exposed 10 Lithium ..ottt e et e s e e s e s e sres e s s sene e Arc-Cast Columbium Tubing Tested in Static Lithium .......cccoeiviiiioecieece e Effect of Welding Procedures on Corrosion ca.. ittt et eva e tee et e Comparison of Methods for Determining the Nitrogen Content of Columbium .........o.ocovvvvuerieernneee. Molybdenum and Columbium Tested in Lithium in Seesaw-Furnace Apparatus .cccoceeeveereerveenennee. Berylium Exposed 10 Lithium .o et eeee e eee s et e e eae e et areaeesnaeas Refractory-Metal-Base Brazing Alloys Exposed to Lithium .....occooiiiiiiiiiieiieeeee e, Determination of the Solution Rate of Metals in Lithium ....o...occcooiiiiiiieciceeee e 1.3, WELDING AND BRAZING ..ot ettt ire st e sre e ba b ans v e sae st eatabssassbasssbssesnsaaensarens Development of Brazing Alloys for Lithium Service........cocooiiiioiiiiei e, 1.4. 1.5. Welding Studies of Columbium MECHANICAL PROPERTIES and Molybdenum ..o .................................................................................................................. Development of Test Equipment for High-Temperature Investigations ...........cccccocoovrieviccivieciiecenaa. Multiaxial Creep Studies.......... Creep Analysis ................ Rupture Analysis................ Strain-Fatigue Studies.............. Temperature Dependence .. Frequency of Dynamic Loa Design Factors ....ccoeueneee. .................................................................................................................. ------------------------------------------------------------------------------------------------------------------ .................................................................................................................. .................................................................................................................. .................................................................................................................. dS ............................................................................................................ .................................................................................................................. .................................................................................................................. .................................................................................................................. VOO NN O WWW W xiii 1.6. 1.7. 2.1 2.2. 2.3. 3.1 3.2, xiv Densification of Beryllium Oxide ..ottt Oxidation Resistance of Boron-Containing Beryllium Oxide Bodies ...c.ccocimiiiiiciiciiiiiiiiiininnne, NONDESTRUCTIVE TESTING oottt et s seere s s e a et e e sarene san e snb bbb b ns Remote XeRay VieWing ....ccooiiiiiiiiiiriniis ettt e s st st b ees s aaa b e n e e ae b e s abesinness DUPLEX TUBING e bbb s e bbb e b s Metal [dentification METEr .. ...cocuueiiiiiiiete et eeeetes bttt e ssees s aans e eeses s besbra s e s bt ae e s e rbaras s sanean treneaeares METALLOGRAPHY Lottt es e et s ans sha s ek bbb e e s s s s e a st st e Results of Metallographic Examination of Small Semicircular Fused-Salt Fuel-to-NaK Heat Exchanger Operated at High Temperatures ..o PART 2. CHEMISTRY AND RADIATION DAMAGE MATERIALS CHEMIS TRY ittt e et st st ere e e ea e e se bbb b s b ibe s e e eas s Preparation of Charge Material for Reduction to YHIUM .cccoooiiiiiiiiieiie e Conversion of Y0, 10 YF 4o Preparation of MgF , ..o Preparation of the LiF-MgF,-YF, Mixture ..o Phase Equilibria in the System LiF-YF ..o s Extraction of Lithium Metal Impurities with Molten Salts ..o ANALY TICAL CHEMISTRY .ottt ettt sttt re et sbe st s srs et e sasasess e s e e e nseasnesmsnbessssnsesias Determination of Oxygen in Fluoride Salts ... s Determination of Yttrium and Magnesium in LiF-Mng-YF3 ................................................................ RADIATION DAMAGE ...ttt ettt e e te s te st e s e e e s e st e e eabe st e e eaeeneesebeesesenbnansaara s ETR Irradiation of Moderator Materials for Use at High Temperatures ..., Creep and Stress Rupture Tests Under lrradiation ..o Radiation Effects in Electronic Components ... it et te s e e meanse st aeesessssas Wide-Range Multipurpose Cryostat.. ..ottt e Grown-Junction Silicon Diode lrradiation ........cooieiieiieieee et s e Low-Temperature Indium-Germanium AllOy .......ccooiiviiiiiiic e PART 3. ENGINEERING COMPONENT DEVELOPMENT AND TESTING ..ccoiiieieeciitceiee et snen s s Irradiation Test of Oil-Lubricated Pump Rotary Element ..o, Test of NaK Pump with Fulton-Sylphon Seal.......cccooiiiiiiircr s Cavitation Tests of Centrifugal PUmps ..ot Thermal Stability Tests of Metal Shells. ... HEAT TRANSFER STUDIES w..oooovoc e SO USSR Studies of the Effect of Thermal-Stress Cycling on Structural Materials ... Liquid-Metal Heat Transfer Experiment ..o, Heat Transfer with Vortes Flow .ot tetaee s e e st sen e senneaeans 3.3. 3.4. 3.5. 4.1. 4.2. INSTRUMENTATION AND CONTROLS 94 Data ACQUISITION Sy STEM ..ottt e 94 Liquid-Metal-Level TranSducers .......cooovuomoooeeeoeoeeoeeoeee e 94 Pressure Transmitter TestS ..ot ee oo e 94 ThermoCOUPIE TeSES ........oiiiuieriirn ettt ettt ee et ee et eoee e e oo e oo eseee e 95 Thermocouple Development STUdIES ............cooemiivviieeoeeeeeeseeeeeeree oo oo oo 95 Behavior of Thermocouples in the Temperature Range 300 t6 1100°C weeveveveoeoooeeoeeooo 95 Oxidation and Cold Work STUdies .....oiuueieeeeceeeeeeeese e eeeeee e 96 Calibration SHUAIES ..ot ettt ee e et eeee e 9% Oracle Program for Thermocouple-Data Handling ..o eeeeeoeevoeooeeoeoeeoeeeoeeoeeoeeeeeeeee 96 APPLIED MECHANICS. . ..ottt es e e 98 Basic Problems in EIastiity oiiuiiriiiiocececeeceeee e et 28 NUMEFICAl ANGIY SIS ©oiviuitiiteicecececet ettt 99 ADVANCED POWER PLANT DESIGN .. ..ot 101 Vortex Reactor Experiments.. ...cooocoviiuiiiiiiieec oo e e et 101 Survey of Design Problems of Auxiliary Power Units for Satellites ..oooomommoooooeoeoooo, 110 Cycle Performance Considerations ..o oo oo 110 PLANNE PROGIAM ..ottt e s et e e e e et 113 PART 4. SHIELDING SHIELDING THEORY ..ottt st ee e aes s ee s e e e s ae sttt s s eee oo eeoe e 119 Analysis of Neutron Physics Division Experimental Study of Gamma Rays Produced by Neutron Interaction i AQr ......o.ooeeeoeee oo e 119 Spectra of Gamma Rays Incident on the Outside of the Collimator .....cocooveeveeeeereveseereereorrorans 121 Spectrum After Attenuation and Buildup in the Water Collimator.o.ovimmmeeoooooeooooo, 122 Calculation of Pulse-Height Distributions from the Detector ... imreoneeeeroseeeeesreseesress oo 123 Monte Carlo Calculation of the Deposition of Gamma-Ray Heating in Stratified Lead and Water SIabs ..o ettt e ettt 123 A Monte Carlo Calculation of the Gamma-Ray Heating in Several Shielding Configurations Adjacent to the Bulk Shielding REQCIOr ...oc.oveiiereeereeeeeeeeeeeee e e ees e e 128 MEthod Of Caleulation ..ocoiiiiiiiii et v e eee e et e est e s eeesenee s seeons 128 CONCIUSTONS .ttt oo es e e e ee e e te et e e s s s e s e oot ee s 131 A Monte Carlo Code for the Calculation of Deep Penetrations of Gamma Rays ......cccooeevveevrennan., 133 IMPOrtanCe SAMPIING ...ttt enesesese e ne s rasesans 134 Double Systematic SAMPling .....ciiiieiiceeeeee e ettt ee s 134 Statistical EStimation ..ottt et e e 135 Weighted ISotropic SCattering ..ottt et et e s oo 135 Splitting and Russian RouUlEtte .. ...ttt st enan e 135 DUBPUL _in.-ID MULLITE TUBE 2-in, VACUUM IORE Rt L E nnnoonnéoo w s b P b g = |INDUCTION COIL 4] T C D 2 o O m o o D D D P D 4-in, DIFFUSION PUMP - D e - QUARTZ TUBE L | SRR ( QQOOOOOOOOODOO‘OODOO [ ";\—GRAPHiTE SUSCEPTOR J o) FORE PUMP \\ — b 5~—— EXTRUSION BILLET ] 7 b . L/ - BRASS ENDPLATE AN /[ A\ Fig. 1.1.1. Layout of High-Temperature High-Vacuum Annealing Furnace. ANP PROJECT PROGRESS REPORT ORNL-LR-DWG 33725 COLUMBIUM OR MOLYBDENUM) V. WELD \ S5 e Fig. 1.1.2. Billet Design for Extrusion of Stainless- Steel-Clad Refractory-Metal Tubing. WELD ) 7 N Purification Experiments D. O. Hobson The possibility of purifying columbium by anneal- ing at high temperatures in vacuum is being investi- gated. A vacuum system is being used in which a columbium wire can be resistance heated to temper- atures up to its melting point, 2415°C. The results of one series of tests for 30-min periods at various temperatures and pressures are presented in Fig. 1.1.3. In future experiments, the effect of varying H. Inouye the heating time at constant temperatures will be studied. Special Alloy Preparations T. K. Roche Special columbium alloy specimens are being pre- pared for evaluation in molten lithium. The avail- able evidence indicates that the mechanism of corrosion of columbium by lithium is related to the presence of oxygen in the “‘pure’’ metal. The addition of alloying materials that are potentially capable of neutralizing the residual oxygen should therefore be beneficial, The alloy additions pres- ently being considered are beryllium, zirconium, and cerium, Columbium melting stock has been pre- pared from sheet material received from DuPont, which has been found, by chemical analysis, to contain 0.032 wt % C, 0.0003 wt % H,, 0.0084 wt % N,, and 0.053 wt % O,. Prior to the preparation of the alloys, the fabricability of the unalloyed ma- terial melted by normal arc-melting procedures will be evaluated. UNCLASSIFIED ORNL-LR-DWG 33726 PRESSURE {mm Hg) (x10°3) 15 15 24 2.4 28 a a5 400 : IMPURITY CONCENTRATION (ppm)} of S —— — ROOM TEMP 1400 1500 1600 1700 1800 1300 2000 (AS RECEIVED) ANNEALING TEMPERATURE (°C) Fig. 1.1.3. Effect on Impurities af Heating Columbium Wire in Yacuum for 30 min at Yarious Temperatures and Pressures. (@OWiamyahmisianaayetiom Effects of Oxygen and Nitrogen Contamination on Corrosion by Lithium J. E. Spruiell H. Inouye Experimental studies of the absorption of oxygen and nitrogen by pure columbium in the temperature range of 800 to 1600°C at pressures ranging from 3x 104 to 1 x 10~ mm Hg were initiated. The preliminary tests of nitrogen pickup consisted of (1) sealing specimens of columbium and hot-pressed boron nitride or hot-pressed silicon nitride in quartz capsules and heating for 20, 50, and 75 hr at temperatures of 900, 1000, and 1100°C; and (2) in- troducing purified nitrogen of a given partial pres- sure into a previously evacuated furnace which contained columbium. The results of the tests in which the columbium specimens and boron nitride or silicon nitride were sealed in quartz capsules were evaluated in terms of weight change and change in chemical analysis. VSummer employee. The weight-change measurements indicated that the columbium specimens had absorbed some gas, but chemical analysis proved the absorbed material to be primarily oxygen rather than nitrogen. The data for one columbium specimen sealed with boron nitride under a pressure of 3 x 10~ mm Hg in quartz are presented in Tables 1.1.3 and 1.1.4, Since the change in weight indicated a total im- purity pickup of only 0.09%, it appears that the analyses are somewhat incorrect, but they do indi- cate the relatively higher pickup of oxygen than of nitrogen. Eleven specimens were given similar tests, and all the results were of the same order of magnitude. The hot-pressed boron and silicon nitrides apparently contained decomposable oxide impurities. In a second series of tests, columbium speci- mens were heated in a purified nitrogen atmosphere in a furnace at temperatures up to 1300°C for 1 hr. No significant weight changes were observed. In these tests nitrogen pressures from 0.5 to 3 x 10~3 were used. The specimens were bright after the tests, but they reflected light in @ way that re- vealed thin transparent films on their surfaces. The effect of the pressure of the nitrogen on film formation has not been studied completely, but it is Table 1,1,3, Weight Change Data for Pure Columbium Exposed to Boron Nitride in an Evacuated Quartz Capsule for 72 hr at 1000°C Columbium Boron Specimen Nitride Original weight (g) 2,8825 0.7263 Weight after test (g) 2,8853 0.7134 Weight change (g) +0.0028 ~0.0129 Table 1.1.4. Results of Impurity Analyses of Columbium Before and After Exposure to Boron Nitride in an Evacuated Quartz Capsule for 72 hr at 1000°C Columbium Impurities (wt %) Specimen 02 N2 H2 As-received 0.0061 0.0046 0.0005 After test 0.12 0.017 0.0046 PERIOD ENDING SEPTEMBER 30, 1958 believed to be small. There may be a critical pres- sure below which absorption occurs without the formation of the nitride film. None of these speci- mens were analyzed because the weight-gain data indicated no measurable absorption of any gas. Furnaces and equipment capable of operating at much higher temperatures are being fabricated so that the range of study may be increased. Oxygen-contaminated specimens were made by introducing purified oxygen into a previously evac- uated furnace containing columbium specimens. The gas was introduced in such a manner as to maintain a constant partial pressure of oxygen. Various concentrations of oxygen were obtained by leaving the specimens at temperature in the oxygen atmosphere for various times. By this method, specimens containing from 61 to 5000 ppm oxygen were prepared by heating at 850°C under a partial pressure of 2 x 10~ mm Hg O, for times ranging up to 16 hr, These specimens are being tested to determine the effects of oxygen contamination on corrosion of columbium by lithium. YTTRIUM AND YTTRIUM HYDRIDE INVESTIGATIONS W. J. Werner T. Hikido Effect of Metal Purity on Hydriding It has been reported by GE-ANPD that con- siderable difficulty has been encountered with re- spect to cracking of hydrided pieces of yttrium. The cracks form both during hydriding and upon thermal cycling of the hydrided piece. There has been some evidence that the cracking is associated with nonmetallic inclusions which may be observed in the microstructure of the metal, and it is hoped that improvement of the purity of the yttrium will alleviate the cracking problem. Conclusive proof of the effect of impurities is still unavailable, how- ever, because of the lack of high-purity metal. At the present time, studies are under way on the zone refining of yttrium by the floating-zone procedure developed by Kneip and Betterton 2 for the zone refining of zirconium, Work has also been started on the evaluation of various procedures for the purification of hydrogen. With pure metal and pure 2G. D. Kneip, Jr,,and J, O, Betterton, Jr., J. Electro- chem, Soc. 103, 684 (1955). ANP PROJECT PROGRESS REPORT hydrogen, it will be possible to compare the prop- erties of a high-purity hydride with the properties of hydrides having various impurity contents, Dissociation Pressure Studies No accurate set of dissociation pressure curves for the YH system is, as yet, available. A dis- sociation pressure rig is therefore being constructed, and measurements will be made of as pure a hy- dride as is obtainable. The effects of impurities on the dissociation pressure will also be determined. Deformation Mechanisms of Yttrium The growth of single crystals of yttrium is being attempted in connection with a program to study the deformation mechanisms of high-purity yttrium. The first method to be tried for the growth of these single crystals is that of growing the crystals from the melt (zone refining). A strain-anneal technique will also be tried. Electron-Beam Melting of Yttrium and Yttrium=Magnesium Alloy W. J. Werner T. Hikido The Temescal Metallurgical Corporation, Rich- mond, California, has developed a technique for melting metals under a high vacuum in which the energy of a focused electron beam is used. The furnace consists essentially of a vacuum pumping system, an electron **qun’’ and power supply, a melting-stock feeding device, and a water-cooled copper crucible with an ingot extraction mechanism. The electron gun has a circular tungsten filament electron source, focusing devices, and a power source which supplies a high-potential electric field through which the electrons are accelerated, The electron gun is designed to concentrate part of the electron beam on the end of the melting stock as it is lowered into the furnace and the balance on the surface of melted metal ingot in the crucible. The metal melted from the stock drips into the crucible where a molten pool is maintained and the ingot is continuously extracted. The major ad- vantages of the Temescal process are (1) the high vacuum (104 mm Hg) maintained during melting and (2) the use of water-cooled copper to minimize crucible contamination, The electron-beam melting technique was used to investigate the possibility of purifying yttrium metal with respect to O, N,, Li, Mg, and C content, and the possibility of consolidating yttrium-magnesium alloy directly into a pure metal ingot without going through the vacuum-distillation step, Calculations by A. R. Miller® showed that it should be thermo- dynamically possible to purify yttrium by melting it under a high vacuum. The consolidation of the yttrium-magnesium alloy was also thought to be possible if the melting were done by using a slow feed rate and a low power input to the electron gun. For the experimental investigation, one bar of yttrium metal made up of six finger castings Heliarc welded together and two bars of yttrium-magnesium alloy ~ 1Y% in. in diameter and 10 in. long were sent to the Temescal Metallurgical Corporation. The bar of yttrium metal was double melted. The first meit was accomplished at a feed rate of ~ ]/3 Ib of metal per hour with maximum power input to the gun. This melt was characterized by a heavy purple plasma which surrounded the electron gun, metal stock, and the melt, The density of this plasma, which was composed chiefly of lithium and mag- nesium, controlled the melting rate for a given power input. At too fast a melting rate, the plasma be- came dense enough to short out the filament to the focusing device. The plasma was considerably less dense during the second melt, and a faster melting rate, was, therefore, possible. The original analyses of the six finger castings and the analysis of samples taken from the bottom of the first melt and from the top, middle, and bottom of the second melt are presented in Table 1.1.5. The results indicate that the electron-beam melting did not purify the metal with respect to oxygen and nitrogen and that the melting may have actually resulted in some additional contamination. There did seem to be some degree of purification between the first and second melts, however, and the method should be very effective in removing magnesium, lithium, and any volatile fluorides. Therefore, consideration will be given to further experiments with electron~beam melting when larger quantities of metal become available., Even if re- moval of the oxygen and nitrogen does not prove to be possible, it will be of considerable interest to minimize the other impurities in order to determine the true effect of the gases. The fluorides and oxyfluorides are currently suspected of contributing heavily to the difficulties encountered in working yttrium metal and with cracking in hydrided yttrium. 3A. R. Miller, A Thermodynamic Study of Metals and Ceramics, report prepared for Stauffer Chemical Company, PERIOD ENDING SEPTEMBER 30, 1958 Table 1.1.5. Chemical Analyses of Yttrium Stock and Melts Prepared by Electron«Beam Melting Techniques at the Temescal Metallurgical Corporation Impurities (ppm) Casting No. 0, N, Mg Li C Yttrium Stock L.9B-1 1300 1100 27 50 66 L-11A-3 990 48 25 16 L-T1A-4 970 30 23 16 L-10A-2 980 170 29 410 L.10A.7 980 100 100 <10 140 L-10A-8 850 110 120 <10 150 Melt T-1B (bottom of first melt) 1600 330 <10 T-2T (top of second melt) 1500 210 <10 T-2B (bottom of second melt) 1300 170 <10 T+2M (middle of second melt) 1100 220 <10 Attempts to process the 80 wt % Y~20 wt % Mg alloy directly to the metal ingot by the electron- beam technique were unsuccessful. It was thought that if a low melting rate were maintained and the power input ke pt low, the magnesium could be driven off and the metal melted. However, when the beam of electrons was foacused on the-alloys the++ magnesium boiled out violently with considerable splattering. Some of the magnesium deposited on the filament and focusing shield and caused short circuiting of the filament. It was not possible to control the splattering and shorting, and therefore the melting attempts had to be abandoned. Expansion of YHrium Preparation Process T. Hikido In March 1958, the Oak Ridge National Laboratery was requested to consider scaling up the yttrium preparation process, which was described in the previous report.* Official approval of the proposed ORNL program was received at the end of April, and the preparation of facilities was begun. The objective set forth for the program was the prep- aration of ytirium metal containing less than 1000- ppm O, in 20-Ib ingots. The equipment for this program, which is installed in Building 9201-2, Y-12 area, consists of two parallel 50-1b fluoride purification systems that supply YF, mixtures to a National Research Corp- oration 150-lb (rated in steel) vacuum-induction- melting furnace. A system for transferring into the furnace the lithium used as the reductant is included in the installation. The control panel and furnaces for one of the fluoride purification systems may be seen in Fig. 1.1.4 beside the vacuum-reduction furnace, which is in the center of the photograph. The lithium and fluoride loading ports on the side of the vacuum furnace may be seen. Another view of the reduction furnace which shows the control panel for the 100-kw motor-generator reduction power supply is shown in Fig. 1.1.5, 47, Hikido and W. J, Werner, ANP Quar, Prog. Rep. March 31, 1958, ORNL-2517, p 4. ANP PROJECT PROGRESS REPORT Fig. 1.1.4. Yttrium Preparation Facility. ekttt 10 PERIOD ENDING SEPTEMBER 30, 1958 Fig. 1.1.5. Yttrium Preparation Facility and Control Panel for Power Supply. (Queesiemsasaiaiesstie - ANP PROJECT PROGRESS REPORT 1.2. CORROSION E. E. Hoffman LITHIUM PURIFICATION E. E. Hoffman Various methods have been investigated for the purification of lithium.! Although the effect of nitrogen and oxygen contamination of the lithium on the corrosion resistance of columbium to lithium at elevated temperatures is not yet well understood, it is generally felt that oxygen and nitrogen concen- trations should be reduced to a practical minimum to eliminate the purity of the lithium as a test vari- able. A method of reducing the oxygen contami- nation is also needed in order to produce the low- oxygen-content lithium required in the production of low-oxygen-content yttrium (see Chap. 1.1, this report). Analyses of lithium from three shipments supplied by the vendor in gas-tight stainless steel containers (under an inert gas) are presented in Table 1.2.1. The specification of less than 100 ppm nitrogen was, as may be seen, not met by the vendor. Ef- forts will be made to obtain improved as-received material, but it is felt that it will probably be nec- essary to purify all lithium used for tests. Three different methods have been used for de- termining the oxygen content of lithium. In order to provide a comparison of the results obtained by the three methods, dip samples were taken from approxi- mately 25 different batches of lithium. These dip sampdes weresthenssuwbmimtedalor analysis by each TE. E. Hoffman, W. H. Cook, and D. H. Jansen, ANP Quat, Prog, Rep. March 31, 1958, ORNL-2517, p 9. Table 1.2.1. Nitrogen and Oxygen Analyses of Recent Lithium Shipments Impurity Content Shipment Quantity (ppm)* No. {Ib) Nitrogen Oxygen 1 20 605 390 2 46 1940 415 3 46 3310 1980 *Each value is an average of four analyses; sample were taken for analysis at approximately 300°C. 12 method: (1) an activation analysis method de- veloped in the ORNL Analytical Chemistry Di- vision, (2) a butyl iodide-iodine method also de- veloped in the ORNL Analytical Chemistry Di- vision, and (3) a methyl alcohol-Karl Fisher reagent method developed by the Nuclear Development Cor- poration of America. Although the agreement of the results obtained with the three methods was not good, the best agreement was obtained between the butyl iodide-iodine method and the activation analysis method. Improved sampling and analytical techniques will be employed in the future to further determine the reliability of these two methods for determining oxygen in lithium. Results of preliminary determinations of the solu- bility of lithium nitride and lithium oxide in lithium were presented previously,] and one additional oxide solubility determination has been made for which an improved sampling technique was used. A total of five analyses made by two methods by the Analytical Chemistry Division on samples taken at 250°C showed good agreement. An averaged value of approximately 1300-ppm oxygen was measured. This value is considerably below that obtained in one previous test, and additional work will be re- quired to resolve the difference. The purification experiments, described previ- ously,! in which titanium and zirconium metals were used as gettering agents for nitrogen and oxygen in lithium, indicated that although fhesg metals are effective in removing nitrogen, they do not getter oxygen from the lithium. In several tests the oxygen content of the titanium getter foil actually decreased, as would be expected, inas- much as lithium oxide is more stable than titanium oxide. Yttrium oxide is much more stable from a free energy of formation standpoint than either ti- tanium or zirconium oxides, and, therefore, tests are being performed to determine the effectiveness of yttrium metal as a getter for oxygen in lithium. A gettering test was conducted in which 88.5 g of yttrium turnings was added to 628 g of lithium, The stainless steel container, which had a tan- talum liner, was held at 1500°F for 72 hr. Yttrium- metal cylinders (0.3 in. in diameter and 2 in, in fength} introduced into the purification vessel at 24-hr intervals were analyzed for oxygen and nitro- gen at the conclusion of the test. The results of PERIOD ENDING SEPTEMBER 30, 1958 the analyses, which were made by the vacuum- and further tests will be conducted to check the fusion method, are presented in Table 1.2.2. usefulness of yttrium for this application. Hardness measurements were made on the yttrium metal specimens tested in lithium and on similar COLUMBIUM EXPOSED TO LITHIUM specimens given the same heat treatment in an E. E. Hoffman evacuated capsule. The purpose of these measure- ments was to determine the extent of gettering of Arc-Cast Columbium Tubing Tested in Static oxygen and nitrogen from the lithium by the yttrium, Lithium Pickup of these elements results in increases in Static corrosion tests of the various grades of bardness in yttrium. The hardness increases which columbium stock received from vendors are being occurred in one of the yttrium cylinders are given performed. Specimens are exposed to lithium for in Table 1.2.3 as a function of depth from the sur- 100 hr at 1500°F and at 1800°F in an effort to face. It may be seen that an appreciable hardness correlate metal history and analysis with any increase occurred to depths of 65 to 105 mils. observed corrosion. Only a few samples have been Specimens No. 2 (Y-6) and No, 3 (Y-8) showed tested thus far, and there are, as yet, no clear cut similar hardness increases near the surface but to quantitative indications of the effects of the nitro- lesser depths, as would be anticipated, since they gen and oxygen content of the metal on corrosion. were exposed to the lithium for shorter times. The tests have shown, however, that the presence Unfortunately, these promising results were not of oxygen is detrimental to corrosion resistance, accompanied by reductions of the oxygen content In general, specimens of welded tubing have been of the lithium, Lithium samples taken at various heavily attacked in the weld and heat-affected time intervals showed a continuing increase in zones, and there has been little or no attack of the nitrogen and oxygen as the test run proceeded. |t base material, is felt, however, that the yttrium metal analyses The effect of surface contamination introduced are less subject to error than the lithium analyses; during processing is itlustrated in Fig. 1.2.1, which Table 1.2,2, Analyses of Yttrium Metal Getter Specimens Before and After Exposure to Lithium at 1500°F Specimen Time Period of Exposure Oxygen Content (ppm) Nitrogen Content (ppm) No. (hr) Before Test After Test Change Before Test After Test Change 1(Y-3) 0-72 1300 3800 +2500 450 1100 + 650 2 (Y-6) 24-72 1400 3900 +2500 570 850 + 280 3(Y-8) 48-72 240 1800 +860 110 170 +60 x . . ] Table 1.2,3. Diamond Pyramid Hardness Values for an Yttrium Cylinder Specimen (Y-3) Used in a LithiumYttrium Gettering Experiment as a Function of Distance from Surface Exposed to Lithium Compared with Values for a Specimen Given a Similar Heat Treatment in Vacuum Distance from Surface (mils) Specimen Treatment 5 (Edge) 25 65 105 145 (Center) Diamond Pyramid Hardness (500-g load) Exposed to lithium for 72 hr ot 1500°F 117 95 71 66 59 Heat treated as above but in vacuum 59 60 58 * 59 59 rather than in fithium ANP PROJECT PROGRESS REPORT UNCLASSIFIED] OUTER SURFACE OF TUBE WALL £ © =3 s} S UNCLASSIFIED Y-26263 0.015 in. ~=—— INNER SURFACE OF TUBE WALL Fig. 1.2.1. Outer and Inner Surfaces of Tube Wall of DuPont Arc-Cast Columbium Specimen Following Exposure to Static Lithium for 100 hr ot 1500°F. Both sides were exposed to lithium. Note attack on outer surface. Etchant: HF-HN03~HZSO‘-H20 (parts: 25-10-10-55). 250X. ((umeeemmming ) shows the outer and inner surfaces of a specimen Table 1.2.4, Diamond Pyramid Hardness of DuPont of DuPont arc-cast columbium tubing that was ex- Arc-Cast Columbium Tubing Before and After Exposure posed to lithium for 100 hr at 1500°F. Since both to Lithium for 100 hr at 1500°F surfaces were exposed to a common lithium environ- S A ment, the lithium purity was eliminated as a vari- oloes BNy low able. It appears that some contamination was intro- Near Outer Middle of Near Inner duced during tubing fabrication and was diffused Surface Wall Surface into the outer surface during a subsequent anneal. Hardness measurements also indicated the possi- Before test 145 121 133 bility of outer surface contamination. The Diamond Aot et 109 % 18 Pyramid hardness values for this arc-cast tubing before and after the test are given in Table 1.2.4. *Values are averages for four determinations. The decrease in hardness may have been due to leaching of impurities by the lithium, Effect of Welding Procedures on Corrosion A series of tests was conducted to study the ef- fect of various Heliarc welding procedures on the corrosion of arc-cast columbium in lithium. Two joining techniques and three types of welding en- vironment were used in weld sample preparation. The arc-cast columbium stock used was 0,035 in. in thickness, and it contained, as determined by chemical analysis, 360-ppm 0,, 20-ppm N,,, and 140-ppm C. The joints were made by using a fusion welding technique in which no filler metal is applied or with a filler metal consisting of an 85% Zr-15% Cb alloy. Sheet material for the fusion welds was ob- tained by splitting and flattening 0.5-in.-OD seam- less tubing. For the welds made with filler metal, the split tubing was not flattened. Three joints of each type were made that were prepared in three different environments. One was made in air with good, conventional inert-gas (argon) coverage; one was made in an inert-atmosphere chamber containing argon gas contaminated with 1600 ppm N, + CO and 400 ppm 02; and the third was made in an inert- atmosphere chamber containing high-purity argen, which analyzed 67 ppm N, + COand 3 ppm 0,. All the specimens were exposed to lithium simultane- ously in a single columbium container for 100 hr at 1500°F. The results of these tests are presented in Table 1.2.5. The specimens joined in argon plus air showed weight losses and those joined in the inert atmosphere chamber showed weight gains, The specimens joined with the filler metal showed es- sentially no attack of the weld metal. The effects of environment on the heat-affected zone and the weld metal are illustrated in Fig. 1.2.2. As may be seen, the heat-affected zone of the columbium ad- jacent to the braze weld of the specimen joined in a contaminated environment (air plus argon) was heavily attacked, but the same zone showed very little attack when the specimen was joined in a high-purity argon environment. The base material was attacked in both cases but to a lesser degree in the case of the specimen joined in high-purity argon, 2' PERIOD ENDING SEPTEMBER 30, 1958 The specimens joined by fusion welding were heavily attacked in both the weld zone and the heat-affected zone. The attack in these areas de- creased in both depth and concentration as the purity of the cover gas increased. The weld zones are shown in Fig. 1.2.3 after exposure to lithium at 1500°F for 100 hr, The attack was extremely heavy on the specimen welded in argon plus air. It is surprising, however, that even with the purest weld- ing environment the specimen was subsequently at- tacked by lithium along the grain boundaries. The introduction of impurities during welding should have been quite low. The results indicate that unalloyed columbium must not only be welded under very pure, inert atmospheres, but that the parent metal must be quite low in nitrogen and oxygen. An enlarged photograph of a crack in the weld zone of the specimen welded in argon plus air is shown in Fig, 1.2.4. The grain-boundary phase shown in Fig. 1.2.4 has not yet been identified; it was not visible in the as-welded specimens prior to the test in lithium. As was expected, the hardness values for speci- mens in the as-welded condition were guite high when the joining operation was performed in impure argon because of pickup of interstitial elements such as oxygen and nitrogen. All test specimens decreased in hardness as a result of exposure to lithium, and thus it appears that the lithium gettered the contaminants responsible for the hardness in- crease, The free energies of formation indicate that co- lumbium nitride is more stable than lithium nitride ? whereas, lithium oxide is considerably more stable than columbium oxide. Therefore, lithium metal would probably reduce any columbium oxide present as a precipitate in the grain boundaries of colum- bium welds, but it should not attack columbium nitride. Specimens selectively contaminated with nitrogen and oxygen are being tested that were prepared as described in Chap. 1.1 in order to identify the contaminant responsible for the cor- rosion of columbium welds, 2E, E, Hoffman and W. D. Manly, Lithium as a Nuclear Reactor Coolant, ORNL-2538 (August 8, 1958), p 31. ANP PROJECT PROGRESS REPORT Table 1,2,5. Results of Exposure of Columbium Weld Samples to Lithium for 100 hr at 1500° F Diamond Pyramid Specimen Hardness Joining Welding Weight Results of Metallographic (500-0z load) Technique Environment Change Examination of Weld Sample . Before After (mg/in. Test Test Braze Welding Argon + air -0.7 Braze weld: 1 mil of attack 361 343 (85% Zr=15% Cb Filler alloy Heat-affected zone: complete grain- 360 94 boundary penetration of specimen Base material*: 4 mils of subsurface 129 112 voids Argon +1.0 Braze weld: no attack 285 246 +1600-ppm N2 Heat-affected zone: 15 mils of attack 114 98 +400-ppm O2 Base material: 4 mils of subsurface 120 104 voids Argon +0.7 Braze weld: no attack 289 242 +67-ppm N2 Heat-affected zone: 6 mils of attack 122 98 +3-ppm O2 Base material: 3 mils of subsurface 113 105 voids Fusion Welding Argon + air -0.4 Weld: complete and very heavy attack 356 128 of entire weld Heat-affected zone: 15 mils of inter- 127 90 granular attack Base material: 3 mils of attack 127 97 Argon +0.6 Weld: complete attack of grain bound- 155 105 +1600- N aries; attack not as concentrated as ppm Ny in fusion weld described above +400-ppm 02 Heat-affected zone: 9 mils of inter- 116 93 granular attock Base material: 3 mils of attack 128 107 Argon +0.2 Weld: complete attack of grain bound- 131 95 4 67-ppm N aries in a few scattered areas; less PP 2 attack than in either of the fusion +3-ppm 02 welds described above Heat-affected zone: 7 mils of inter- 118 92 granular attack Base material: 1 mil of attack 127 93 *The as-received arc-cast columbium contained 360-ppm 02, 20-ppm N2, and 140-ppm C, and it had a Diamond Pyramid hardness of 137. 16 PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED Y-26784 WELD AREAS 85% 2r =15 % Cb FILLER METAL ‘\‘ 0500 -in.-0D, 0.035-in.— WALL TUBING \COLUMBIUM BRAZE WELDED IN ARGON +AIR BRAZE WELDED IN ARGON + 67ppm Ny + 3ppm O, HEAT-AFFECTED AREAS Fig. 1.2.2. Effect of Welding Environment on the Corrosion Resistance of Columbium Braze Welded with an 85% Zr—15% Cb Alloy and Exposed to Lithium at 1500°F for 100 hr. Unetched. (Swmmgt with caption) UNCLASSIFIED (@) (6) (e) Vezeris 0.03 in. Fig. 1.2.3. Effect of Welding Environment on the Corrosion Resistance of Columbium Fusion Welded and Ex- posed to Li ium at 1500°F for 100 hr. (a) Weld zone of specimen welded in argon plus air. (b) Weld zone of speci- men welded in argon plus 1600-ppm N, plus 400-ppm O,. (c) Weld zone of specimen welded in argon plus 67-ppm N, plus 3-ppm O,. Etchant: HF-HNO-H,S0,-H,0. Wwmmpt with caption) 17 ANP PROJECT PROGRESS REPORT UNCLASSIFIED | Y-26663 0.010in. Fig 1.2.4. Enlorgement of Crack Shown in Fig. 1.2.3a COMPARISON OF METHODS FOR DETERMINING THE NITROGEN CONTENT OF COLUMBIUM E. E. Hoffman The validity of the vacuum-fusion method for de- termining the nitrogen content of columbium has been evaluated by checking ¥he results against those obtained with the micro-Kjeldahl method, Several columbium samples were analyzed by the Special Analysis Laboratory of the ORNL Ana- lytical Chemistry Division by both methods. The results obtained by the two methods were in agree- ment, as may be seen in Table 1.2.6. The micro-Kjeldah| method consists of dissolving a sample of approximately 0.1 g in concentrated H,S0,, transferring the solution to the Kjeldahl upparmus, adding NaOH, and distilling the ammonia. The distilled ammonia was analyzed colorimetri- cally with the use of Nessler's reagent. In the vacuum-fusion method, a platinum bath in a graphite crucible is used for sample dissolution. The technique used was standard except that pieces of platinum weighingatotal of about 10g are dropped into the bath after the samples have been added in order to thoroughly mix the bath material. In addi- tion, the columbium metal samples are wrapped in platinum foil so that when dropped they will sink 18 Table 1.2.6. of Columbium by Vacuum-Fusion and the Micro-Kieldahl Methods Determination of the Nitrogen Content Nitrogen Content (ppm) Columbium Samplo Code BY Vacuum-Fusion By Micro-Kjeldahl Method Method Cb Wire 35 35 No. 11 Cb Wire 3500 3500 No. 12 3400 Cb Wire 34 33 No. 13 WC-1A 160 150 160 160 160 130 deep into the bath and not float on its surface. The evolved gases are passed over hot copper oxide, and the CO, which forms is taken out in a cold trap. The H is converted to water and removed in a magnesium peychlora'e absorption tube. The re- maining gas is N,, which is defermined in a McLeod gage. MOLYBDENUM AND COLUMBIUM TESTED IN LITHIUM IN SEESAW-FURNACE APPARATUS E. E. Hoffman A series of dynamic corrosion tests of molybdenum in contact with lithium were conducted with the use of a seesaw-furnace apparatus. The test assembly is shown in Fig. 1.2.5. In the tests, a 15-in.-long molybdenum pipe (0.80 in. OD x 0.10 in, wall) was partially filled with lithium, and the molybdenum pipe was placed inside an Inconel pipe (1.31 in. OD x 0.07 in. wall). The annular space between the molybdenum and the Inconel was partially filled with sodium, The Inconel pipe protected the mo- lybdenum from oxidation, while the sodium acted as the heat-transfer bond between the two containers. The test assembly was then placed in the seesaw funace, where it was rocked up and down (% cpm) with the desired temperature gradient between the ends of the capsule. Hot-zone and cold-zone -~ UNCLASSIFIED ORNL— LR-—-DWG 30678 SPECIMEN 1.0x0.60x0.065 in. INCONEL TC WELL Fig. 1.2.5. Seesow-Furnace Test Assembly for Evalu- ating the Corrosion Resistance of Molybdenum in Lithium. Tflble ]n207. PERIOD ENDING SEPTEMBER 30, 1958 specimens were suspended in the ends of the mo- lybdenum tube, so that weight changes and the attack in each zone could be determined. The extruded molybdenum pipes used for the six tests conducted thus far were made from unalloyed vacuum-arc-cast material, The specimens sus- pended in the hot and cold zones were made from arc-cast molybdenum-titanium alloy sheet, The unalloyed stock was found by analysis to contain 0.02 wt % C, 0.0012 wt % 0,, and 0,0005 wt % N,,. The alloy stock contained, in addition to the 0.4§ wt % titanivm, 0.02 wt % C, 0.0031 wt % G,, and 0.0005 wt % N,. A summary of the results obtained to date in the experiments is given in Table 1,2.7, The results show that molybdenum has very good resistance to lithium at quite high temperatures. The results of test S5-514 were not in agreement, however, with those obtained in the other five tests. The cold-zone specimen showed a weight gain and metal-crystal deposition. Spectrographic Results of Seesaw-Furnace Tests of Molybdenum Exposed to Lithium T Specimen emperature . o Duration - : Weight Change ("F) Lithium Analysis .2 ) Test No. of Test (mg/in.) Results of Metallographic Hoet Cold (hr) N2 (ppm) 02 {ppm) Hot Cold Examination of Specimens Zone Zone Zone Zone §$S-514* 1700 1300 500 284 210 —25.8 +9.0 Hot zone: surface pits to a depth of PRSCE 7L R 0.4 mil Cold zone: metal crystal deposition to a depth of 0.3 mil 55-525* 1700 1000 500 4770 1450 -0.1 +0.4 Not yet examined 55-523** 1800 1000 100 284 910 —-0.3 —0.7 Hot zone: surface roughened slightly, less than 0,2 mil of attack Cold zone: no attack or deposition $S-5246** 1800 1050 500 4770 1450 ~0.1 +2.2 Not yet examined §S-522** 1900 1000 100 284 910 —0.4 —0.8 Hot zone: surface roughened slightly, less than 0.2 mil of attack Cold zone: no attack §5-527** 1900 1100 150 4770 1450 -0,3 +0.4 Not yet examined *The tested specimens were hand polished for metallographic examination; the polishing removed ™~ 0,25 mil of surface. **The tested specimens were surface ground for metallegraphic examination; the grinding removed ™~5 mils of p 9 R Lo g p g g surface, 19 ANP PROJECT PROGRESS REPORT analysis of the specimen surface did not reveal any differences between as-tested and before-test speci- mens. Test $5-525 was a recheck of test $5-514, and the weight-change data were in good agreement with results obtained in the tests at higher temper- atures. At the present time the anomalous results of test 55-514 are not understood, and further work is planned in an effort to explain the previous re- sults. The ends of the molybdenum capsule and the hot- and cold-zone specimens from test $5-526 are shown in Fig. 1.2.6, and an enlargement of the hot zone of the molybdenum capsule is shown in Fig. 1.2.7. No molybdenum capsule failures have occurred to date. Test $5-527 was terminated by a sodium leak through the Inconel protective capsule wall, and, although the molybdenum capsule was attacked in the region of the leak, it did not fail. That the molybdenum capsule did not fail may have been due to gettering of the available oxygen by the sodium during the time required for the furnace to cool. Hot- and cold-zone specimens from test 55-522 are shown in Fig. 1.2.8. Very little, if any, attack may be detected even at very high magnifi- cation. The hot-zone specimen partially recrystal- lized during the test. The cold-zone specimen sur- face shown in Fig. 1.2.8 is quite similar to that of the metal before test. In all these seesaw-furnace tests, some dis- similar-metal mass transfer from the Inconel pipe wall to the outer surface of the molybdenum capsule walls in the hot zone was observed. Mass transferred crystals were also found on the cold- zone wall of the Inconel pipe exposed to sodium. Thermal-convection loop tests of molybdenum- lithium systems are planned in order to check the corrosion resistance of molybdenum in a system with unidirectional flow. A series of columbium-lithium seesaw-furnace tests is under way for which a test configuration similar to that shown in Fig. 1.2.5 is being used. In the columbium tests, the outer protective cap- sule is type 316 stainless steel rather than Inconel as in the molybdenum tests. The columbium tubes are 15 in. in length, 1.25 in. in outside diameter, and 0.032 in. in wall thickness. The only test per- formed to date was terminated after 300 hr by a sodium leak in the stainless steel pipe. The co- lumbium tube did not fail as a result of the failure in the protective pipe. The hot- and cold-zone specimens were at 1600 and 1100°F, respectively. 20 The hot-zone specimen showed a slight weight gain (0.1 mg/in.2), which may be attributed to gettering of residual nitrogen from the lithium, while the cold-zone specimen showed no weight change. Further tests are planned at higher tem- peratures and for longer test periods. = HOT ZONE, 1800°F WEIGHT CHANGE, —-0.4mg/in2 weLassiEs Vi COLD ZONE, 1050°F WEIGHT CHANGE, +2.2 mg/in2 pialiiin o 0.5 10 INCHES Fig. 1.2.6. Hot- and Cold-Zone Molybdenum Capsule Sections ond Specimens from Test 55-526. Lithium cir- culated inside the seesaw capsule for 500 hr under the temperature conditions indicated. Sodium flowed over the outside of the molybdenum capsule (see Fig. 1.2.5). (@ww————! with caption) wepssnes ——oesin——f Fig. 1 Enlargement of Hot-Zone End of Molyb- denum Capsule From Test $5-526. Welding operations were performed in an inert-atmosphere chamber. This section of the test capsule was at 1800°F for 500 hr. (Swmtlm—! vith caption) UNCLASSIFIED Y+26675 INCHES PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED Y-26676 “.' (@) HOT-ZONE, 1900°F 1500 st 1000°F Fig. 1.2.8. Molybdenum Specimens from Seesaw-Furnace Test §5-522. Etchant: 50% NH,OH-50% H,0,. Specimens nickel #Y8F8® to preserve edges during metallographic polishing. (UMl with caption) BERYLLIUM EXPOSED TO LITHIUM E. E. Hoffman In previous corrosion tests? a beryllium speci- men in an iron capsule exposed to lithium at 1830°F for 400 hr showed extensive attack and large weight losses. Tests have recently been conducted in iron capsules and in all beryllium systems, and the results are compared in Table 1.2,8. In the recent tests, as in the earlier test, the specimen tested in an iron capsule ot 1830°F suffered extensive solu- tion attack as a result of dissimilar metal mass transfer of beryllium from the specimen to the cap- sule wall. The beryllium in an all beryllium system showed very good corrosion resistance to static lithium at elevated temperatures. The serious effect of dissimilar metal mass trans- fer on the corrosion resistance of the beryllium is 3A. deS. Brasunas, Interim Report on Static Liquid- Metal Corrosion, ORNL-1647 (May 11, 1954). o illustrated in Fig. 1.2.9. The rectangular shape of the beryllium specimen was altered appreciably in the test conducted in an iron capsule. The slight weight gains of the beryllium specimens fested in beryllium containers are probably due to gettering of impurities such as oxygen and nitrogen from the lithium. The surfaces of the beryllium specimens tested recently at 1830°F are shown in Fig. 1.2.10. Considerably more porosity and surface roughening is apparent on the specimen tested in an iron con- tainer than on the specimen tested in a beryllium container. Some indication of the extent of dissimilar metal mass transfer is illustrated in Fig. 1.2.11. The Be,Fe intermetallic compound on the surface was identified by x-ray analysis. This phase has a hexagonal structure, and its anisotropy is apparent in the photomicrograph taken with polarized light (Fig. 1.2.118). 21 ANP PROJECT PROGRESS REPORT Table 1.2.8, Corrosion of Beryllium Specimens by Lithium in Static Tests Attack (mils) Temperature Time o Specimen Weight Change - (°F) () Commotun (mg/em?) Intergranular Solution 1830% 400 ron ~101.8 127 30 P 1500 100 Iron -3.9 4 0 1500 100 Beryllium +0.3 2 0 1830 100 Iron —66.1 5 2 1830 100 Beryllium +03 3 [) *Previous test, ref 3. UNCLASSIFIED Y-26285 [¢] 0.5 10 INCHES BEFORE TEST TESTED IN Fe CAPSULE TESTED IN Be CAPSULE WEIGHT CHANGE, WEIGHT CHANGE, p —66.4 mg /cm? +0.3 mg/cm? Fig. 1.2.9. Effect of the,Expasure.of Beryllium to Static Lithium for 100 hr ar 1830°F. (q@g@mwith caption) UNCLASSIFIED UNCLASSIFIED Y-26413 Y-26409 Fig. 1.2.10. Surfaces of Beryllium Specimens Shown in Fig. 1.2.9 Which Had Been Exposed to Static Lithium for e 100 hr ot 1830°F. (a) Tested in an iron capsule. (b) Tested in a beryllium capsule. As-polished. 500X. Wlwmme with caption) 22 PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED Y-26416 Fig. 1.2.11. Wall of Armco Iron Capsule in Which a Beryllium Specimen Was Tested in Contact with Static Lithium for 100 hr at 1830°F. Note dissimilar metal mass transfer of beryllium to capsule wall. As-polished. (a) Bright*fi@ldeM|umination. (b) Polarized light illumination. REFRACTORY-METAL-BASE BRAZING ALLOYS EXPOSED TO LITHIUM D. H. Jansen Most commercially available brazing alloys contain constituents used to alter the melting point, lower the flow point, increase flowability, or im- prove the ductility of the alloy. These constituents include the precious metals, copper, and manganese, to mention a few. The additives are beneficial for their specific purposes, but they may produce alloys that possess limited corrosion resistance to the liquid metals, especially lithium. A typical ex- ample, a 60% Mn—40% Ni brazing alloy which failed completely when exposed to lithium for 100 hr at 1500°F, is shown in Fig. 1.2.12. Analyses of microdrillings from the remaining alloy fillet indi- cate that the manganese was attacked to the same extent as the nickel. Another example, an 86% Fe-5% Si—5% Cu-4% B brazing alloy on a type 347 stainless steel tube-to-header joint which was tested in lithium in a seesaw-furnace apparatus for 100 hr at 1500°F, is shown in Fig. 1.2.13. The copper concentration of the lithium bath after this test was very high. As a result of the poor corrosion resistance of the alloys described above, refractory-metal-base ®wmy with caption) brazing alloys are being investigated for use in high-temperature lithium systems, Columbium tubing (‘/2 in. OD) has been used for capsules in tests in which small (5 g) brazing alloy samples (buttons) have been corrosion tested in static lithium at 1700°F, Conventional methods for loading and testing the brazing alloys were used. The corrosion testing program on these refractory- metal-base alloys has involved, initially, static tests of zirconium- and titanium-base binary and ternary alloy buttons. The alloys which appear promising with respect to corrosion resistance will be used as a basis for developing other alloys with lower flow points or better flow characteristics. (Melting-point and flow-point determinations and flow-characteristic studies are being made by the Welding and Brazing Group, see Chap. 1.3.) The more promising alloys will then be used for brazing molybdenum or columbium T-joints which will be given more severe corrosion tests (seesaw-furnace tests or tests in thermal-convection loops with brazed sections in the hot legs). The results for all the refractory-metal-base alloys corrosion tested in lithium to date are presented in Table 1.2.9. The alloy button con- taining vanadium (70% Zr-30% V) exhibited the 23 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 094 Y-7776 005 BASE MATERIAL BASE MATERIAL Lo Fig. 1.2.12. A 60% Mn—40% Ni Brazing Alloy on a Type 316 Stainless Steel T-Joint After Exposure to Static Lithium for 100 hr at 1500°F. Complete penetrationgafethie brazing alloy is evident. Unetched. 150X. (iiiimes! with caption) EEE] INCHES. i i fuone foor| - Lo fooef @ foef o | § Lo [t Fig. 1.2.13. (o) An 86% Fe-5% Cu—5% Si—4% B Brazing Alloy on Type 347 Stainless Steel Tube-to-Header Joint After 100 hr of Exposure to Lithium in a Seesaw-Furnace Apparatus at a Hot-Zone Temperature of 1500°F. (b) En- largement of Area Enclosed by Dotted Lines in (o). Unetched. 50X, 250X. Reduced 28%. Mmmiissmmb with caption) 24 PERIOD ENDING SEPTEMBER 30, 1958 Toble 1.2.9. Results of Static Corrosion Tests on Refractory-Metal-Base Arc-Cast Brazing Alloy Buttons Exposed to Lithium in Columbium Capsules at 1500 or 1700°F for 500 hr Braze Material Test Composition Temperature Weight Change* Metallegraphic Results and General (wt %) (°F) (%) Condition of Button After Test 95 Zr-—-5 Be 1700 +0.005 Subsurface void formation to a depth of less than 84 Zr-16 Fe 1500 -0.573 No attack; numerous small cracks in specimen -0.62 82 Zr.18 Cr 1700 -1.81 No attack observed 75 Zr-25 Cb 1700 -0.227 No attack cbserved -~0.11 75 Zr—15 Cr=10 Fe 1500 —-2.5 Edge of specimen attacked nonuniformly toa depth -1.7 of 1 mil 70 Zr-30 V 1700 —-10.87 Second phase leached to a depth of 9 mils -2004 69 Zr-31 Mo 1700 -1.34 Neo attack observed - 1030 80 Co-20 Cb 1700 -2.6 Specimen uniformly depleted to a depth of 0.5 mil; spotty attack found that varied from 0 to 10 mils 55 Ti—45 Zr 1700 -0.063 No attack observed +0.079 72 Ti~28 Co on Cb Spotty attack to less than 1 mil in depth T-jcint PP *Each weight change is the result of one test; the buttons weighed approximately 5 g. most weight change during the test. The micro- structure of the edge of the alloy before and after the test is shown in Fig. 1.2.14, Both vanadium and zirconium had previously exhibited good cor- rosion resistance to lithium,**> and a microspark traverse on the as-tested 70% Zr-30% V alloy button showed essentially no alteration of compo- sition from the edge into the interior of the alloy. It appears therefore that the zirconium and va- nadium may both have gone to the columbium cap- sule by a process of dissimilar metal mass transfer, This particular test is being repeated to determine whether the large weight loss is due to mass trans- fer. 4D. H. Jansen and E. E, Hoffman, ANP Quar. Prog. Rep. Sept. 30, 1957, ORNL-2387, p 215. SNuclear Pyppulsion Pyogram, Engineering Progress Report, April 1, 1958—]June 30, 1958, PWAC.582, The only alloys that showed severe corrosion at- tack were the 75% Zr-15% Cr—10% Fe and the 80% Co-20% Cb (Fig. 1.2.15) alloys. Attack on these alloys was quite spotty and nonuniform, The in- homogeneity of the alloy buttons, which may be the cause of nonuniform attack, is being studied. Previous static tests on pure cobalt in lithium under slightly more severe conditions (1832°F for 447 hr) showed intergranular attack and a weight loss of 7% (0.7 g/in.2). Consequently, the attack seen on the 80% Co-20% Cb alloy (Fig. 1.2.15) is understandable, in view of the high cobalt content of the button. Titanium-base brazing alloys in button form and in the form of braze fillets on columbium T-joints were tested and the results are being evaluated at present, One titanium-base alloy (72% Ti-28% Co) on a columbium T-joint tested in 1700°F lithium for 100 hr showed approximately 1 mil of nonuniform attack on the exposed surface (Fig. 1.2.16), ANP PROJECT PROGRESS REPORT Fig. 1.2.14. Alloy Button of 70% Zr—30% V, () Before and (b) After a Static Corrosion Test in Lithium at 1700°F for 500 hr. Note depletion of the second phase fo o depth of 10 mils. 250X. (@mmissssswmb with caption) 2 Ni PLATE ; UNCLASSIFIED EEE T INC‘!ES T f L EEEEREER Fig. 1.2.15. An 80% Co-20% Cb Alloy Button After Exposure to Static Lithium for 500 hr at 1700°F. The attack shown at the exposed edge was nonuniform and varied from 0 to 10 mils. Unetched. 250X. (Gamimimmbind with caption) 26 PERIOD ENDING SEPTEMBER 30, 1958 Fig. 1.2.16. A 72% Ti-28% Co Brazing Alloy () Before and (b) After Testing in Static Lithium at 1700°F for 100 hr. Dark areas at the braze and base material interface are microcracks. Corrosion of the exposed edge is li (dSmmiinimtionh vith caption) No attack of the type observed on the 80% Co-20% Cb alloy button was seen on this 28% Co alloy on a T-joint. Attack may be seen at only a few, scat- tered places along the surface of the fillet, The results of this test, as compared with those for the alloy button containing 80% Co, indicate that cor- rosion resistance is decreased when large per- centages (80%) of cobalt are used. DETERMINATION OF THE SOLUTION RATE OF METALS IN LITHIUM® E. E. Hoffman A method was developed at NDA for measuring the rate at which container metals dissolve in liquid lithium under controlled nonequilibrium con- ditions.” The method consists essentially of im- mersing a thin test specimen into a comparatively large volume of liquid metal held at the desired temperature in an inert container. With suitable choices of surface area and immersion time, the SSubcontract with Nuclear Development Corporation of America, July 1, 1957 to July 31, 1958, 7B. Minushkin, Determination of the Solution Rate of Metals in Lithium, NDA-44 (June 30, 1958). ited fo intermittent attack to a minimum depth of 1 mil. Etchant: 46 H,0, 46 HNO,, 8 HF. 100X. Reduced 21%. test specimen can be made to lose about 1 mg of weight in a test period of less than one day. This weight loss does not significantly increase the concentration of test specimen material in the liquid metal, but the loss can be determined with ample accuracy and precision on a semimicro balance. The solution rates determined in this manner are proportional to the specific solution rate constant, . Based on experimental evidence from thermal- convection loop tests, it is believed that inhibitors and impurities exert their effect on mass transfer rates by altering the specific solutionrate constant. Therefore, this method is suitable for preliminary studies of the effects of additives and impurities on mass transfer. The data from fests on stainless steel indicate that variables can be detected which affect the initial solution rate by as much as +20%. The re- sults indicate that nickel-bearing stainless steel dissolves in lithium at 1600°F at an initially high rate of 3.0 mg/in.Zshr because of the preferential leaching of nickel. Within a relatively few hours, depletion of nickel in the specimen surfaces re- sults in a rapid decrease in the solution rate to a 27 ANP PROJECT PROGRESS REPORT value of about 0.3 mg/in.2+hr at solute concen- trations of a few hundred parts per million. It was shown that this solution rate accounts for the mass transfer rates observed in thermal-convection loop tests. The solution rate tests confirm the dele- terious effects of nitrogen and oxygen and the beneficial effects of aluminum additions on mass transfer rates that were observed in thermal-con- 28 vection loop tests. In addition, it appears that Misch metal and tantalum additions may have beneficial effects. Solution rate studies of molybdenum in lithium at 1600°F indicate that this material is not signifi- cantly attacked by lithium. Tests on niobium yielded conflicting results which have not been resolved. W o & PERIOD ENDING SEPTEMBER 30, 1958 1.3. WELDING AND BRAZING P. Patriarca DEVELOPMENT OF BRAZING ALLOYS FOR LITHIUM SERVICE R. G. Gilliland G. M. Slaughter Refractory-metal-base brazing alloys are being developed for application in high-temperature re- actors that use lithium as a coolant. As a starting point, a literature survey was made of the binary systems of columbium, zirconium, titanium, tantalum, and molybdenum in order to find eutectics or minimums in the approximate temperature range of interest. The data obtained for columbium and zirconium were presented previously, ! and data for titanium, molybdenum, and tantalum are summarized in Tables 1.3.1, 1.3.2, and 1.3.3. Small buttons of approximately fifty binary or ternary alloys of in- terest have been prepared for testing by arc melt- ing, and results of tests of a number of these alloys in lithium are presented in Chap. 1.2, Corrosion. A vacuum-tube fumace that has been modified for ultrashigh-temperature work is being used for preparing brazed joint specimens for testing. The furnace and the quartz muffle for use at temper- atures up to 1200°C are shown in Fig. 1.3.1. A ceramic muffle made primarily of zirconium oxide is available for use at 1350°C. Vacuums of less than 1 u are consistently obtained. A number of the arc-melted alloys have been tested for flowability at various temperatures by using the T-joint sample design with both columbium and molybdenum as base metals., Flow points were obtained from these tests, and metal- lographic examinations of the resulting joints were conducted. The flow points and general brazing characteristics of the several alloys tested thus far are presented in Table 1.3.4. The brittleness of the brazing alloy in the as- brazed condition is an important problem in this study, as may be seen in Fig. 1.3.2, which shows a brazed molybdenum T-joint that contains numerous fillet cracks. The brazing alloy used was 84 wt % Zr and 16 wt % Fe. In comparison, the molybdenum T-joint brazed with an alloy containing 70 wt % Zr and 30 wt % V, shown in Fig. 1.3.3, is considered to be satisfactory, 16, M. Slaughter, ANP Quar, Prog. Rep. March 31, 1958, ORNL-2517, p 35. A study of the diffusion of brazing alloys on re- fractory metals has been initiated. Lap joints made from columbium, and possibly molybdenum, brazed with the different altoys will be used in the study, and particular attention will be paid to brazing alloys containing refractory metals, The joints will be aged just befow, at, and just above Table 1,3.1. Titanium Binary Systems of Potential Interest As Brazing Alloys Eutectics or Minimums* System Composition Melting Point (wt %) (°F) Ti 3035 Ti-Ag 96 Ag, eutectic 1650 Ti-Au 15 Au, eutectic 1530 82 Au, eutectic 2345 Ti-Be 3 Be, eutectic 1750 90 Be, eutectic 2300 Ti«Cr 47 Cr, minimum 2550 Ti-Co 28 Co, eutectic 1870 81 Co, eutectic 2075 Ti«Cu 50 Cu, eutectic 1740 70 Cu, eutectic 1600 Ti-Ge 19 Ge, eutectic 2570 Ti-Fe 31 Fe, eutectic 1985 85 Fe, eutectic 2415 Ti=Pb 26 Pb, eutectic 2250 TieMn 42 Mn, eutectic 2150 Ti=Ni 28 Ni, eutectic 1750 65 Ni, eutectic 2030 83.8 Ni, eutectic 2350 Ti=5i 9 Si, eutectic 2430 51 Si, eutectic 2700 78 Si, eutectic 2430 Ti-Y 30 V, minimum 2950 Ti=Zr 50 Zr, minimum 2900 Ti-Th 88 Th, eutectic 2200 *A. D. McQuillan and M. K. McQuillan, Titanium, New York Academic Press, New York, 1956. ANP PROJECT PROGRESS REPORT Table 1.3,2. Molybdenum Binary Systems of Potential Table 13,3, Tantalum Binary Systems of Potential Interest As Brazing Alloys Interest As Brazing Alloys Eutectics or Minimums* Eutectics or Minimums* System Composition Melting Point System Composition Melting Point (wt %) °F (wt %) (°F) Mo 4800 Ta 5425 Mo-Ni 47 Ni, eutectic 2400 To-Co 67.6 Co, eutectic 2325 Mo-Co 62 Co, eutectic 2444 Ta-Si 94 Si, eutectic 2525 Mo-Si 12 Si, eutectic 2600 To-Fe 80 Fe, eutectic 2570 Mo-Al 22 Al, eutectic 4200 TaeNi 62 Ni, eutectic 2480 Mo-B 1B, eutectic 3540 39 Ni, eutectic 2552 Ve D0y Anctlc 2624 *M. Hansen, Constitution of Binary Alloys, McGraws Mo-Mn 99 Mn, eutectic 2191 Hill, New York, 1958. Mo-Cb 53 Cb, eutectic 2400 *M. Hansen, Constitution of Binary Alloys, McGraw- Hill, New York, 1958. UNCLASSIFIED PHOTO 44634 Fig. 1.3.1. Vacuum Tube Furnace and Quartz Muffle for Refractory-Metal Brazing Alloy Studies. The furnace is mounted on a track fo permit rapid heating and cooling of the test sample. 30 PERIOD ENDING SEPTEMBER 30, 1958 Table 1,3.4. Data Obtained in Refractory-Metal«Base Brazing Alloy Study Metallographic Observations . Flowability Brazing Alloy Flow of Joints Composition Point Brazed on Brazed on (wt %) ) Ch Mo Brazed on Brazed on Cb Mo 84 Zr—16 Fe 9347 Good? Good Severe cracks Severe cracks 65 Zr~25 V-10 Fe >1300 75 Zr~15 Cr-10 Fe >1300 Good Poor® Severe cracks 82 Zr18 Cr 13007 Good Goad Severe cracks No cracks 65 Zr~25 V-10 Ge ~ 1300 Fair? Good Severe cracks No cracks 67 Zr—29 V4 Cr >1300 63 Zr=-27 V-10 Cr >1200 60 Zr—-26 V14 Cr >1200 57 Zr~24 V-19 Cy >1200 70 Zr~-30 V 1230% .Good Good No cracks No cracks 67 Zr—29 V-4 Fe 1300 Good Good 65 Zr--28 V-7 Fe >1200 60 Zr-26 V-14 Fe > 1200 L S B 4 . 63 Zr-27 V-10 Fe > 1300 50 Zr—-21 V=29 Mo >1300 57 Zr—-24 V-19 Mo >1200 63 Zr~27 VY-10 Mo >1200 60 Zr-25 V-15 Cb 1300 Goed Good No cracks No cracks 69 Ti-31 Fe 10857 Good Good Slight cracks Severe cracks 72 Ti~28 Co 10254 Goad Gaad Na cracks No cracks M. Hansen, Constitution of Binary Alloys, McGraw-Hill, New York, 1958. bGood indicates continuous filleting and extensive spreading en joint. “Poor indicates wetting only at contact points. dFair indicates intermittent filleting and little spreading on joint. the brazing temperature. After 8- and 100-hr aging WELDING STUDIES OF COLUMBIUM periods the specimens will be examined to determine AND MOLYBDENUM the extent of diffusion of the brazing alloy elements R. L. Heestand into the base metal or alloying of the base metal with the brazing alloy. If such diffusion or alloy- A small, glass dry box containing a stationary ing occurs, it is hoped that it will increase the re- torch and a movable table was assembled, as shown melt temperature of the brazing alloy and thus in- in Fig. 1.3.4, with vacuum equipment capable of crease the maximum service temperature attainable. maintaining vacuums as low as 4 x 105 mm Hg, for LY I -4 31 ANP PROJECT PROGRESS REPORT UNCLASSIFIED Fig. 1.3.2. As-Brazed Molybdenum T-Joint Brazed with an Alloy Containing 84 wt % Zr and 16 wt % Fe. The joint contains numerous fillet cracks. As-polished. 150X. (Gumdigssm®l with caption) . Molybdenum T-Joint Brazed with an Alloy Containing 70 wt % Zr and 30 wt % V. ®Swmtigiauin| with caption) 32 UNCLASSIFIED Y-26820 As-polished. kR INCHE S ETTTTRITTTRITITEIL: o 1s0x 50X. PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED PHOTO 43496 Fig. 1.3.4. Welding Dry Box with Stationary Torch and Movable Table. welding studies of columbium and molybdenum. The equipment was designed to make both fusion welds (no filler metal addition) and welds with filler metal, and it can be used for the evaluation of the weldability of refractory metals and their alloys in controlled atmospheres. The setup for preparing edge-fusion and butt joints in the con- trolled atmosphere chamber is shown in Fig. 1. Several typical butt welds made on arc-cast columbium sheet in the automatic welding dry box in a purified argon atmosphere are shown in Fig. 1.3.6. Free-bend tests at room temperature showed samples of all these welds to be ductile. Metal- lographic studies of the samples are under way. Five different heats of arc-cast molybdenum sheet for use in welding studies were analyzed for carbon, nitrogen, and oxygen. Samples of the 0.040-in.- thick sheet were then welded in the dry box de- scribed above and machined into bend specimens. The bend specimens are being tested to determine the effect of the contaminants on the ductile- brittle transition temperature of material from each heat. All specimens tested thus far have been found to be brittle at room temperature. 33 ANP PROJECT PROGRESS REPORT UNCLASSIFIED PHOTO 43497 Fig. 1.3.5. Close-Up of Dry Box Equipment Showing Setup for Preparing Edge-Fusion and Butt Joints. UNCLASSIFIED Y-26115 INCHES Fig. 1.3.6. Butt Welds in Columbium Sheet Made with (a) the Addition of 85 wt % Zr-15 wt % Cb Filler Wire, (b) the Addition of 82 wt % Zr—15 wt % Cb=3 wt % Mo Filler Wire, (c) the Addition of 100 wt % Cb Filler Wire, and (d) Without Filler Wire. (@uWniiggmsim with caption) The selection of structural materials capable of withstanding service loads at the high temperatures at which reactors for the propulsion of aircraft will operate has necessitated determinations of the basic strength properties of new materials and the development of accurate and precise analytical pro- cedures for the evaluation of materials. Equipment is now available which permits the rapid testing of materials in controlled environments at temperatures up to 2500°F. Means for obtaining quantitative de- sign information on the behavior of metals under complex states of stress and strain have been studied, and it has been found that experimental creep data are in substantial agreement with the results of calculations based on accepted theories of creep and fracture. DEVELOPMENT OF TEST EQUIPMENT FOR HIGH-TEMPERATURE INVESTIGATIONS D. A. Douglas In the development of equipment for studying the mechanical properties of both metals and nonmetals in the temperature range from 1800 to 3000°F, it was necessary to provide control of the test en- vironment, which would be gas or a vacuum, to provide control of the temperature gradient and the temperature fluctuation comparable to that provided UNCLASSIFIED. ORNL=LR-OWG 33728 U-CUP SEALS FOR PULL ROD. ~PULL ROD WATER-COOLED FLANGE |— CHAMBER (10in. Dia X 12in. LONG) FURNACE (THREE SPECIMEN WITH PLATINUM ELEMENTS P1-Rh THERMO- SEPARATELY CONTROLLED) COUPLES —— 1 __TO VARIAC AND ALuNDUM | | [CONTROLLER 110 v EMFFLE === | O VARIAC AND i [[CONTROLLER 1 v U E===—rocroumo TO VARIAC AND CONTROLLER 110 v Fig. 1.4.1. Diagram of an Apparatus for Creep and Tensile Tests in Controlled Environments at 2000°F and Above. PERIOD ENDING SEPTEMBER 30, 1958 1.4, MECHANICAL PROPERTIES D. A. Douglas for conventional creep and tensile tests, to assure accurate strain and load measurements, and to keep the design simple and compact so that specimens could be tested quickly. A schematic diagram of the apparatus designed within this framework is shown in Fig. 1.4.1, and a photograph of the parts ready for assembly is shown in Fig. 1.4.2. The use of U-cup O-rings as sliding seals simplifies the application of the load and accommodates the usual displacement in the tensile and creep tests. y Lo h Vason ) Fig- 1.42. Specimen, Pull Rods, and Grips Used in Apparatus Shown in Fig. Ld.1. 35 ANP PROJECT PROGRESS REPORT The furnace element currently being used in the apparatus is platinum, but molybdenum or tungsten could be used for operation at higher temperatures. The three elements of the furnace are separately controlled by Variacs, and temperature deviations along the gage length can be held to +5°F, High-purity argon has been used as the test en- vironment for all of the tests conducted to date. The apparatus is amenable to evacuation, but the increased time and trouble involved in vacuum tests make high-purity inert gases more desirable as test environments, The reliability of the ap- paratus has been checked in three tests on mo- lybdenum specimens at 2000°F, and the results indicate that a suitable device has been achieved. MULTIAXIAL CREEP STUDIES C. R. Kennedy Analytical studies have been made of creep data obtained for Inconel tubular specimens in tests in which internal pressure, as well as axial loading, was applied in the apparatus described previously.’ The internal pressure required to produce a given tangential stress was calculated by using the thin- wall formula. The axial stress was calculated conventionally, with the axial stress produced by the internal pressure being taken into consideration. The average radial stress was obtained by dividing the internal pressure by 2, After-test measurements of the tube dimensions were made to determine the three principal strains at rupture. In those cases where end effects or bulging occurred, tangential strains measured at regions of highly localized de- formation could not be compared directly with the over-all axial rupture strain in connection with subsequent data analysis. |t was necessary, there- fore, to use average tangential strains for this purpose. These were obtained by averaging five measurements of the tangential strain taken at equidistant positions along the 21/2-in.-goge length, An observation that the sum of the three measured average strains often was greater than zero sug- gested an increase in volume during the test, Since a volume change is inconsistent with the assump- tions used in the mathematical analysis of creep flow and fracture, this matter was investigated by c. R Kennedy and D, A, Douglos, ANP Quar. Prog. Rep. Sept. 30, 1957, ORNL-2387, p 185. 36 measuring the specific gravities of after-test speci- mens by using a Jolly balance. In these tests, ob- served density changes were found to be less than the sensitivity of the balance, and thus the apparent volume change based on after-test strain measure- ments was not real. The error was believed to be associated primarily with the radial strain measure- ments and was attributable to the inaccuracies in- herent in measuring small changes in the sizes of small sections with curved surfaces. In the data analysis which follows, the ‘‘measured’’ radial strains at fracture were determined from axial and tangential strains based on constancy of volume, Creep Analysis A workable formulation for steady-state creep is that of Soderberg,? which is based upon the follow- ing assumptions: (1) the material is originally isotropic and remains isotropic during the creep process so that the stress effect on creep rates canbe expressed in terms of the principal stresses; (2) the principal stresses coincide with the principal strain directions; (3) the volume remains constant so that hydrostatic pressure has no influence and the stress effect must therefore be a function only of the difference of the principal stresses (that is, shear stresses); (4) the stress state remains constant with time; and (5) the effective creep stress, 7, and the effective creep strain rate, €, for all stress states at a given temperature are related through the material constants K and » by the relation (1) E=KG" . According to this formulation, the principal steady- state creep rates are 2¢. R, Soderberg, Trans, Am, Soc. Mech, Engrs, 63, 737 (1941), where the subscripts Z, 6, and R denote axial, tan- gential, and radial stress or strain rates, respec- tively. Based on the assumptions stated, only the principal stresses and the relation between creep stress and creep strain for simple tension are re- quired in order to compute multiaxial creep rates according to Eq. 2, The value of & depends on the flow criterion selected to correlate the data. Based on the von Mises (distortion energy) criterion, it is given by 1 3) © =———{(az-—08)2+ (ag-oR)2+ V2 and, by the Tresca (maximum shear stress) criterion, it is given by (4) g zamax —Omln . The effective creep strain rate, €, is the creep strain rate associated with & for a value of uniaxial stress equal to the calculated effective stress. The design criterion based on creep rate in many applications is the time to reach a limiting strain. For the purpose of comparing experimental results with those predicted by Eq. 2, the average creep rate to 2% strain was used. In those cases where rupture occurred before 2% axial strain had oc- curred, the average creep rate to one-half of the total strain was used, Comparisons of the measured axial creep rates with those predicted from Eq. 2 by using both the von Mises and the Tresca criteria are shown in Figs. 1.4.3, 1.4.4, and 1.4.5. In general, the corre- lation of the experimental data obtained by using the von Mises criterion is slightly better than that obtained by using the Tresca criterion. In most design situations, however, the stresses are known only approximately, and, in view of the many inde- terminate variables which influence creep, the simpler and more conservative Tresca criterion is clearly adequate for engineering purposes, The advantages of using the Tresca criterion for multi- axial creep have been demonstrated by Wah!? for the case of rotating disks. According to Wahl, the use of the von Mises criterion predicted much lower creep rates than those obtained experimentally. 3A. M. Wahl, J. Appl. Mechanics 23, 231 (1956). PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED ORNL-LR-DWG 31530 -1 10 N S ,717211’ - = ’4177 pa— :‘ - e 2 0, 6000 psi — - | ~ B L = 4 —F r E e o //* i;[ o7y =5000psi ~ ] o e . S B g, e A ! é TO .. — /Il 4 o S S—— i 4000 psi ] fl — = T | 5 [ A = ,’ 3 . 2 s, =3000psi wn o, = psi ——— T Pt 4 e e e LI = "~ " _I ~— 1V g > < ——-I7 = TRESCA CRITERION ~——[¥ = von MISES CRITERICN . 104 } @ 4 2 1.33 1 0.8 0.67 0.57 STRESS RATIO (a7 /) Fig. 1.4.3. Axial Tension Creep Rates vs Stress Ratios for Four Principal Axial Stresses. UNCLASSIFIED ORNL-LR-DWG 34534R 10 % 1 5 2 rad L 0= 6000 psi 77 4 4 ‘l I, > 7 5 7 7 AR // v Wa 2 ,/ - [5=4000psi ] 7/ - ® 107! LA s 77 > 7 r i d P 5 r/ /" rd —~ / 7 2V 1 7 9,=3000 psi 7 r P > /// // Py ,1/ / L] 7/ c/ ,/ 1g72 4 A 'I‘II, ,ll I,I 5 17— 7 p 7 Pl / ) 7 ) * 1/ / v 7 1672 ’ 5 7 ' — 4 ~=—— T =TRESCA CRITERION _| 4 V'=VON MISES CRITERION 2 H T G0.25 o] -0.25 -0.5 -075 -4 -14.33 STRESS RATIC (O}/Oé) Fig. 1.4.4. Axial Compression Creep Rates vs Stress Ratios for Three Tangential Stresses. 37 ANP PROJECT PROGRESS REPORT UNCLASSIFIED ORNL-LR-DWG 34532 10 [ I T i ! 1 I ] ] 1 T 1 5L == 7=TRESCA CRITERION | em—— =VON MISES CRITERION T2 ~ B2 o 5g=2000psi '_ <1 — (0 /I [a w a4 0.5 g 7 / 2 /S o / / ) Q 0.2 7 74 o / /@ = / o oA / © ) 77— - A 4 < YA 4 % 0.05 // 0.02 0.01 4 -2 -3 -4 -5 STRESS RATIO (g /) Fig. 1.4.5. Axial Compression Creep Rates vs Stress Ratios for Stress Ratios of Less than —1 with a Constant 2000-psi Tangential Stress. Rupture Analysis Rupture data are presented in Fig. 1.4.6 in which the stress ratios calculated by the thin-wall formula are plotted against the rupture life. The rupture data are also presented in the form of 100-, 500-, and 2000-hr isochronous fracture envelopes in Fig. 1.4.7. The envelopes were obtained from cross plots of the data shown in Fig. 1.4.6. The 100-hr fracture envelope predicted by the maximum princi- pal tensile stress criterion is also plotted in Fig. 1.4.7. It may be seen that the results calculated and plotted in the manner described show deviations from the simple maximum principal stress criterion. There are, however, at least two important factors which should be considered in a realistic analysis of the data, Both of these factors involve in some 38 UNCLASSIFIED ORNL LR- DWG 31533 © . TTT T A 6000 - psi MAXIMUM 2.00 |- PRINCIPAL STRESS — “_#_ \__H b@ & 100 — (@] 5 050 - %—1 a o (5] 2 ]_T \ @ Q — - - + = = [ m . . H ' H H L - f£3000- psi MAXIMUM | 1/ | [T PRINCIPAL STRESS —-0.50 — —A 1 B J H o W il —1.00 L L] H\ 1N [ J_J___ 20 1000 10,000 TIME TO RUPTURE (hr} Fig. 1.4.6. Rupture Times vs Stress Ratios for Three Average Maximum Principal Stresses, UNCLASSIFIED ORNL-LR-DWG 31611 8000 .-\c\ i —— | 6000 [— I—— | GI\ ."'"‘--.‘ [ ~— | 4000 T3 '--...______-.‘.-‘ ‘ .\_T > P 1 ! | 2000 } | ' .’(7‘) ™Y ® | e ’ » I o s ’ a J | o M | = , 1 ! w g I ; | % 2000hr § I H | < -2000 .' 7 / . P / | [ ] ® -4000 100 hr 1 - 6000 100-hr FRACTURE ENVELOPE PREDICTED BY MAXIMUM /} PRINCIPAL STRESS CRITERION | -8000 | | 0 2000 4000 6000 8000 10,000 TANGENTIAL STRESS {psi} Fig. 1.4.7. Rupture Data Presented as 100., 500-, and 2000-ht Isochronous Fracture Envelopes. manner the changes in stress state associated with deformation of axially loaded pressurized tubes. The first factor is concerned with the stress-state changes in uniformly deformed tubular specimens. For an equivalent amount of strain under these con- ditions, the increase in the axial stress is less than the increase in the tangential stress over the entire range of stress ratios studied. For example, after 10% axial strain in a specimen tested in simple tension the '‘true’’ axial stress becomes 1.1 times as great as the original axial stress; however, after 10% outer-surface tangential strain of a speci- men subjected to tangential stress only, the aver- age tangential stress becomes 1.18 times as great as the original tangential stress, For this example, the increase in the tangential stress is greater than that of the axial stress by a factor of 1.07. This is to be compared with the data plotted in Fig. 1.4.7, where it is shown that the axial stress required to produce rupture in 100, 500, or 2000 hr is greater than the corresponding tangential stress by a factor of about 1.25. This discrepancy can be rationalized, however, by noting that the effect being considered is intensified by the second important factor which is associated with localized deformation or bulging, particularly in tension-compression states, The relative amounts of bulging are demonstrated in Fig., 1.4.8, which is a plot of the difference in the maximum and the average tangential strain at rupture versus the stress ratios. It is shown that the bulging between stress ratios of « to ‘/2 is insignificant; however, the extent of bulging in- creases substantially as the stress ratio drops from ‘/2 to ~1. The result of this bulging, if only the increase in diameter of the specimen were con- sidered, would be to increase the tangential stress, which is the maximum principal stress in this range. The relation between the degree of bulging and the deviation of results from the maximum principal stress criterion may be seen by comparing Figs. 1.4.7 and 1.4.8. For example, in Fig. 1.4.8 the degree of bulging obtained by testing at 3000 psi is shown to be less than that obtained by testing at 4000 and at 6000 psi. This is consistent with the rupture results presented in Fig. 1.4.7, which show that the deviation of the 2000-hr frac- ture envelope from a vertical line representing a maximum principal stress criterion is much less than that for the 500- or 100-hr curves (vertical reference lines not shown). Accordingly, the re- sults of this study appear to support the maximum PERIOD ENDING SEPTEMBER 30, 1958 principal stress criterion for time-dependent frac- ture under multiaxial stress conditions. It has been shown, however, that in applying this cri- terion it is important to compensate for the re- duction in rupture life caused by changes in the stress state associated with deformation, The role of the maximum principal stress is dem- onstrated by crack patterns (Fig. 1.4.9) on the out- side surfaces of specimens. A close inspection of these patterns reveals that the cracks propagate in a stair-step manner, and, in all cases, the general direction of the intergranular cracks is normal to the maximum principal stress, |t was observed that all cracks were intergranular and that none of the specimens exhibited necking prior to failure, The effect of multiaxial stress states on the rupture elongation was also investigated. Fracture strains based on after-test measurements of speci- mens tested with a 4000-psi maximum principal stress are shown in Fig. 1.4.10, in which the aver- age axial, tangential, and radial strains are plotted versus the stress ratio. The strain at failure, €/, for tests with the same maximum principal stress appears to be related to the deviatoric stress in the following way: f ' 06+0R ALY f OZ+09 (o5 A comparison of these relationships with the data in Fig. 1,4.10 indicates good agreement, The value of B is constant for all the stress states tested with the same maximum principal stress, but it varies with the maximum principal stress, as shown in Fig. 1.4.11. The rupture strains from all tests are compared with Eq. 5 in Fig. 1.4.12, which is a plot of the ratio of rupture strains to the axial strain for simple tension vs the stress ratio, Again, there appears to be good agreement with Eq. 5. The ratio of the shear strain at rupture, as deter- mined from yf: Ef --Ef max min 7 to the shear strain at rupture for specimens tested 39 ANP PROJECT PROGRESS REPORT UNCLASSIFIED ORNL-LR-DWG 31535 y2 /20 ¥ / 6000 - psi MAXIMUM PRINCIPAL STRESS o // ] // ! > 4000 - psi MAXIMUM/ PRINCIPAL / 52 z a o |_ 0 J g = z w > < 6 = STRESS w T} & w 5 - > a o, / = ®2000-psi MAXIMUM 5 s PRINCIPAL STRESS | = - L z ( Sl w L (0] / L a 0 / -1 -2 J -3 © q 2 £33 { 0.75 0.5 0.25 0 -0.25 -0.5 -075 -{ STRESS RATIO (0'2/0'8) Fig. 1.4.8. Difference in Average and Maximum Tangential Strains at Rupture for 3000-, 4000., 6000-psi Maximum Principal Stresses at Various Stress Ratios. 40 184 UNCLASSIFIED Y-26205 6000,/4000 6000,/6000 } TANGENTIAL STRESS AXIAL STRESS s 1500,/6000 Fig. 1.4.9. Crack Patterns of Outside Surfaces of Specimen Tested with a Maximum Stress of 6000 psi at 1500°F in Argon. Dye penetrant applied after test. 8561 ‘06 ¥39W3Ld3S INIONI dOI¥3d ANP PROJECT PROGRESS REPORT UNCLASSIFIED ORNL—-LR-DWG 31586 14 A 2 5 // 10 o, + O [ f_ 8 i /< o,+ o, e, =80, - —) F_ Z' R @) - 7 7 (? 0 b@ 4 2 / AN é 2 y. N }_ J g o . :. -1 T / W / \‘ g _o A N // /T \‘ -~ Kl Xefi:g(%— > [ ] -6 - ) \i I i/ \‘ s N\ [ ] \‘ —10 \ ® AXIAL STRAIN \ -4 O TANGENTIAL STRAIN N B RADIAL STRAIN €,=—(€, +€,) \? L © 4 p 1.33 { 075 05 025 O -025 -05 -075 -1 STRESS RATIO (g, /0) Fig. 1.4.10. Strain at Rupture vs Stress Ratio for Specimens Tested with a Maximum Principal Stress of 4000 psi. 42 UNCLASSIFIED ORNL-LR-DWG 31587 £ 10000 wy wy w - e 5 — W 5000 // 5 - z /' W & - — = o < a 2 2000 a a = 2 = % 1000 3 0,015 0.020 0.025 0.030 8 Fig. 1.4.11. Variation of B {Constant of Equation Re- lating Strain at Failure and Deviatoric Stress) with Maximum Principal Tensile Stress for Specimens Tested at 1500°F in Argon. under simple tension is also plotted against stress ratio in Fig. 1.4,12, This shows that specimens which obey Eq. 5 with the same maximum stress in the tension—~tension stress state fail with the same shear strain regardless of rupture life, Thus, for specimens of this type, the rupture strain in the tension—tension state is a function of the maximum principal stress only, It should be realized that al- though the total elongation in a particular direction may be less under combined tension—tension stresses, the shear strain or total deformation at rupture is the same. Specimens tested in the tension—compression stress state cpparenfly can sustain a much greater shear strain before fracture than those in the tension—tension state. As shown in Fig. 1.4.12, specimens tested under pure shear sustained twice the shear strain at fracture as the specimens tested in simple tension, STRAIN-FATIGUE STUDIES R. W. Swindeman Many of the failures encountered in engineering devices result from dynamic loads which are im- posed on the structure either mechanically or thermally. Most often, fatigue failures are con- sidered to be the result of rapidly fluctuating stresses which introduce damage in the material on a very microscopic scale. Thus in low-tempera- ture fatigue, no bulk plastic straining of the metal can be discerned. However, at elevated tempera- tures, where relaxation of stress can occur, meas- vrable amounts of plastic strain may be induced PERIOD ENDING SEPTEMBER 30, 1958 during each stress reversal. This is particularly true in the case of restrained structures when they are subjected to large thermal fluctuations, Quite often, under such conditions, the thermally induced stresses occur over relatively long time cycles and are large enough to exceed the elastic limit of the metal. Thus, in studying the behavior of materials loaded in this manner it is more meaningful to think of the metal as experiencing a number of strain reversals which ultimately lead to failure, rather than to attempt to base such failures on the inde- terminate stress state, Recent studies have revealed that the plastic strain history of a metal under dynamic load con- ditions provides extremely useful data for calcu- lating the metal life consumed and the service life remaining under expected operating conditions. This idea was conceived by Manson? and then dem- onstrated experimentally by Coffin® by thermally cycling stainless steel under conditions of restraint until failure occurred. The relationship between plastic strain and cycles to failure was found to be of the form N%€_ = K, where N is the cycles to failure, € _is the strain per cycle, and @ and K are constants which depend on the material and test conditions. Most of the subsequent work has been within the temperature range where the rate of work hardening was appreciably greater than the recovery rate. The work described below concerns metals at tem- peratures where creep and relaxation are the domi- nant factors. The fact that creep and relaxation are dominant in the temperature range of interest made it both feasible and attractive to substitute mechanical loads for thermal fluctuations as a means of producing strain cycles. The apparatus used is described in detail in a separate report.’ Most of the investigation was conducted with Inconel as the test material; but, a few tests were made with Hastelloy B and with beryllium, Thus, representative data can be compared for metals of quite widely different degrees of strength, duc- tility, and fabrication history. Conventional plots of the plastic strain per cycle, €p1 VS N, the number 4s, s, Manson, Bebavior of Materials Under Conditions of Thermal Stress, NACA-TN<2933 (July 1953), 3. F. Coffin, Jr,, Trans, Am. Soc. Mech. Engrs. 76, 931 (1954), 6C. R. Kennedy and D, A, Douglas, Plastic Strain Absorption as a Critetion for High Temperature Design, ORNL-2360 (April 17, 1958). 43 ANP PROJECT PROGRESS REPORT UNCLASSIFIED ORNL—LR—DWG 31536R 2.0 . Z & ’(7) S i o %S : el | | = 1.5 — L — | © 9 | | | = ’ : xr o y! e: eg ' P~ - — =( — = — ) FOR o, =CONSTANT a Yo €7, @ 2 @ G I I N N 1.0 J J T 1.5 S S R SR R 1 1 : 1 i ! ; ‘( i Tr + 0, ' Eeé’=5 (0'9—- u’z—”> FOR o,=CONSTANT o = >£ 1.0 ‘ | L. il - 0.5 Z8 278 ) FOR ¢, =CONSTANT }_ b 97% pure, with silicon being the pre- dominant impurity. DENSIFICATION OF BERYLLIUM OXIDE R. A. Potter Seven different compositions, given in Table 1.5.1, were mixed in a jar on rolls for 18 hr, and the material was then pressed at 20,000 psi in stee| dies, broken up, granulated by passing through a 30-mesh screen, and remixed for 2 hr, Right-cylinder shapes approximately 3‘/4 X 3/4 in. ]L. M. Doney, R. L. Hamner, and R. A. Potter, ANP Quar. Prog, Rep. March 31, 1958, ORNL-2517, p 45. Table 1.5.1. Compositions of Densified BeQ Bodies Composition (Parts by Weight) Composition L. M. Doney were made by pressing the materials at 15,000 psi. These shapes were then fired to six different tem- peratures (1650 to 1900°C in increments of 50°C). The furnace consisted of an induction-heated graphite tube with a positive flow-through of helium. In each case a fast firing schedule of approximately 2 hr was used. Screening tests indicated that composition 2 was superior to the others with respect to density. The densities of this body at the six test temperatures are compared with those of the 100% BeO control sample in Table 1,5.2. Magnesium oxide alone, as in composition 3, acted as a densifier, but at about 1700°C it began to volatilize rapidly and caused distortion of the shape of the cylinder. At 1650°C the density of composition 3 was found to be 2.9 g/cm?; how- ever, there was a slight distortion in the shape. For a further test, composition 2 was pressed into right cylinder shapes approximately ]3/8 x 1 in. in height. Forming pressures on these pieces were 6,000 and 10,000 psi. The slugs were fired in an oxidizing atmosphere to a temperature of 1500°C. Densities of these samples were in the range 2.67 to 2.75 g/cm>. Density values for similar 100% BeO control samples were 2,02 to 2.04 g/cm3, Specimens in the shape of bars approximately 5x 'll/4 X 3’/4 in. have been fabricated for additional testing. Table 1.5.2. Comparison of Densities at Various Temperatures of a 100% BeO Body and a Body =W%ith 94 wt % BeO, 5 wt % MgO, and 1 wt % B,C No. BeO* Mg0 Fe,0, B,C 1 100 2 94.0 5.0 1.0 3 94.0 5.0 4 94.0 1.0 5 98.0 1.0 1.0 6 98.0 1.0 7 98.0 1.0 *G. C. grade, Brush Beryllium Company. Sintering Density (g/cm3) Temperature ©C) 100% BeO 94% BeQ-5% MgO-1% B4C 1650 2.71 2.91 1700 2.79 2.88 1750 2.83 2.89 1800 2.84 2.88 1850 2.85 2.90 1900 2.85 2.86 ANP PROJECT PROGRESS REPORT OXIDATION RESISTANCE OF BORON-CONTAINING BERYLLIUM OXIDE BODIES R. L. Hamner Screening tests were also run to determine the oxidation resistance of hot-pressed boron-contain- ing BeO bodies being considered for GE-ANPD applications. The oxidation characteristics of the mixtures tested are compared in Fig. 1.5.1. The additives TaB, (9.35 wt %) and BN (2.29 wt %) not only have poor oxidation resistance above 1000°C but also apparently have very volatile oxidation products, Ta, O, and 8,0, which ac- count for the comparatively large weight losses. It was expected that the BeO-CrB, body would show an over-all weight loss upon oxidizing, be- cause the oxides of chromium, Cr203 and Cr03, tend to be volatile. A green deposit adjacent to the specimen indicated that a chromium oxidation product slowly volatilized along with the B,0; an over-all weight loss was detectable after about 125 hr at 1300°C. The borides of Ti (2.31 wt %), Zr (5.22 wt %), and Hf (9.25 wt %), had good oxidation resistance when incorporated in dense BeQ, as evidenced by appearance and over-all weight changes, with the ZrB,-BeO and HfB,-BeO combinations being slightly better than the TiB,-BeO combination. It is not known why the HfB,-BeO compact lost weight despite the known stability of the oxida- tion product, HfO,. The thermal stability of ZrB, in air during long- term oxidation testing at 1300°C is illustrated in Fig. 1.5.2. A specimen of pure BeO of comparable ORNL-LR-DWG 33729 Dl [ TiB2tBe0 - o o s e S -y 18,1+ Be0 R~ . — CrB,+ Be0 [ © HiBy+BeO 2 5 -05 . : ‘ ! & [ ' i v . o g o | i o 5 o ; = -1.5 ""‘—l‘b: N S - —_ | ! 5 1000°C| & I|300°c\ [\L Y oo L NG BN + BeO s ! ‘ ! T ———y -25 | - s [ - R— by ‘ 1 f 30 pt—t \JoB; * 220 ; N 0 25 50 75 10C 125 150 175 200 225 250 275 300 TIME (hr) Fig. 1.5.1. Results of Oxidation Tests of Boron-Con- taining BeO Bodies. 50 o) ORNL—;_R-DWG 33730 0.20 i 5 /”lo—’ & = o6 L ZrB, +BeD 1 Z o P T : g ¢ Pl o ] 3 P 5 008 7 I ' S .04 //‘ _ " i 1L o ¢ . re—— [, BeO e g -004 , ~0.08 [ o) 100 200 300 400 500 600 700 TIME (hr) Fig. 1.5.Z Results of Oxidation Tests of BeO and BeO + Zr82 in Air at 1300°C. density was tested with the ZrB,-BeO compact to give a more complete picture of weight changes involved, since BeO is somewhat volatile in the presence of water vaporat 1300°C, X-ray and chemical analyses showed the boron to be present after the long-term oxidation test. There was not sufficient ZrO, present as an oxidation product to be detectable by x-ray analysis, Information was received during this reporting period that the Brush Beryllium Company had dis- continued the manufacture? of ““Luckey S. P.” grade BeO, which has been used aimost exclusively in this laboratory for BeQ development work. |t was, therefore, necessary to establish fabrication characteristics of a new material representative of future supplies; Brush's G. C. grade BeO, which is derived from the sulphate and calcined at ap- proximately 1100°C, was received for this purpose. Specimens of G. C. grade BeO and BeO contain- ing 5.22 wt % of laboratory synthesized ZrB, were hot pressed at temperatures of 1550, 1650, and 1750°C under a pressure of 2000 psi for 15 min, The G. C. grade of BeQ was found to be much more reactive than the Luckey S. P. grade; notice- able compaction began at approximately 250 to 300°C lower (at about 1250°C). This difference in reactivity of the two materials can be explained by the difference in grain size, as shown by elec- tron photomicrographs, Figs, 1.5.3 (S, P. grade) and 1,5.4 (G. C. grade). Densities of pure BeQ obtained at 1550, 1650, and 1750°C were 92, 95, and 98%, respectively. Densities obtained for the BeO-ZrB,, mixture were approximately the same for all the temperatures, 2 . . . Private communication to R, L. Hamner. Fig. 1.5.3. Brush Beryllium Company's S. P. Grade BeO. 23,000X. Reduced 57%. being 96.7, 96.4, and 96.4%, respectively; thus ZrB,, appears to be somewhat effective as a densi- fier at the lower temperature. Above 1750°C a rather vigorous reaction occurred between the ZrB,-BeO mixture and the graphite die, the attack being severe enough to prevent removal of the piece from the die. Attempts were made to fabricate large hot- pressed blocks of the ZrB-BeO mixture, 3 x 3 x 1 in., for physical and mechanical properties meas- urements. A commercial grade of ZrB,, was used because of the large quantities of boride involved. Upon cutting the blocks, which were pressed at 1550°C and 2000 psi, an irregularity in color was noted that was roughly in the shape of a “bow- tie’’ and which approximated the temperature profile of the piece in the graphite die. The top and center PERIOD ENDING SEPTEMBER 30, 1958 W s b 1 Fig. 1.5.4. Brush Beryllium Company's G. C. Grade BeO. specimens shown in Fig. 1.5.5 are typical of this defect. It was not eliminated by time-temperature variations, although it was less pronounced at lower temperatures. Analysis of the ZrB, revealed a large quantity of iron (approximately 10%). When iron-free ZrB, was used, either laboratory synthesized or an acid-treated commercial grade, the appearance was uniform, as illustrated by the bottom specimen shown in Fig. 1.5.5. Six ZrB,-BeO blocks are being fabricated at ORNL for testing at GE-ANPD. ANP PROJECT PROGRESS REPORT UNCLASSIFIED PHOTO 31988 ARananaml T \\\-\‘\w\3 2 ° \NCHES Fig. 1.5.5. Hot-Pressed BeO-ZrB, Blocks. Top block and center blocks made with commercial ZrB,. Bottom block made with iron-free laboratory synthesized ZrB,. SulSl with caption) 52 PERIOD ENDING SEPTEMBER 30, 1958 1.6. NONDESTRUCTIVE TESTING J. W. Allen REMOTE X-RAY VIEWING J. W. Allen R. W. McClung Further studies were made of the use of x-ray sensitive Vidicon' in a closed-circuit television system for remote viewing of x-ray images. As reported previously, the sensitivity of the present selenium photoconductor Vidicon system is con- siderably less than that which can be obtained with existing film techniques. The primary ad- vantages of this system (in contrast to those con- sisting of closed-circuit television applied to fluoroscopy) are the inherent high resolution, which approaches the resolution attainable with a fine-grained radiographic film, and the inherent magnification of images. Recently it was discovered that the sensitivity of the selenium photoconductor to x-radiation could be improved in many situations by admitting visible light to the photoconductor surface. Although the phenomenon is not well understood, it appears that the sensitivity is increased merely by raising the level of total incident radiation, including both x-radiation and visible light. This effect is readily observable and has been used to increase the con- trast sensitivity by as much as a factor of 3. The wave length of the enhancing light does not seem to be critical, as evidenced by the use of both white fluorescent lamps and incandescent lamps as the light source. By using this light-enhancing technique, thick- ness changes of about 2,5% can be detected in s/a-in.-fhick aluminum. This indicates that practical inspections could be accomplished at contrast sensitivities of 5 to 6%. Although this figure is considerably lower than that attainable with film techniques, it compares favorably with those pub- lished for other television systems which use fluorescent screens. Further investigation of the light-sensitizing technique is planned. In addition the use of photo- conductor materials other than selenium is being studied. DUPLEX TUBING R. W. McClung The increased interest in the fabrication and use of duplex tubing has created a need for adequate inspection techniques for this configuration. The problems which are common to the inspection of conventional tubing, such as dimensional gaging and the detection of cracks, seams, laps, and other discontinuities, are also present in duplex tubing. In addition, an evaluation must be made of the bond quality between the layers, and measurements should be made of the individual layer thicknesses. To detect the more common discontinuities, the conventional techniques? of encircling-coil eddy- currents, pulse-echo vltrasonics, radiography, fluorescent penetrants, and visual inspection should be adequate for complete evaluation. Also some of these techniques may provide methods for determining some of the properties peculiar to duplex tubing. The use of the pulse-echo ultra- sonic technique has detected the presence of lami- nations in welded and drawn tubing.3 The detection of such laminations is probably dependent upon the size and location of the lamination, the ultra- sonic frequency, and the dimensions of the tube. At present, not enough is known about the mecha- nism of detection of these laminations to assure a high level of confidence in their detection under all conditions. Many of the areas of nonbonding will resemble simple laminations, and, under proper conditions, may be located by the pulse- echo ultrasonic technique. If a gross separation exists between layers, radiography may offer a means for detection, either as a function of the decreased metal thickness due to the voids when passing the radiation perpen- dicularly through the lack-of-bond, or by observing the poor bond in profile with the radiation passing tangentially through the separation between layers. However, as the separation between layers de- creases, the confidence level of such an examina- tion decreases rapidly. The resonance ultrasonic technique is being evaluated as a means for the detection of non- bonded areas. This technique has been used in 'R. B. Oliver and J. W. Allen, ANP Quar. Prog. Rep. 2R, B. Oliver et al., ANP Quar. Prog. Rep. Dec. 31, 1956, ORNL-2221, p 260. 3R. B. Oliver, R. W. McClung, and J. K. White, Am. Soc. Testing Materials, Spec. Tech, Publ. 223, 62-79 (1957). 53 ANP PROJECT PROGRESS REPORT the past to measure wall thickness. On duplex tubing with a well-bonded interface, the total wall thickness of both layers would be indicated as a function of a fundamental or harmonic resonance of an ultrasonic frequency. Under proper conditions, if a lack of bonding exists, this indicated thick- ness would be that of the clad or outside layer only. However, to achieve this detection of non- bonding, it is necessary to make a careful selec- tion of frequency range and calibration to avoid confusion of indications. A few small batches of duplex tubing have been examined to determine the bond quality. These include approximately 14 ft of 0.504 x 0.042 in. tubing and 24 ft of 0.375 x 0.036 in. tubing. These lots were fabricated by cladding about 0.025 in. of low-carbon steel on types 304 and 347 stainless steel, respectively. Radiography, resonance ultra- sonics, and pulse-echo ultrasonic tests revealed the presence of gross discontinuities in the 0.504-in.-dia tubing. Metallographic sectioning of several typical defects revealed considerable lack of bonding, with gross voids and separation pre- valent. The worst bond condition detected is shown in Fig. 1.6.1. The transverse cracks seen in Fig. 1.6.2 are probably responsible for the many indications noted in the pulse-echo ultra- sonic examination. The same ultrasonic and radiographic techniques as those used on the 0.504-in.-dia tubing, when applied on the 0.375-in.- dia tubing indicated the presence of a good bond, similar to that shown in Fig. 1.6.3. The resonance ultrasonic technique seems to provide an evalua- tion for bond quality which is the best and most consistent of any of the current techniques. Addi- tional development effort is needed and will be made to provide a high level of confidence for the complete evaluation of duplex tubing. METAL IDENTIFICATION METER R. A. Nance J. W. Allen A transistorized model of the Metal Identification Meter? (MIM) has been developed. Although the internal operation has been altered, the new instru- ment also utilizes eddy-current methods for identi- fication and retains the simple external operation “R. B. Oliver, J. W. Allen, and R. A. Nance, ANP Quar. Prog. Rep. June 30, 1957, ORNL-2340, p 256. UNCLASSIFIED 26826 Fig. 1.6.1. Very Poor Bond Between Type 1030 Steel Cladding and Type 304 Stainless Steel in }-in.-OD, 0,042- in.-Wall Tube. 54 PERIOD ENDING SEPTEMBER 30, 1958 Fig. 1.6.2. Transverse Cracks at Bond Interface Between Type 1030 Steel Cladding and Type 304 Stainless Steel in ‘/Z-rn.-on, 0.042-in.-Wall Tube. 4 S INCHES Fig. 1.6.3. Good Bond Between Type 1008 Steel Cladding and Type 347 Stainless Steel in %-in.-OD, 0.036sin.- Wall Tube. 55 ANP PROJECT PROGRESS REPORT of the vacuum-tube model. The MIM-Mark 1I, shown in Fig. 1.6.4, is 7% x 3% x 2% in., and it weighs approximately 20 Ib, representing an 80% reduction in size and weight. In developing the transistorized model, several changes were made which greatly increased its versatility. These are: (1) the instrument has been made truly portable by the incorporation of a battery power supply, (2) the range of metals identi- fiable has been expanded to include all ferromag- netic and nonferromagnetic metals, (3) the effec- tive depth of penetration of the eddy-currents has been decreased by increasing the operating fre- quency and thereby reducing the minimum metal thickness necessary for accurate identification, and (4) the diameter of the probe coil has been de- creased to reduce the error produced when a flat reference standard is used in conjunction with the identification of a specimen having a curved sur- face. A block diagram illustrating the internal opera- tion of this instrument is shown in Fig. 1.6.5. The frequency of the probe oscillator is determined by the inductance of a probe coil, which, in tum, is determined by the conductivity and permeability of the specimen. The frequency of the tuning oscilla- tor is controlled by the dial on the front of the instrument. The signals from these two oscillators are heterodyned to produce a signai having a fre- quency equal to the difference of the frequencies UNCLASSIFIED ORNL=LR-DWG 302938 REGIENCY i, L, ThansiTons N ] TUNING ETERODYNE | (/=%)| AMPLIFIER ILLATOR \| ,'-0355(:.:“5 wl & MIXER |AND DETECTOR] 20k 12140 TRANSISTOR PROBE OSCILLATOR PROBE COIL =200 320 ke 774 specmen | Dflmwmm METER Fig. 1.6.5. Block Diagrom Describing Operation of Metal Identification Meter-Mark IL. UNCLAsSIFIED Y220 Fig. 1.6.4. Metal Identification Meter-Mark Il of the two original signals, The designs of the amplifier and detector are such that only signals of a narrow band of frequencies, centered about 120 kc, will be amplified, detected, and presented on the meter, Since the eddy-current flow in a specimen is a function of its conductivity, magnetic permeability, and geometry, it is possible by controlling geom- etry factors to utilize the conductivity or con- ductivity and permeability characteristics of a metal as identifying features. This is accomplished by tuning the instrument for @ maximum meter de- flection with the probe on a reference standard. The difference frequency between the two oscil- lators is then 120 kc. If the probe is then placed on a specimen having a different conductivity or permeability, the frequency of oscillation of the probe oscillator will shift so that the difference frequency between the two signals is no longer 120 kc and the meter will no longer indicate a maximum. |f the difference between the metals is great enough, the meter will cease to indicate altogether, The response of this instrument to changes in conductivity in nonferromagnetic materials is illustrated in Fig. 1.6.6. The dial settings 565, 545, and 495 represent the tuning dial settings that were necessary to obtain maximum meter deflection when the probe was placed on samples of Hastelloy B, Inconel, and type 316 stainless steel, respectively. It is evident that, if the instrument is tuned with the probe on the Inconel sample (dial setting 545), the instrument will not respond when the probe is placed on the other two samples. [t can also be seen in Fig. 1.6.6 that the slight variations of conductivity normally found among samples of the same alloy will not ap- preciably decrease the meterreading from a maximum. UNCLASSIFIED ORNL-LR-DWG 30634R 10 T T T | T T T DIAL SETTING 565 ' ' A A LT | ; 275 =—CIAL SETTING 545 "~ [ f—F - 2 [ ‘ / | w DIAL SETTING 495 - -+t ' [1d 5 ‘ - Y U o { w = ! w i 225 - / — \ L Y g LN 1.2 1.4 16 1.8 2.0 2.2 2.4 2.6 2.8 3.0 CONDUCTIVITY {7 IACS) Figo Mark Tuning Dial Setting. 1.6.6. Il to Conductivity Changes as a Function of Response of Metal Identification Metere PERIOD ENDING SEPTEMBER 30, 1958 It is important to note that this instrument cannot be used to separate two nonferromagnetic metals of differing composition which have the same conduc- tivity or overlapping ranges of conductivities. Ac- curate identification can only be made when it is known that none of the metals in the group to be identified have the same or overlapping conductivi- ties. A similar problem is encountered in identify- ing ferromagnetic materials, since different com- binations of conductivity and permeability can produce the same instrument response. Variations in the geometry of a specimen can adversely affect its identification when a thick, flat reference standard is used. The extent of the error produced by the curvature variations on the surface of a specimen is illustrated in Fig. 1.6.7. For all practical purposes an item may be con- sidered to present a flat surface to the probe if its radius of curvature is greater than 1.5 in. Variations in the thickness of the specimen will produce an error in identification only if the speci- men is thinner than the effective depth of the eddy- current penetration, which varies with the con- ductivity and permeability of the part and the operating frequency of the probe oscillator. The variations in meter reading caused by thick- ness variations of a specimen can be used as a measure of thickness. This has been satisfactorily accomplished in the gaging of 3/4 x 0.05 in. type 304 stainless steel tubing with an accuracy of 2'/2%. This instrument can also be used to separate specimens from the same heat of an alloy which have been subjected to different mechanical and thermal treatments, since they produce changes in the conductivity and in some cases the magnetic permeability of the part. [n addition, this instru- ment can be used to locate cracks normal to the inspected surface, since they interrupt the eddy- current flow and cause an apparent conductivity change in the part. UNCLASSIFIED ORNL-LR-DWG 30295 50 Y \ g a0} \ —‘ . g £ 30 | I % T 820 . e - _ bal \ } [ 10 — . ™ — ———t N O \ 0 0.5 1.0 1.5 2.0 2.5 3.0 @ CYLINDER RADIUS {in.) Fig. 1.6.7. Effect of Cylinder Radius on Error of Metal Identification Meter-Mark Il. 57 ANP PROJECT PROGRESS REPORT 1.7. METALLOGRAPHY R. J. Gray RESULTS OF METALLOGRAPHIC EXAMINATION OF SMALL SEMICIRCULAR FUSED-SALT FUEL-TO-NaK HEAT EXCHANGER OPERATED AT HIGH TEMPERATURES R. J. Gray J. E. VanCleve A metallographic examination has been com- pleted of a 20-tube semicircular heat exchanger fabricated of Inconel in which a fused salt flow- ing around the tubes exchanged heat to NaK flow- ing in the tubes. Details of the construction and operation of this Inconel heat exchanger were pre- sented previously.'*? Testing of this heat ex- changer was terminated after 1139 hr of high- temperature operation at 1200°F and above because of a leak of NaK into the fused salt. At the time of termination the heat exchanger had operated at design conditions for 258 hr with a temperature differential from fused-salt inlet to outlet of 350°F (1600 to 1250°F) and a temperature differential from the NaK inlet to outlet of 430°F (1070 to 1500°F). There had been 142 hr of operation under transitional conditions and 739 hr of isothermal operation. The heat exchanger had experienced 194 thermal cycles between isothermal operation at 1200°F and operation at design conditions. The general area of the failure was readily located after the shell was removed because of a change in the color of the fluoride salt from green to gray in the NaK outlet header region. The gray material ignited during sectioning of the header, and the presence of NaK mixed with the fluoride salt was thus evident. Two of three fractured Inconel tubes found in the high-temperature header area are shown in Fig. 1.7.1. The cracks were radial and were in the so-called ““tension” side of the tubes. The cracks were found in the tubes with shortest tube befld-'o-?ube sheet length. This length ranged from 1) % to 3/ in. The side of the tube that cracked was in fms:on because of the temperature differ- ence between the tube and the shell. 1). C. Amos et al., ANP Quar. Prog. Rep. March 31, 1957, GRNL- 22740 o 43, 2y C. Amos, R. L. Senn, and D. R. Ward, ANP Quar. Prog. Rep. Szpl 30, 1957, ORNL-2387, p 38 and esp Fig. 1.2.7, 58 The typical condition of the tube wall in all three ruptured tubes is shown in Fig. 1.7.2. The tube wall shifted laterally 0.002 in. after the largest crack completely penetrated the wall. The other large crack completely penetrated the tube wall but it did not open as did the first. A third crack may be seen that propagated from the braze alloy surface. This is the only instance of a small crack penetrating the material, and it suggests the UNCLASSIFIED 3 823 0.101n./DIV | | Fig. 1.7.1. Cracks in Two of Three Cracked Inconel Tubes Found in the High-Temperature Header Area of o Fused Cracks are in the tension side of the tubes. Semicircular Salt-tosNaK Heat Exchonger. Fig. 1.7.2 Typical Cracks Found in Tension Side of a Tube. Etchant: possibility that a mechanism other than the progres- sion of grain-boundary voids had operated to aid in the fracture of the tubes. The largest of the cracks shown in Fig. 1.7.2 is shown at a higher magnification in Fig. 1.7.3. The surface of the tube which was exposed to the fused salt was generally corroded and leached to a depth of 0.007 in. The grain-boundary voids which were found in large numbers in similar areas of other heat exchangers of this type were not present in this unit, The smaller of the large cracks shown in Fig. PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED Y-26166 10% oxalic acid. 100X. 1.7.2 is shown in Fig. 1.7.4 ot a higher magnifica- tion. The crack follows the grain boundaries, and the greater width of the crack at the inner, or NaK contacted, surface indicates that it originated at the inner surface. A few grain-boundary voids may be seen that extend from the penetration. The grain-boundary voids do not seem to be as numer- ous or large as those found in previously examined heat exchanger tubes. The maximum general cor- rosion found in the heat exchanger is shown in Fig. 1.7.5. On the fused-salt side the attack is to a depth of 0.006 in. and on the NaK side the attack is to a depth of 0.002 in. 59 ANP PROJECT PROGRESS REPORT -{\: UNCLASSIFIED SN Y-26165 Fig. 1.7.3. The Larger of the Cracks Shown in Fig. 1.7.1 at a Higher Magnification. General corrosion and leaching to a depth of 0.007 in. may be seen on the surface that was exposed to the fluoride salt. Etchant: 10% = oxalic acid. 200X. UNCLASSIFIED |2 : Fig. 1.7.4. The Smaller of the Large Cracks Shown in Fig. 1.7.2 at o Higher Magnification. Complete pene- 3 tration of the tube wall may be seen. Small grain-boundary voids subtend the large crack. Etchant: 10% oxalic acid. 200X. 60 Fig. 1.7.5. Maximum General Corrosion: * oxalic acid. 200X. 0.006 in. on Salt Side and 0,002 in. on NaK Side. Etchant: 10% Part 2 CHEMISTRY AND RADIATION DAMAGE 2.1. MATERIALS CHEMISTRY W. R. Grimes Chemistry Division PREPARATION OF CHARGE MATERIAL FOR REDUCTION TO YTTRIUM G. J. Nessle A pilot plant has been constructed that will provide for the monthly production of about 200 Ib of massive ytrium metal. The Chemistry Division is responsible for the preparation of the suitable fluoride mixture, either LiF-Mng-YF3 or LiF-YF,, and the Metallurgy Division will carry out the re- duction, distillation, and sparge-melting steps that will yield the final product. The pilot plant was completed during August, and the first batch of fluoride-salt mixture (LiF-YF,) has been prepared. Phase equilibrium studies of the LiF-YF, system, as reported below, have indicated a single eutectic with 19 mole % YF , that has a melting point of 682°C. Therefore an attempt will be made to prepare low-oxide-content yttrium from this binary mixture and thus eliminate the Mg-Y alloying step and subsequent distillation of Mg from the alloy that would be required if the LiF-MgF ,-YF_ mixture were used, If this attempt is unsuccessful, it will be necessary to revert to the ternary system. The metal reduction operation is described in Chap. 1.1 of this report. Conversion of Y,0, to YF, J. Truitt About 400 Ib of YF, was produced by direct hydrotluorination of dry Y,0, in a large-scale (250-1b capacity) converter. In addition, about 100 Ib of YF, was produced in a small, pilot-scale unit (10 Ib per batch). These units are operated at approximately 1100°F, and the raw material is treated for about 10 hr with HF. Analyses of the batches produced in the pilot-scale unit are given in Table 2. 1.1, Preparation of MgF, J. Truitt No satisfactory grade of MgF . was available commercially, and therefore six 10-lb batches of MgF , were prepared by dry hydrofluorination of MgO. Previous results had shown that by extended treatment at 600°C a product of better than 99.5% could be obtained. However, extended hydro- fluorination is not economically desirable if, as was later shown, 95% assay material is satisfactory for charging to the final molten salt purification step. The range of product analysis and conversion efficiency of the six batches is given in Table 2,1.2. It is now possible to purchase a satisfactory grade of MgF, and sufficient material has been ordered for several months of operation. Preparation of the LiF-MgF,-YF ; Mixture J. Truitt J. E. Eorgan Several small-scale batches of the ternary mixture LiF-MgF ,-YF, were purified for yttrium-metal pro- duction. Although it was known that the raw materials contained oxides to some extent, the standard 3-hr hydrofluorination time used in other Table 2.1.1. Analyses of Batches of YF3 Obtained by the Conversion of Y203 with HF in a Pilot-Scale Unit Minimum Maximum Mean (%) (%) (%) Analytical results* Yfl'rium 59.7 6]-3 60-7 Fluorine 38.1 39.0 38.7 Oxygen 0.09 1.5 0.7 Conversion efficiency* 97.2 100.0 99.0 *Individual analyses for Y and F and conversion effi- ciencies are not necessarily from the same batch. Table 2.1.2. Analyses and Conversion Efficiencies Obtained in the Batch Conversion of Mg0 to MgF2 Minimum Maximum Mean (%) (%) (%) Analytical results* Magnesium 39.0 40.0 39.5 Fluol’ine 59.4 60.7 60.3 Conversion efficiency* 95.0 99.6 97.7 *Individual analyses for Mg and F and conversion effi- ciencies are not necessarily from the same batch, 65 ANP PROJECT PROGRESS REPORT salt purifications was used on the ternary mixture in order to test the new method of oxide deter- mination in fluoride salts and to determine the reproducibility of oxide removal. The analytical results obtained for ten batches are presented in Table 2.1.3. Petrographic and x-ray diffraction examination of these mixtures did not detect oxides. As can be seen from the data, however, the chemical analyses indicated approximately 1000-ppm oxygen in the salt, There is some reason to believe that the oxide analyses gave high results because of sampling techniques, but, on the basis of these analyses, the hydro- fluorination time will be increased gradually until a product purity of 500-ppm oxygen or less can be assured. Table 2.1.3. Analyses of Ternory Mixtures of LiF-Mu_:ng-YF3 Analytical Results* Constituent Minimum Moximum Found Found Mean Y 30.6% 33.5% 32.3% Mg 8.70% 9.23% 8.95% Ni 40 ppm 145 ppm 75 ppm Cr 30 ppm 200 ppm 100 ppm Fe 80 ppm 200 ppm 150 ppm 0, 1000 ppm 1500 ppm 1100 ppm *Individual analyses are not necessarily from the same batch. Extending the hydrofluorination will have a marked effect on the processing time per batch of mixed salt because there will be increased dis- solution of the reactor vessel, usually nickel, and of the accumulated impurities left from previous batches. Therefore, care must be taken to use only the HF needed in order to prevent the processing time (or cost per batch) from becoming unrea- sonable, if possible, PHASE EQUILIBRIA IN THE SYSTEM LiF-YF3 R. E. Thoma Determinations of the character of the phase equilibria occurring in the system LiF-YF have been deterred by the presence of considerable 66 amounts of a phase which is probably an oxyfluoride of yttrium, This phase appears as a biaxial negative material with an optic angle of 10 deg, and it has refractive indices just above and below 1.700. All samples of YF, used previously had been prepared by solid-state hydrofluorination of Y,0,. R ecently a LiF-YF, batch was hydrofluorinated in the liquid state in order to minimize the content of the oxyfluoride., This batch was subsequently analyzed chemically and used as the beginning m aterial for several thermal-gradient quenching experiments, The quenched samples from these experiments were apparently free of oxygen- containing phases. The quench results are thus believed to be more nearly representative of L iF-YF, equilibria than any results previously available, A portion of the phase diagram of the system has been constructed from these results and is shown in Fig. 2.1.1. A single eutectic occurs in the system at 19 mole % YF,; mp, 682°C. This melting temperature is lower than any of the values previously reported, ! and it is believed to be the most accurate value yet attained., 'R. E. Thoma, ANP Quar. Prog. Rep. March 31, 1958, ORNL-2517, p 54. TEMPERATURE (°C) UNCLASSIFIED ORNL-LR-DWG 33937 900 LIQUID 800 \ LiF + LIQUID YLIF-XYFy + LIQUID 700 3 o ;[ o - LiF + YLiF - XYFy 600 0 10 20 30 YF3 (mole %) Fig. 2.1.1. A Portion of the System LiF-YF, (0~30 Mole % YF3)0 EXTRACTION OF LITHIUM METAL IMPURITIES WITH MOLTEN SALTS G. M. Watson J. H. Shaffer A tentative procedure for purifying lithium metal by extracting oxide, nitride, and carbide impurities with a eutectic mixture of lithium halide salts, such as LiF-LiBr (12-88 mole %; mp, 453°C) or LiF-LiCl (20-80 mole %; mp, 485°C), was described pre- viously.? The effectiveness of the method depends on the differences in solubilities of impurities in the salt and metal phases. Indirect measurements in this laboratory indicate that the solubility of Li,O is relatively high (over 1 mole %) in molten salts. The solubility of the oxide in liquid lithium is not known, however, because of difficulties en- countered in sampling and analysis, even though some progress has been recently reported? in the determination of oxygen in lithium, It is expected but not necessarily assured, how- ever, that the differences in the solubilities of the impurities in the salt and metal phases are favorable toward their accumulation in the salt phase and that a suitable extraction method for lithium purification may be developed. Five experiments were performed on two different batches of lithium in attempts to gain familiarity with the procedure and to determine whether a cor- relation could be obtained between the number of extractions on a given batch of lithium and the amounts of impurities extracted. The results of these experiments are summarized in Table 2.1.4. No particularly good correlation has been obtained to date between the number of extractions performed on a given batch of lithium and the concentration of 2G. M. Watson and J. H. Shaffer, ANP Quar. Prog. Rep., Marchk 31, 1958, ORNL-2517, p 55. 3N. I. Sax and H. Steinmetz, Determination of Oxygen in Lithium Metal, ORNL-2570 (Oct. 15, 1958). Table 2,1.4. PERIOD ENDING SEPTEMBER 30, 1958 impurities extracted. |t is true that the second ex- traction in each of the series shows a considerable lowering in concentration of the impurities extracted as compared with the initial concentration; however, the concentration of 50 ppm shown is probably in error, since in this experiment a considerable amount of bismuth-lithium alloy was trapped in the salt. Substantial corrections determined by measuring the hydrogen evolved from the salt when treated with water had to be applied. Thus the value of 50 ppm was obtained as the net result of the difference between two relatively large numbers. Accordingly this result can be dis- counted for the present. The result of the third extraction of batch number 1 appears to be totally due to nitrogen impurities, as given by chemical analysis of the salt. It is not immediately apparent why no oxygen was found (by difference) unless the analytical results for the nitrogen determination are erroneous. For future experiments greater care will be exercised in minimizing contamination in the experimental assembly. Also the lithium metal will be heavily contaminated with known amounts of oxide, introduced as CuQ, in order to attempt to determine on a much firmer basis whether the partition coefficient for the oxide extraction is favorable for the accumulation of impurities in the salt phase. Recent improvements in analytical techniques have made it feasible to analyze lithium samples before and after the salt extraction, and it is no longer necessary to depend entirely on the complex procedures required for analysis of the salt phase. The liquid lithium samples will be obtained as emulsions of lithium in paraffin. 4 4A, S. Meyer, Jr., and R. E. Feathers, ANP Quar. Prog. Rep. March 31, 1958, ORNL-2517, p 57. E xtraction of Impurities from Liquid Lithium by Molten LiF-LiX Eutectic Mixtures Total Impurities Batch Extraction Salt Extracted (Calculated Nitrides No. Number Mixture as ppm 0) {as ppm N) 1 1 LiF-LiBr 400 Not determined 2 LiF-LiBr 50 <50 3 LiF-LiBr 484 500 2 1 LiF-LiCl 2860 <20 2 LiF-LiCl 467 <20 ANP PROJECT PROGRESS REPORT 2.2. ANALYTICAL CHEMISTRY J. C. White Analytical Chemistry Division DETERMINATION OF OXYGEN IN FLUORIDE SALTS A. S. Meyer, Jr. G. Goldberg The method for the determination of oxygen in fluoride salts by fluorination with KBrF ,, as de- scribed in a previous report, was applied success- fully to determinations of oxygen in the mixed salt LiF-MgF -YF ;. Seven salt mixtures were analyzed during this period. The results for duplicate samples are presented below: Sample No. Oxygen Content (%) Al 0.11, 0.11 A2 0.21, 0.22 A3 0.11, 0.10 A4 0.11, 0.10 AS 0.14, 0.15 Aé 0.24, 0.26 A7 0.11, 0.13 The coefficient of variation of the results is 6%. Since Sheft, Martin, and Katz? had reported that KBrF, was too basic to react with sufficient satis- faction with an oxide as basic as MgO and had recommended the use of an acidic addition com- pound, SbF BrF2, for such oxides, the validity of the results ?or oxygen in LiF-MgF ,-YF , was cor- roborated by analyzing standard samples of MgO and Y,0,. The recovery of oxygen from these oxides exceeded 98% with excellent precision (coefficient of variation, 2%). Highly refractory MgO, however, did not react appreciably with KBrF .. The method described is now being used for routine determi- nations of oxygen in the mixed salt and also to determine the extent of conversion of oxides of magnesium and yttrium to the fluorides. TA. S, Meyer, Jr., and G. Goldberg, ANP Quar. Prog. Rep. Sept, 30, 1957, ORNL-2387, p 150. 2|, Sheft, A. F. Martin, and J. J. Katz, J. Am. Chem. Soc. 78, 1557 (1956). 68 DETERMINATION OF YTTRIUM AND MAGNESIUM IN LiF-Mng-YF3 J. P. Young R. F. Apple A potentiometric method for the determination of both yttrium and magnesium in LiF-MgF ,-YF . was developed and evaluated. Yttrium and magnesium form stable complexes with ethylenediaminetetra- acetic acid (EDTA) and can be determined by direct titration by applying the technique of Reilley and Schmid,® who use a modified mercury electrode. The electrode system is essentially an amalgamated gold wire as the indicator electrode and calomel as the reference electrode. Both yttrium and magnesium are determined in the same sample because of the difference of the stability constants of their respective EDTA com- plexes. Yttrium, the more stable complex, is determined at a pH of 4.0 £ 0.5. Then the pH of the sample solution is increased to 9.7 £ 0.3, and magnesium, the less stable complex, is determined. An example of the resultant titration curve is shown in Fig. 2.2.1. When titrating yttrium and magnesium with 0.01 M EDTA, at least 5 mg of yttrium and at least 1 mg of magnesium should be present in order to obtain a satisfactory potential break at the end point. Although macro amounts of iron, nickel, and chromium would definitely interfere with these determinations, the method is quite insensitive to the micro amounts of these corrosion products that would be expected in routine samples of LiF- MgF -YF . Up to approximately 400 ug of fluoride ion can be tolerated in the sample solution without affecting either determination. The titrations can be per- formed in either a nitrate or sulfate medium; large concentrations of chloride ion interfere with the response of the mercury electrode. Typical results for the determination of yttrium and magnesium in synthetic standard sample solutions which con- tained nitric or sulfuric acid are given in Table 2.2.1. 3C. N. Reilley and R. W. Schmid, Anal. Chem. 30, 947 (1958). The precision of the method is about 1%. The results of these determinations are accurate to with- in 2%. This method is considerably less time- consuming than the gravimetric methods that were previously used for the determination of these metals. The procedure has also been applied to the determination of yttrium in yttrium oxide and in yttrium fluoride. UNCLASSIFIED ORNL-LR-DWG 33938 R C T T 1T Mg CONTENT: 2.18 mg ,/ pH: 9.7 £ 0.3 [ -Sva I i [ | POTENTIAL (WITH REFERENCE TO A SATURATED CALOMEL ELECTRODE) 6 8 10 12 19 16 18 20 22 24 26 0.01# EDTA (mi) Fig. 22,1, Results of Potentiometric Titration of Yttrium and Magnesium with Ethylenediominetetraacetic Acid. PERIOD ENDING SEPTEMBER 30, 19 Table 2.2.1. Typical Results for the Determination of Yttrium and Magnesium in Solutions Which Contain Yttrium, Magnesium, and Lithium by Medns of a Potentiometric Titration with Ethylenediaminetetraacetic Acid 58 Yttrium (mg) Magnesium (mg) Present Found Present Found With 1 M Sulfuric Acid Present 7.24 7.35 1.45 1.44 7.30 1.44 10.9 10.9 2.18 2.16 11.1 2.17 14.4 14.6 2.90 2.84 14.4 2.89 With 0.2 M Nitric Acid Present 7.50 7.58 1.45 1.49 7.58 1.49 11.3 11.4 2.18 2,15 11.3 2.16 15.0 15.3 2.90 2.84 15.2 2.87 69 ANP PROJECT PROGRESS REPORT 2.3. RADIATION DAMAGE G. W. Keilholtz Solid State Division ETR IRRADIATION OF MODERATOR MATERIALS FOR USE AT HIGH TEMPERATURES W. E. Browning, Jr. R. P. Shields J. E. Lee, Jr. Experimental apparatus is being prepared with which to test the stability of yttrium hydride and beryllium oxide at temperatures up to 1800 in the high-level irradiation environment of the Engi- neering Test Reactor (ETR), where the gamma-ray heating is approximately 25 w/g. Temperature gradients will be induced in the specimens, and the reactor programing will provide thermal cycling. The yttrium hydride samples will be checked for changes in crystal structure, as well as for the effects of stresses resulting from temperature gradients. Before the samples of the moderator materials are irradiated, x-ray examinations will be made of the yttrium hydride samples in order to determine the crystal structures, the modulus of elasticity and shear modulus of each sample will be measured, marked regions on the samples will be photographed, and exact measurements of the pieces will be made. The radioactivities of the samples should be low enough for the examinations and measurements to be repeated after irradiation. The heat generated in the irradiated capsules will be conducted across a small annulus by a helium- argon or a helium-nitrogen mixture which will be essentially static as far as heat transfer is con- cerned. There will be just enough flow to permit changing the mixture in order to maintain a constant temperature at the outside surface of a capsule. Helium has about four times the heat transfer capacity of argon and will make up the largest part of the gas mixture. Since the intensity of gamma-ray heating in the reactor is uncertain it is necessary to provide means for raising the experimental equipment to a region of lower gamma-ray intensity if the heating is found to be too great. In order to accommodate the necessary raising devices, the apparatus for the initial experi- ment will include only one yttrium hydride and one beryllium oxide capsule. The beryllium oxide sample will contain three right cylinders 0.636 in. in diameter and 1 in. in 70 height. These will be placed end to end in a type 410 stainless steel capsule. The yttrium hydride sample is similar, but the diameter of each cylinder is 0.800 in. and the cylinders will be enclosed in type 430 stainless steel. The yttrium hydride cylinders and the steel capsule prior to assembly are shown in Fig. 2.3.1. As may be seen, one cylinder, which will be at the top of the assembly, has a weil for a thermocouple with which to monitor the central temperature. The top beryllium oxide cylinder temperature will be monitored similarly. The assembled capsules and the tubes which pro- vide the gas annuli are shown in Fig. 2.3.2. Thermocouples will be attached to the sides of the capsules midway between the ends. In order to pass these thermocouples through the thin gas annulus, sheathed, swaged magnesia thermocouple leads will be silver soldered into the small metal plugs shown in Fig. 2.3.2. The thermocouples will be spot- welded to the capsule. Flexibility in the thermo- couple wires where they pass through the gas annulus will be provided by helical coils. After the outer tube assembly is completed, it will be encased in an aluminum spacer that will provide a ]/] -in. water annulus and will also anchor the sheathed thermocouple leads and gas tubes. The spacer unit will then be fitted into an aluminum tube which will be attached to a 1-in.-dia stainless steel pipe. This pipe, which will extend to a point about 10 ft above the ETR lattice, will support the experi- mental apparatus. All thermocouple leads and gas lines will be inside the pipe and will leave the reactor cell through a 1-in.-dia flexible steel tube that will connect with the support tube and the flange on the reactor. The steel cable attached to the support pipe and a drum near the reactor flange will provide the means for raising the apparatus to a position of lower gamma-ray flux if the termperature becomes too high. The design work on the thermocouple installation was the major developmental effort required for this experiment. Each thermocouple must be capable of operating at a high thermal stress in the gas annulus. Also, the sheath which encloses the leads and is sealed at the junction with the tube that forms the gas annulus must be capable of withstanding the high pressure of the reactor cooling water. PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED PHOTO 44085 Fig. 2.3.1. Yttrium Hydride Cylinders and Steel Capsule Ready for Assembly and lrradiation in the ETR at High Temperatures. (Secret with caption) UNCLASSIFIED PHOTO 44558 Fig. 2.3.2. Capsule Irradiation Assemblies Showing Tubes Which Form Gas Annuli. A series of mockup devices was used in designing a thermocouple arrangement and developing the technique for installing the thermocouples, and the operating conditions were simulated as nearly as possible in the mockup tests. The mockup used to test the thermocouples is shown in Fig. 2.3.3. The heating element (bottom of Fig. 2.3.3) is a 10-mil- wall Inconel tube attached to copper electrodes. At 900°F, approximately 2 kw of heat can be conducted from this element across the 10-mil annulus to a water jacket. The thermocouple wires are spot- welded to the heating element, and the Inconel sheath containing the wires extends through a small cap which anchors the sheath to the water jacket. Cycling the heating element from room temperature to 900°C flexes the thermocouple wires; the life of 71 ANP PROJECT PROGRESS REPORT Fig. 2 for High-Temperature Irradiation of Moderator Material s. the thermocouple is a function of the number of such cycles. It was found that annealing the wires after removing the sheath and before bending them greatly increased the life. This mockup device also served to develop a temperature control method. With argon flowing through the annulus at a very low rate, a Moore controller, activated by a thermocouple, introduces helium into the argon in varying amounts to increase or decrease the heat conductivity of the gas. Varying the conductivity of the gas mixture allows a constant temperature to be maintained on the out- side of the capsule although considerable changes in gamma-ray flux may take place. 72 UNCLASSIFIED PHOTO 44814 . Mockup Device Used in Designing and Testing the Thermocouple Installation Used in the Apparatus The compatibility of yttrium hydride and beryllium oxide with various probable encapsulating materials was also investigated. The materials tested were Inconel, nickel, and types 446, 430, and 410 stai less steel. An Inconel capsule containing yttrium hydride that was heated for 200 hr at 900°C showed no attack. However, when a second capsule was tested at 1000°C, the Inconel reacted with the hydride, and o large opening was formed in the capsule in a matter of 10 to 15 hr. The stainless steels mentioned were all satisfactory and the final selection was made on the basis of most favorable machining, welding, and heat properties. At the present time the apparatus for the experiment and a reactivity mockup required for the ETR critical facility are being assembled. CREEP AND STRESS RUPTURE TESTS UNDER IRRADIATION J. C. Wilson W. E. Brundage W. W. Davis N. E. Hinkle J. C. Zukas The first MTR tube burst tests ' were carried out in a helium atmosphere. The short times to rupture were attributed to thermocouple errors arising from contamination of the helium atmosphere by thermal insulation. A second series of tests have now been completed in an air atmosphere in the MTR, and the resultant data are being evaluated. Equipment is being constructed for operation of tube burst tests in the lattice of the ORR and for carrying out creep rate experiments on the pool side. RADIATION EFFECTS IN ELECTRONIC COMPONENTS Wide-Range Multipurpose Cryostat 0. E. Schow In conjunction with studies of the effects of radi- ation on semiconductor barriers, a cryostat has been constructed with a temperature range of ~200°C to +350°C. Requirements, other than the temperature range, were long-term reliability and ‘‘fail-safe’’ operation, that is, cooling water failure or electrical failure should shut the cryostat down without harmful effects to any of the system components. The cryostat consists of concentrically nested metal ‘*cans,’’ suspended from the top, as shown in Fig. 2.3.4. The innermost container, the sample chamber, is in two sections. The part which actually holds the sample is a 4-in.-ID copper cylinder having %-in. walls. The bottom of the cylinder is c|ose3 with Y -in. copper plate recessed 2 in. from the end. This recessed bottom is neces- sary to reduce the heat loss through the bottom of the cylinder. The over-all length of the copper section is 12 in., of which 10 in. is available for use. The open end of this section is silver-soldered to a 4-in.-ID stainless steel cylinder, 14 in. long, having 0.019-in. walls. The stainless steel 11, C. Wilson et al., ANP Quat Prog. Rep. Dec. 31, 1957, ORNL-2440, p 207. PERIOD ENDING SEPTEMBER 30, 1958 cylinder is, in turn, attached to a brass plug which bolts to the suspension plate. The entire sample chamber, along with its associated thermocouples and heaters, is completely removable for servicing. Around the sample chamber and separated from it by a ]/2-in. heat-exchange-medium chamber is the liquid- nitrogen cooling jacket. This jacket is filled automatically and provides the heat sink for the sample chamber. Around the cooling jacket is a ]]/2-in.-high vacuum chamber to isolate the system from the ambient temperature. The brass suspension plate is supported by the outermost container, a I/B-in. wall, stainless steel cylinder, 30 in. high. Each chamber, with the exception of the liquid- nitrogen cooling jacket, is connected by means of holes in the suspension plate to a gas-and-vacuum manifold panel, from which it is possible to intro- duce vacuum, any desired gas, or air into each chamber. These operations may be performed independently or simultaneously. The system is monitored by gages with which it is possible to determine pressures in the chambers from 2 atm to 10~% mm Hg. Also connected through this panel is a mechanical roughing pump to reduce the system pressure sufficiently to prevent damage to the diffusion pump. The diffusion pump for the system is a water- cooled 300-liter/sec fractionating pump. Silicone oil is used to prevent pump damage if the system goes to atmospheric pressure while the pump is on. This pump is suspended below the cryostat and connected through the outer cylinder by means of a 4-in. ‘‘O’'-ring-sealed flanged joint. The diffusion-pump cooling coil and the boiler coil are connected to a monitoring and control panel. Loss of water pressure, reduction of water flow, or failure of electrical power will automatically initiate an emergency shutdown. The pump cannot be started again until corrective measures have been taken. Temperature control of the cryostat is effected by means of a Speed-O-Max ‘“‘PAT’’ temperature con- troller. This controller uses, as reference, one of seven thermocouples which have been embedded in the copper section of the sample chamber and held there by copper amalgam. Heat is applied to the sample chamber by means of five heaters wound around the chamber. The total electrical input to this heater system is controlled by the controller, but separate adjustments can be made to determine what portion of the total electrical power goes to each heater. Temperatures in the sample chamber 73 |74 UNCLASSIFIED ORNL-LR-DWG 33939 r LIQUID NITRCGEN -~m—HIGH VACUUM \> STAINLESS STEEL HIGH VACUUM TEST CHAMBER HEAT EXCHANGER yd l L) I —< AIR DC HEATER SAMPLE CHAMBER POWER S N NITROGEN TEMPERATURE MANIFOLD PANEL CONTROLLER = :”” —— HELIUM | [ ‘ THERMOCOUPLE 1: I 1 | MECHANICAL | [ PUMP — | f / 1 | I | E THERMO - | o =~ COUPLE DISCHARGE 1 | SR VACUUM VACUUM [ | R GAGE GAGE B —3 S _] DIFFUSION PUMP [ 3 - COOLING PANEL - L % ION GAGE | : : ,Ljf : R . 2 I L | £ i o | ; § } : I j‘\f/’ DIFFUSION AND MECHANICAL PUMP CONTROL PANEL WATE RJ e SUPPLY COPPER POWER INPUT EMERGENCY DIFFUSION PUMP SHUTDOWRN CONTROL PANEL < COMPRESSED AIR g ——=—+= TO DRAIN ) 7 Fig. 2.3.4. Wide-Range Multipurpose Cryostat. MECHANICAL PUMP LY0dI Y S53490¥d LIITFr0dd NV can be maintained to within ]/4°C with a gradient along the length of the copper section of the sample chamber of ]/4°C Grown-Junction Silicon Diode Irradiation J. C. Pigg C. C. Robinson 0. E. Schow An experiment has been performed to determine the effect of fast-neutron radiation on a grown-junction silicon diode. The facility used for the irradiation was the same as that used for bombardment of a point-contact silicon diode and an alloy-junction silicon diode so that comparisons of the three types of diodes would be possible. The sample studied in this irradiation was 1.6 ¢m long, 5 mm wide, and 5 mm thick. The junction was about 8 mm from the end of the p side. The sample was ground and carefully etched. Nickel was plated to the ends, and lead wires were attached. It was then placed in a glass container which was evacuated, flushed with helium, and filled to a pressure of ]/2 atm with helium. The leads and thermocouple wires were brought out through Kovar seals. The irradiation was performed in hole 51 of the ORNL Graphite Reactor. This hole is a fast-flux facility with a fast flux of about 8.3 x 10! neutrons/cm?.sec when the reactor is at 3500 kw. The flux is proportional to the power level. The facility is water-cooled to maintain a temperature of about 23°C. Because the sample was not in intimate contact with the side of the facility, its temperature during the irradiation rose to 33°C. The irradiation was begun with a reactor power of 100 kw, which was raised to a maximum of 2100 kw. At this power the fast flux was approximately 5 x 101! neutrons/cm2.sec. To further assure bombardment by fast neutrons only, the sample was wrapped in several layers of 4.5-mil cadmium. A bias of 1 v was applied in the forward and reverse directions alternately, and the current was measured. A rapid increase in forward current and a small decrease in reverse current which occurred during the very early part of the irradiation is illustrated in Fig. 2.3.5, that is, up to about 5 x 1013 neutrons/cm?. Although not shown conclusively, experience with previous irradiations indicates that this may possibly be attributed to photo-emf's pro- duced in both the sample and in the lead wire insulation by the gamma-ray flux of the reactor. From the dosage of about 1 x 10" to about 1 %103 neutrons/cm?, the changes in the reverse PERIOD ENDING SEPTEMBER 30, 1958 current may be attributed to the increase in temper- ature. |nasmuch as the temperature levels off after 1x10'3 neutrons/cm?, while the reverse current continues to increase, this change in reverse current is believed to be the result of radiation. The irradiation was terminated at 1.8 x 10'® neutrons/cm? because the ratio of forward to reverse current had reached a magnitude which suggested other studies of the sample. A preliminary comparison of the results of this irradiation with those of previous irradiations? show a change in this sample on the order of a factor of 100 increase in reverse current as compared with an increase of 200% for a comparable irradiation of an alloy-junction diode and an increase of 30% in a point-contact diode. Low-Temperature Indium-Germanium Alloy J. C. Pigg C. C. Robinson 0. E. Schow During some preliminary work on the problem of alloying indium into germanium for the purpose of making electrical barriers, some heretofore un- familiar phenomena were noted. This portion of the work dealt with attempts to alloy at temperatures which were not greater than 280°C. On a phase diagram this temperature represents a concentration of indium of about 5%. Alloying times were on the order of 15 min to 1 hr, with additional 10-min periods to bring the samples to temperature and 10-min periods to cool the samples to room temper- ature. Inasmuch as no continuous time-temperature recordings were taken, it would be difficult to estimate the cooling rate of the sample as it passed through various significant temperatures, that is, the eutectic or the freezing point of the indium. |t is believed, however, that the thermal inertia of the heater would prevent extreme, 50°C/min or faster, temperature changes. The alloying was carried out in a helium atmosphere. After the alloying cycle, the major part of the indium was removed by careful cutting, and the remainder of the sample was etched in nitric acid only. This served to remove the indium without significantly attacking the germanium. Microscopic examination of the surface showed no major amount of penetration of the indium into the germanium; 2J. C. Pigg and C. C. Robinson, Solid State Semiann. Prog. Rep. Aug. 30, 1954, ORNL-1762, p 78. ANP PROJECT PROGRESS REPORT UNCLASSIFIED ORNL-LR-DWG 33940 3 | 34 SAMPLE TEMPERATURE (°C)—| el —— » ] .—i_. a .—_.% —al ® S—— — 32 - /- x10 % om 3 p L] b { . — 30 1 ® 21— S ~ | /.—. —1 28 G a o £ / " S o—0 A o = o W 5 % _ — 26 E:_j S [a o so—o-0.0-* A m ~ - wees A ® R - 24 100 kw / fl Va z 500 1000 1400 2100 kw > kw kw kw Q v - REACTOR POWER LEVEL - o1 /‘ —] 22 oJ WA W /. AT {-vbias (x10 % amp) A [, AT {-vbias (x10™% amp) A ® SAMPLE TEMPERATURE (°C) T‘:“ — 20 oA 0 1x10" 5x 10" 1x10'"° 1.5 x 102 2x10% FAST-NEUTRON DOSE (neutrons/¢m?) Fig. 2.3.5. Irradiation Effects on a Grown=Junction Silicon Diode. indeed, most of the surface appeared to be un- touched. (Microscopic examination of germanium etched with nitric acid alone does show some attack, but the action is slow and does not compare with the action of nitric acid on the indium and is not believed to detract from the phenomena observed.) In the areas where penetration was noted, it was seen to be, for the most part, penetration which conformed to rather precise geometric configurations. Some of these may be seen in Fig. 2.3.6. Geometric crystal growths on the surface of the germanium may also be seen. Other examples are shown in Fig. 2.3.7. The sample shown in Fig. 2.3.7 is not the same as the one shown in Fig. 2.3.6. On still another sample, well-defined triangular *‘screw growth’’ was seen. In one instance, a pyramidal crystal was growing from the bottom of a hexagonally sided pit. The pyramidal and triangular shapes of 76 the growths and pits might be attributed to the crystal structure of indium, but it is suspected that such growths would have been removed by nitric acid. Hexagonal pits and growths might be attributed to the germanium. If so, a pit in germanium conforming to the shape of the germanium crystal would not be expected unless the germanium had been dissolved by the indium and the resulting mixture removed by the nitric acid. This leads to speculation about the action of nitric acid on different mixtures of germa- nium and indium. Triangular pits which seem to be progressing towards a hexagonal shape can also be seen in the photographs. The spotted or ‘‘dusty’’ surface of the germanium is thought to be a result of the previously mentioned action of nitric acid on germanium. Fi Penetration. 35X. It is well known that indium-germanium alloy junctions do not result in an even mixing or diffusion of one metal into the other. Recrystallization of single crystal germanium on the indium side of the barrier has been seen, and the irregularity or “raggedness’’ of the junction is still a problem to those interested in junction preparation. Examination of germanium alloy transistors by the above technique, that is, removal of the indium by nitric acid, shows a number of indentations in the germanium surface, but they are not of a geometric configuration. Other samples show no significant pits, but the surface is smoothly curved. The pits of these commercial units may be due to entrapped gas, and, if so, it could be said that the slight curvature of the surface is a result of more complete melting than in samples alloyed at low temperatures. Since a considerably higher temperature is involved PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED | PHOTO 45703 2.3.6. Photomicrograph of Germanium Surface Showing Crystal Growth (Indicated by Arrows) and Indium in the preparation of the commercial units, it may be that samples alloyed at low temperatures indicate the beginning of the alloying process, that is, penetration takes place at some particular spot on the germanium surface in a disordered region of some sort such as, perhaps, a dislocation or an etch pit. Since etching rates along crystal faults are different than on a normal surface, the prefer- ential attack might be limited to a crystal fault inherent in the crystal or introduced by cutting and not removed by etching. Such a conclusion does not appear to be completely unfeasible, since it is known that ““screw growths”’ occur at crystalline faults. If such is the case, it would lead to specu- lation about alloying indium into germanium which has been ground and lapped but not etched. Further speculation would suggest that recrystallization of indium-germanium alloy junctions begins from these growths. Pl UNCLASSIFIED PHOTO 45704 | Fig. 2.3.7. Dark-Field llluminated Photomicrograph Showing Germanium Surface Crystal Growth. 250X. The germanium used in this work was n-type, and it had a resistivity of approximately 2 ohm/cm, which is not extremely different from that of com- mercial units. Before alloying it was etched in CP-4, washed in distilled water, and kept in ethyl alcohol until used. The indium was lightly etched in nitric acid long enough to produce a shiny surface, washed in distilled water, and stored in ethyl alcohol until used. 78 Examinations of the samples after exposure to air for a period of 1to 2 weeks showed that the pits and growths had undergone some change. They were still on or in the germanium, but the edges had been rounded as if they had been lightly etched. Another sample kept in ethyl alcohol for a period of two weeks showed no changes in the shape or sharpness of the pits and growths. Part 3 ENGINEERING A. P. Fraas A. L. Boch Reactor Projects Division 3.1. COMPONENT DEVELOPMENT AND TESTING H. W. Savage E. Storto Reactor Projects Division IRRADIATION TEST OF OIL-LUBRICATED PUMP ROTARY ELEMENT D. L. Gray The device which consists of the rotary element of an oil-lubricated pump and which is installed in a gamma-irradiation facility in the MTR canal, as previously described, ! has continued to operate with downtime limited to 20 min as a result of a power failure and 1 hr as a result of a low=lube-oil- flow alarm. As of August 19, the lower bearing and seal area had received a total gamma-ray dosage of 8.9 x 107 r, and, by September 30, a total gamma-ray dosage of 1 x 1010 ¢ is expected, A large oil reservoir is used with this type of rotary element and thus only a small portion (about 2%) of the lubricant (Gulfcrest 34) is exposed to radi- ation at any time. The average gamma-ray dosage to the total quantity of oil in the lubricating system is therefore expected to be about 2.2 x 108 ¢ by September 30. Insofar as can be ascertained, the operation of the seal has been comparable to that of a seal op- erated without irradiation for a similar period of time. The leakage of oil from the seal into the catch basin has usually been at a rate of 5to 7 ml/day, although higher rates have been observed for short periods from time to time. Results of lD. L. Gray and W. K. Stair, ANP Quar. Prog. Rep. Sept. 30, 1957, ORNL.-2387, p 32. analyses of samples of oil taken from the test equip- ment at periodic intervals are presented in Table 3.1.1. The viscosities of both the bulk and the leakage oil have increased but not deleteriously so, as indicated by operation of the apparatus, The in- crease in the bromine number indicates that irradi- ation has displaced some hydrogen atoms from the hydrocarbon molecules, also without noticeable effect on operation. The acidity number indicates very little contamination of the system. TEST OF NoK PUMP WITH FULTON-SYLPHON SEAL D. L. Gray A sump-type centrifugal pump designed to operate at a shaft speed of 1200 rpm and to circulate NaK at 1200°F at a rate of 1200 gpm was modified to in- clude a Fulton-Sylphon seal and placed in test operation as described previously.? During more than 7500 hr of operation, the oil leakage past the test seal, which separates the lubricating oil from the pump tank helium cover gas, has been of the order of 10 cm? per week, which is exceptionally small for a seal of this type. The similar upper seal, which separates the lubricating oil from the atmosphere, has been equally satisfactory. 2p. L. Gray, ANP Quar. Prog. Rep. March 31, 1958, ORNL.-2517, p 67. Table 3,1.1. Results of Analyses of Oil Samples from MTR Test Apparatus Bulk Circulating Oil Lower Seal Oil Leakage Sample Sample Sample Control Sample Accumulated Taken Accumulated Sample of 7-31-58 3-13 to 3-27-58 4-28-58 524 to 6-21-58 Yiscosity, SUS 90.2 118.5 108.0 112.6 140.2 at 250°C Bromine No., mg of 0.87 5.6 3.4 4,1 7.3 Br/100 g of oil Acidify, mg of 0 0.013 0 0 0 KOH/g of oil 81 ANP PROJECT PROGRESS REPORT CAVITATION TESTS OF CENTRIFUGAL PUMPS P. G. Smith W. E. Thomas One of two pumps being tested, as described previously,? to study the effect of operation in the cavitation region was stopped because of leakage of NaK from the system drain valve, This sump- type centrifugal pump had circulated NaK in the temperature range of 1200 to 1450°F for 2136 hr at a flow rate of 645 gpm, a pump speed of 2700 rpm, and a pump tank cover gas pressure of 2.5 psig. At the time of termination of the test, the pump had been operating under cavitation conditions for 1096 hr. The pump appeared to be in good con- dition after the test; there was no evidence of cavitation damage, The data accumulated during this period are being analyzed. A similar pump that is being operated with a fused salt as the pumped fluid has logged more than 10,300 hr of continuous operation at 1200°F, in- cluding about 8200 hr under cavitation conditions. The oil leakage from the lower seal that separates the lubricating oil from the pump tank cover gas has averaged 10 cm>/day during the past six months. Seal leakage from the upper seal that separates 3P, G. Smith and W, E. Thomas, ANP Quar. Prog. Rep. March 31, 1958, ORNL-2517, p 69. 82 the lubricating oil from the atmosphere has averaged 5 cm3/day. THERMAL STABILITY TESTS OF METAL SHELLS J. C. Amos R. S. Senn A third thin-shell model was tested under high- temperature thermal cycling conditions similar to those previously imposed on two other thin shell models.?*> This shell had the same geometry as those used previously, but it was fabricated by welding over-size sections and reducing the shell to final dimensions by machining the weldment. Testing of the third shell was continued beyond the scheduled 300 thermal cycles to 602 thermal cycles. At this point the test was terminated be- cause of a test loop failure. There was no indication that the loop failure was in any way associated with the test shell. The shell is currently being examined for dimensional changes, and it will be metal- lurgicaily examined for effects of thermal cycling on the welds, 45, C. Amos and L. H. Devlin, ANP Quar. Prog. Rep. March 31, 1957, ORNL-2274, p 50. 3), C. Amos and R. L. Senn, ANP Quar. Prog. Rep. Sept. 30, 1957, ORNL-2387, p 45. PERIOD ENDING SEPTEMBER 30, 1958 3.2. HEAT TRANSFER STUDIES H. W. Hoffman Reactor Projects Division STUDIES OF THE EFFECT OF THERMAL-STRESS CYCLING ON STRUCTURAL MATERIALS J. J. Keyes A. |. Krakoviak Experimental studies of the effect on Inconel of surface thermal-stress cycling in a fused.salt en- vironment (NaF-ZrF -UF ,, 56-39-5 mole %) were continued with the use of the high-frequency pulse- pump loop.! In order to obtain data for comparison with data obtained in the University of Alabama bulk thermalcycling studies? and to establish the frequency dependence of cycles to crack initiation, tests 9 through 11 were conducted at a pulse fre- quency of 0.1 ¢cps and test 12 at 1.0 cps. Tables 3.2.1, 3.2.2, and 3.2.3 present summaries of the operating conditions and results for these runs. The wall thicknesses of the Inconel specimens used for these tests were selected to give just perceptible temperature fluctuations at the outside surface. Further, as seen from Table 2.3.3, tests 9 through 11 constitute a comparison of the effect of exposure time (or total number of cycles) for nearly equiv- alent stress conditions at a given frequency. ]J. J, Keyes, A. |, Krakoviak, and J. E. Mott, ANP Quar, Prog. Rep., Dec. 31, 1957, ORNL-2440, p 54. 2J. F. Gonee, Jr,, and W. D. Jordan, Thermal Fatigue Tests II, MR7, Bureau of Engineering Research, University of Alabama, June 1957, Table 3,2, 1. Operating Conditions of High~Frequency Thermal-Stress-Cycling Test of Inconel Pipes Test Number Test Variables 9 10 11 12 Average fluid temperature, °F 1407 1412 1411 1415 Maximum temperature difference between 550 569 557 575 hot and cold streams, °F Frequency of temperature oscillations, cps 0.1 0.1 1.0 Total pulsing time, hr 200 75 25 100 Total cycles 72,000 27,000 9,000 360,000 Total fluid flow,rate, gpm 3.6 3.6 3.7 3.8 Table 3,22, Results of Metallographic Examination of Test Pieces from High«Frequency Thermal«Stress«Cycling Tests of Inconel Pipe Test No. Metallographic Results 9 Severe intergranular cracking found throughout test section; depths of cracks increased from 0.128 in. at top of test section to 0,207 in. in center section 10 Intergranular cracking found throughout; depths of cracks increased from 0,012 to 0.023 to 0.103 in. in top, center, and bottom of test section, respectively 11 No evidence of cracking 12 No evidence of cracking 83 ANP PROJECT PROGRESS REPORT Table 3,2,3. Summary of Results of High«Frequency Thermal-Stress | ; 3 eI | o i ! : W ooz T L Fu S U S S 1[ , I [ - e = : LY it = AR (I | =1 el o i ] o de= -P ] 5 0.05 Il Pl 1 P a ° 0.4 cps, MEDIUM-GRAINED INCONEL ROD oA 1| * 0.t cps, MEDIUM-GRAINED INCONEL ROD E —4—4 1.0 cps, MEDIUM-GRAINED INCONEL ROD +— E [ ORNL METALLURGY DIVISION FATIGUE ‘ J 7002 F— DATA, MAXIMUM AND MINIMUM ¢, FOR — BERR g COARSE-GRAINED 1.0 cps INCONEL ROD ' = TESTED AT {400°F AND 1.G cps l ‘ 0.01 | T | 104 2 5 103 2 5 108 N, NUMBER OF CYCLES Fig. 3.2.4. Effect of High-Frequency Thermal Cycling on Inconel in a Fused-Salt Environment. A comparison with mechanical fatigue data in terms of total strain. PERIOD ENDING SEPTEMBER 30, 1958 couples were located at the positions B in this figure. Power was supplied for electrical resistance heating of the mercury at 5 v and 1900 amp, maxi- mum. The 60-cycle frequency of the power supply ensured a uniform current distribution with radius (uniform radial volume heat generation). Axial variation of the heat generation as a result of the temperature dependence of the mercury electrical resistivity was neglected, since the variation about the mean was estimated to be only +2%, Heat losses from the heated section were minimized by guard heating and insulation. The experimental results are presented in Table 3.2.4. The Reynolds modulus covered a range from 29,000 to 165,000, and the volumetric heat gener- ation rate varied from 1.4 x 107 t0 8.3 x 107 Btu/hr-ft3. The radial temperature differences given in the final column of the table are based on 4G. L. Muller, ANP Quar. Prog. Rep. Sept. 30, 1957, ORNL -2387, p 106; ASME preprint No.58-HT-17, Second Annual ASME-AIChE Joint Heat Transfer Conference, August 1958, Table 3.2,4. Experimental Results for Mercury Forced-Circulation Heat Transfer with Internal Heat Generation Velumetric . Radial Heat Removed Axial Heat Reynolds Prandtl Electrical Temperature Run from Test in Temperature Generation Number, Modulus, Power Input, . ) Difference, No. R P Bto/h Section, 90t 7 Rise ate, g 3 NRe NPr 9in (Btu/hr) (Bru/hr) out (°F) two- t (Btu/hr-ft ) ( F) x 107 1 3.48 73,300 0.0200 13,880 13,040 0.939 104.47 3.46 2 1.65 29,000 0.0200 6,700 6,060 0.905 127.05 3.90 3 2.55 45,200 0.0200 9,940 10,280 0.969 125,51 4.90 4 7.53 147,000 0.0192 29,500 27,400 0.929 113.35 4.08 5 5.84 102,900 0.0192 22,670 20,530 0.907 125.41 3.25 6 8.30 164,500 0.0191 32,700 29,900 0.915 113.30 5.87 7 5.10 95,700 0.0200 20,600 18,100 0.878 119.03 4.57 8 2.68 88,500 0.0225 10,870 9,710 0.894 63.71 2,23 9 3.20 58,400 0.0205 13,060 11,860 0.908 122,42 3.33 10 2.25 40,000 0.0204 9,060 8,510 0.939 124, 11 2,83 11 1.44 39,200 0.0224 5,660 5,200 0.919 75.49 1.81 12 6.51 124,000 0.0207 26,450 24,100 0.91 116.46 3.78 87 ANP PROJECT PROGRESS REPORT UNCLASSIFIED ORNL—LLR—DWG 25908R _—MIXING CHAMBER AND //THERMOCOUPLE WELL / T ] uT _‘ aNt : o0 208000 A COOLER// POWER } N POWERSTAT TRANSFORMER / — THIN-PLATE ORIFICE IN v = ol«“NT / Al OO TROL\" )y W o0 Q 0 [ SUMP SURGE Z CRAMBER . N AUXILIARY 7! c0 e SN L N T Ter FI F1 1L LI @0-600 psig TC "QUICK-CONNECT” COUPLINGS —&— @0-600 psig LAB. TEST GAGE 0-150psig £ | =600 psig < TC 0 in. o} g =10 H (5)° I 17 FI 600 psig — 5 4 TUBE NO. 2 Mp WALL PRESSURE, psia — Q.02 — M o 65 [ ] 115 A 165 0.0 o} 0 0.004 0.008 0.0t2 0.0t6 0.020 7 (Ib/sec-ft} Fig. 3.5,8. Variation of Ratio of Yortex Tangential Mach Number and Jet Exit Mach Number with Mass Flow Rate per Unit of Tube Length and Tube Pressure for Tube No. 2. ORNL—LR—DEG 34571 0.8 1 | I 1 07 -flm‘%—* —"— TUBE NC.1, 777,= 0.02 Ib/sec-ft | £,= 585 psia; p,= 85 psia MJ’,-=1,9O 08 &= Tuse NO. 2, M, = 0.0097 Ib/sec-ft — p,= 640 psio;, p, = 65psia M/-32.16 0.5 DASHED CURVES ARE FOR IDEAL "FREE VORTEX" M/A/Jlf. 04 0.3 0.2 o 0.2 0.4 0.6 0.8 1. RADIAL POSITION, r' Fig. 3.5.9. Mach Number Profiles for 2«in, Plastic Vortex Tubes. 107 ANP PROJECT PROGRESS REPORT i ORNL-LR-DWG 32293 The local Mach number, M, at r’, is obtained ' directly by differentiation of the pressure data as applied to a relationship for circular flow of an . ideal gas that neglects the small contributions 0.8 . . ° ° due to a radial velocity term: a [w) L ] a 2 1 r’ dp rd 0.6 y p’ dr’ e where p”is the ratio p/p_ . In phase 1-B of the experiments, the diameter of 0.4 TUBE NO 1 the tube was decreased to 1 in., and the experi- waLL PREE?URE (psia) mental Mach profile shown in Fig. 3.5.12 was o 15 obtained. Also shown is a result from a phase 2 0.2 o 165 7 experiment with the 2-in. metal tube under nearly the same conditions. 0 0 0.004 0.008 0.012 0.016 0.020 m1 (Ib/sec - ft) 16 ORNL-LR-DWG 34572 . . . . +e Fig. 3.5.10. Variation of € in the Expression v oc r | | with Mags Flow Rate per Unit of Tube Length for Tube 4 ] No. ]o ll iR 12 i Y + \\ L \ 10 ORNL-LR-DWG 32292 \ | —o— PLASTIC TUBE No.3, fin. 1D ' \ M= 0028 Ib/sec-ft ‘ 1 1.0 \ PO = 386 psio; £,= 52 psio \ My =196 h\ 0.8 o \\ —e— METAL TUBE, 2in. ID MM, 0.8 X 71, = 0028 Ibfsec-ft o P . \\ A = 745 psie, £, = 68 psic \ M-2m 0.6 ‘ \\ (REPLOT FROM Fig. 3.5.13) E o o 06 A e \ ) N o N " " 04 \ \\,\,.FREE VORTEX 0.4 b NS TUBE NO. 2 N WALL PRESSURE (psic) N (o] 65 e o 15 SN o 165 0.2 ‘\\"‘"---_(_ — o \/h-________" 0 0.004 0008 0.012 0.016 0.020 . 77?1 (Ib/sac - ft) 0 02 0.4 06 0.8 1.0 RADIAL POSITION, + Fig. 3.5.11. Variation of € in the Expressiony o r € with Mass Flow Rate per Unit of Tube Length for Tube No. 2, Fig. 3.5.12. Comparison of Mach Number Profiles for 1=in.=ID and 2-in,-ID Vortex Tubes with no Bleed Flow. 108 The objective of the phase 2 experiments was to investigate the effect of two methods of bleeding off a fraction of the inlet flow at some radial position away from the tube center. In the first method, the bleed was from the tube periphery through a uniformly porous wall, in the hope that sufficient boundary layer stabilization could be achieved to significantly lower the eddy losses induced by shear forces at the wall and thus in- crease the vortex strength. The Mach profile obtained with a uniform wall bieed ratio of 2.8 is compared in Fig. 3.5.13 with that obtained with no bleed at the same inlet mass flow. Results are presented in Fig. 3.5.14 for an alternate method of bleed in which the excess flow is removed radially at a position 7 = 0.25 so that the exit ORNL - LR-DWG 34573 0.9 0.8 4 a =0 M, = M, = 0.028 Ib/sec- ft 0.7 | & | \ 1 l \ 06 il | |I |‘\ :—:2: m] = 0.028 |b/SeC « ft ALL FLOW RADIAL AT TUBE CENTER By = 715 psia; A, = 68 psia M = 248 ‘\ Mg = 0.0074 lo/sec - ft ) 0.5 \‘l UNIFORM WALL BLEED RATIO = 2.8 ] M/M, \ - /M, \ M, = 1.98 ‘\ 04 \ \ ) DASHED CURVES ARE FOR IDEAL \\ "FREE VORTEX" 0.3 0.2 / O % ‘M ' —&< ; 0 0 02 0.4 0.6 0.8 10 RADIAL POSITION, r/ Fig. 3.5.13. Mach Number Profiles for o 2-in.-ID Metol Vortex Tube Showing Effect of Uniform Wall Bleed. PERIOD ENDING SEPTEMBER 30, 1958 ORNL-LR-DWG 34574 0.5 0.4 | /% 7= 0029 Ib/sec ft M= 0.0062 Ib/secft Po=T40 psic;pw= 61 psia M;=2.27 \ 0.3 \ MM, \ c.z2 \ . o \_____ .AP.\ 0 0.2 04 0.6 0.8 1.0 RADIAL POSITION, r' Fig. 3.5.14. Mach Number Profile for o 2«in.«ID Metol Vortex Tube with Radial Bleed 0,25 in. from Center, flow at r” = 0 is held within allowable limits. The main effect here is that of increasing the inlet mass flow parameter m /27y over a large fraction of the radius. The foilowing general observations from the data are noteworthy: 1. The vortex strength, M_/M ., for a 2-in.-diam- eter increases from about 0.63 for plastic tube No. 2 with m, = 0.0045 Ib/sec-ft (Fig. 3.5.8) to about 0.08 for the metal tube with m, = 0.028 Ib/sec-ft (Fig. 3.5.13). Thus, the vortices which were generated at low mass flows appear to be significantly weaker than would be required in vortex reactor applications. The plots of Fig. 3.5.9 for the 2-in. plastic tubes show that the maximum value of M/M. which can be achieved at low mass flows is severely limited. The small peak in the profiles which occurs near the wall for all the 2-in. tubes is not understood but may be related to the nature of intfroduction of the gaseous |et. 2. As indicated in Figs. 3.5.10 and 3.5.11, the value of the exponent € in the relationship v« r*€, where v is the tangential velocity, in- creases from —0.4 at m, = 0.0045 Ib/sec ft to about -0.8 at m, = 0.011 Ib/sec-ft; a further in- crease in m, appears fo produce little change in €, Since deviation of € from —1 is a measure of the deviation from the ideal ‘‘free vortex,’’ it is 109 ANP PROJECT PROGRESS REPORT concludgd that although the vortex strength is low the velocity distributions roughly approximate free vortex behavior at the higher mass flows. 3. The effect of decreasing the tube diameter from 2 to 1 in., as illustrated in Fig, 3.5.12, is to increase significantly the velocity at a given radial position, 7”. [n the 1-in. tube a value of M/M . of 1.25 was achieved (M = 2.45) at "= 0.085; whilé at the corresponding absolute radius in the 2-in, tube at 7 = 0.043 in., M/M.= 0.6 (M =1.31). At a radius of 0.25 in., the 1-in. tube produced M/M. = 0.25, while for the 2-in. tube M/M.=0.27. ], 1. Thus it appears that by decreasing the radius the performance of a given vortex tube is not impaired, and the advantage is gained of enabling more tubes to be included in a given volume of reactor, 4. When a uniform wall bleed is applied at a given inlet mass flow, as illustrated in Fig. 3.5.13, the effect is to increase slightly the velocity near the periphery of the tube at the expense of that nearer the center. The rapid rise at r* <0.06 for the profile with bleed is a conse- quence of the small exit hole employed in this run as contrasted with the larger annular opening in the run with no wall bleed (see Table 3.5.1). This effect is discussed in a forthcoming report.3 It is felt that the failure of wall bleeding to achieve a significant improvement may be due to the high degree of turbulence induced by the jets in the vicinity of the wall so that a laminar boundary layer cannot be achieved with practical bleed ratios. 5. As an alternate method of bleed-off, a run was made in which the excess mass flow was removed at a radial position 0.25 in. from the center of the 2-in. metal tube; no wall bleed was employed. The results are depicted in Fig. 3.5.14, which indicates when compared with Fig. 3.5.13, that for 7 > 0.1 the velocity profile agrees well with that obtained when all the flow is removed at the tube center. This technique shows promise for increasing velocities and vortex strength at the expense of lowering exit mass flow rates. In addition to the basic experiments discussed here, numerous runs have been carried out to 3). L. Kerrebrock and J. J. Keyes, Jr., An Experi- mental Study of Vortex Tubes for Gas Phase Fission Heating — Part I, ORNL CF-58-7-5 (July 5, 1958). 110 study the effects of number, size, and location of the inlet nozzles, tube length, methods of bleed- off, gas properties, and pressures. Some pre- liminary separation experiments have also been made using He-Br, and He-C Flé gas pairs, ds discussed in ref 3. Since these results are not conclusive, it is intended to continue the separa- tion studies. SURVEY OF DESIGN PROBLEMS OF AUXILIARY POWER UNITS FOR SATELLITES A. P. Fraas Reactor Projects Division The performance of various types of auxiliary power unit that might be used in satellites has been reviewed during the past several months. Informgtion from the SNAP (Systems for Nuclear Auxiliary Power) program indicates that the weight of either solar cells or thermoelectric (or thermo- ionic) power units employing radicactive isotopes as a heat source for a reconnaissance satellite will be about 500 to 1000 Ib/kw. While lighter weights (350 Ib/kw) might be achieved with solar cells continuously oriented toward the sun, satellite missions which involve spending almost half of the time in the earth’s shadow immediately reduce the efficiency of such solar cells by roughly 50%. Further, it is difficult to effect the orientation of the solar cells toward the sun when mission considerations entail orientation of other elements of the satellite toward the earth. In addition, the specific weight of either the solar cells or the radioisotope power sources is not reduced as the power output of the unit is in- creased, and the voltage obtainable from such units is low. From preliminary estimates it appears that the use of a reactor as a heat source might make possible auxiliary power units of only 20 to 40 |b/kw for power outputs in the range of 20 to 200 kw net electrical output. Cycle Performance Considerations One SNAP || power plant proposal suggests three fluid circuits: a sodium circuit cools the reactor, mercury serves as the thermodynamic cycle working fiuid, and terphenyl is used to transfer heat from the condenser to the radiator. This approach entails three heat exchangers, three pumps, three expansion tanks, and three sets of instruments and controls. A very large improwament in reliability and a substantial saving in weight should be possible through the use of a simpler system employing a single fluid. In the mercury-vapor cycle used in both the SNAP | (radioisotope source) and SNAP || (reactor) power plants, there will probably be a substantial holdup of droplets of mercury on the condenser surfaces, and therefore the weight of the mercury in the condenser will probably be greater than the dry weight of the condenser. It would also be desirable to increase the operating temperature level in the condenser to save con- denser weight, and this is not practicable be- cause it would imply too high a pressure level for the high-temperature portion of the mercury- vapor cycle. Further, the low specific volume of the mercury vapor leads to so small a turbine that the shaft speeds tend to be high, and the aerodynamic efficiency of the turbine is low. Cycles employing either aluminum chloride or rubidium vapor should have several important advantages relative to mercury for a light-weight power plant, In both cases, a single fluid can be used to cool the reactor and carry out the thermodynamic cycle. Flow diagrams for power plants employing these cycles are presented in Figs. 3.5.15 and 3.5.16. The thermodynamic properties of these fluids have been estimated and reports are being prepared to present the UNCLASSIFIED ORNL—LR—DWG 34575 EVAPORATING - DRUM / TURBINE I X Y GENERATOR REACTOR ! COMNDENSER ‘ FEED PUMP CIRCULATING PUMP Fig. 3.5.15. Flow Diagram of Nuclear Power Plant Employing o Rubidium-Vapor Cycle. Wimany with caption) PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED ORNL—LR—DWG 34576 GENERATOR S // TURBINE COOLER REACTOR | COMPRESSOR Fig. 3,5.16. Flow Diagrom of Nuclear Power Plont Employing an Aluminum Chloride Gos Cycle, (MDY with caption) resulting data. Some preliminary calculations for some typical cycles have been made and are in- cluded below. The rubidium vapor cycle would operate much like a boiling-water reactor with a pump circu- lating a weight flow through the reactor of roughly ten times the weight of rubidium vapor- ized per pass. Of course the vapor bubbles formed in the reactor will pose serious reactivity and control problems, but it seems likely that the reactor can be designed to reduce these effects to tolerable levels. A gas-turbine cycle utilizing aluminum chloride has some unusual characteristics. Aluminum chloride dissociates from Al,Cl, to AICI, in the temperature range between 700 and 1600°F, with the bulk of the dissociation taking place in a 400°F range that increases with pressure. Pre- liminary calculations indicate that the cycle can be designed so that the gas will be mostly in the form of Al ,Cl, during compression, while during the expansion process it will be mostly AICIl,. This, in effect, will cut the compression work in half and thus produce a marked improve- ment in cycle efficiency. The nature of this effect can be visualized by examining the P-V diagrams of Fig. 3.5.17, which compare similar ideal gas turbine cycles for helium and aluminum chloride. The work involved in each compression 11 ANP PROJECT PROGRESS REPORT RANKINE CYCLE (RUBIDIUM) BRAYTON CYCLE (Al,Clg—AICIg) ORNL— LR-DWG 34577 BRAYTON CYCLE (HELIUM) ° ] C] \ g 2 g N\ w @ & \ 3 2 @ § IDEAL WORK (NPUT = 0.45 Btu/Ib ? IDEAL WORK INPUT =25.1 Btu/Ib & IDEAL WORK INPUT = 477 Biu/Ib o o 1l a a a Y | ! | | [ I i T T [ y IDEAL WORK OUTPUT = 144 Blu/Ib IDEAL WORK OUTPUT = 627 Btu/Ib \ 3 3 2 g \ 2|\ g N N 0 \ " 2 \ 2 ? & \ & % IDEAL WORK OUTPUT = 604 Btu/Ib 5 N & \\ & \_{-—_.__ \1 P [ [ ! [ | I ! [ I | IDEAL NET WORK =113.8 Btu/Ib IDEAL NET WORK = 37.6 Btu/Ib \ 3 3 2 oA e : O & i \ W \ > \ D =) [72] w (72} & \ & \ \ @ | IDEAL NET WORK =127 Btu/lb \_ \\‘-—_____ VOLUME (ft3/1b) VOLUME (ft3/1b) VOLUME (#t3/1b) Fig. 3.5.17. P-V Diagrams for Typical ldeal Thermodynamic Cycles (Data from Tables 3.5.2, 3.5.3, and 3.5.4). or expansion process is directly proportional to the area of the P-V diagram, and the net work is proportional to the net area for the cycle. A Rankine cycle utilizing rubidium vapor is also included in Fig. 3.5.17 to show that the proposed aluminum chloride cycle is roughly midway be- tween a gas turbine (or Brayton) cycle and a Rankine cycle in its requirements for work input during the compression process. The diagrams of Fig. 3.5.17 were prepared for ideal cycles with no ailowances for losses. The most important losses are associated with the efficiencies of the compressor and the turbine, which are likely to be of the order of 80%. This means that with an 80% efficient compressor the ideal work input will be 80% of the actual work input, while the actual work output of the turbine will be only 80% of the ideal. In addition, pres- sure drops between the compressor and the 112 turbine will cause losses in net output from the cycle. The nature of these effects can be seen readily in Fig. 3.5.18. If allowances are made for these losses to obtain the actual net outputs for the cycles of Fig. 3.5.17, the diagrams of Fig. 3.5.19 result, The relatively large work input required for the compression process of the Brayton cycle makes it very sensitive to compressor inlet temperature, because the compression work increases rapidly with the initial temperature. As a result, the net work output and over-all thermal efficiency of the Brayton cycle drop off so rapidly with in- creasing temperature at the compressor inlet that for most plants the compressor inlet temperature must be held below 200°F. At the same time, the turbine inlet temperature must be at least 1200°F, and preferably should be above 1400°F. For the helium cych_gifig. 3.5.19, even with a 1540°F - a4 UNCLASSIFIED ORNL-LR-DWG 34727 ! I I T I 80 PRESSURE LOSS THROUGH HEATER TURBINE INLET TEMPERATURE = 1500°F 60 ! t r—NET EITFECTIVE AREA INE 40 20 [EDDY LOSSESIN A COMPRESSOR — ‘ Tl-llROUGH CIIOOLER I l I 'PRESSURE LOSS THROUGH HEATER | 1 TURBINE INLET TEMPERATURE = w (@] PRESSURE {psi) (o] | 1200°F 60 | i NET EFFECTIVE AREA 40 EDDY LOSSES IN TURBINE EDOY LOSSES IN 20 I~ COMPRESSOR “1 PRESSURE LLOSS THROUGH COOLER 0 a4 8 {2 16 20 24 28 SPECIFIC VOLUME (ft¥1b) Fig. 3.5.18. P-V Diagrams for ldeal Gas-Turbine Air Cycles with Cross-Hatched Areas to Indicate the Magni- tude of the Principal Losses. turbine inlet temperature, the net work from the cycle is negative when the compressor inlet tem- perature is raised to the level required to give a reasonable radiator specific weight. The aluminum chloride possesses another important advantage over conventional gases for use in a gas turbine; the heat transfer coeffi- cient during the heating and cooling processes should be exceptionally high. The large amounts of energy involved in dissociation appear to give effects analogous to those responsible for the high heat transfer coefficients characteristic of PERIOD ENDING SEPTEMBER 30, 1958 boiling and condensing heat transfer conditions, This should reduce the heat fransfer area re- quired in comparison with the area required for gases such as helium or nitrogen. A further advantage of aluminum chloride is that it could be used in the vapor phase throughout the cycle so that there would be no free liquid surface at any point in the system, and hence no expansion tank would be required. While it is possible that a free liquid surface could be stabilized by surface tension forces or other factors, it is also true that a great deal of difficulty has been experienced at ORNL with stabilization of free liquid surfaces in pumped systems, To date no one has yet sug- gested a means of carrying out laboratory tests which would approximate free-flight conditions. Thus there is a major advantage to the use of a working fluid which would avoid any difficulties with liquid surface stabilization. Summaries of the calculations on which Figs. 3.5.17 and 3.5.19 were based are presented in Tables 3.5.2, 3.5.3, and 3.5.4. Planned Program The next phase of this work will entail a com- prehensive series of analytical studies to de- termine the performance obtainable from rubidium and aluminum chloride as working fluids as com- pared with more conventional fluids such as mercury, sodium, helium, air, and water, The studies will also disclose the major problem areas most deserving of attention. If this phase of the work indicates a substantial superiority for either rubidium or aluminum chloride as a working fluid in an auxiliary power plant unit it will be fol- lowed by experimental determination of heat transfer characteristics of the working fluid chosen, investigation of corrosion problems associated with the fluid, and more complete de- sign studies, 113 ANP PROJECT PROGRESS REPORT RANKINE CYCLE (RUBIDIUM) BRAYTON CYCLE {Al»Clg-AICl3) e rnd ORNL-LR -DWG 34578 BRAYTON CYCLE (HELIUM) ACTUAL WORK INPUT = 0.3 Btu/1b ACTUAL WORK INPUT = 31.2 Btu/Ib % @ @ ] ] a w w w @ 1 [1ef =} 2 pn } @ & @ ‘ | w v w w w W 14 @ [1 4 a a %@h Q@ — ACTUAL WORK INPUT =597 Btu/Ib —— R— [ | 1 T | | | I l ACTUAL WORK QUTPUT =79.8 Btu/Ib ACTUAL WORK OUTPUT=50.2 8tu/Ib 2 2 z w w E 5 5 z w ()] w 2 Q @ ] ‘ w l&J w T a : ,‘% &l ACTUALWORK QUTPUT= 493 8tu/ib —| 1 H l % ACTUAL NET WORK=79.5 Btu/Ib B ACTUAL NET WORK=19Btu/Ib ACTUAL NET WORK=-114 Btu/Ib — -—_ | — ‘@ ‘@ ‘@ s & 2 & % & & o 2 =} w w w (73] w w & %\ 2 L Q. Q o %%h—. VOLUME {ft/1b) VOLUME (£t ¥1b) VOLUME (ft3/ib) Fig. 3.5.19. P-V Diagrams for Typical Ideal Thermodynamic Cycles with CrosssHatched Areas to Represent the Losses Entailed by Compressor and Turbine Efficiencies of 80%. Table 3.5.2. Rubidium Vapor Cycl -] Temperature Pressure, Specificm Enthalpy, b Quality Change in Condition (°R) b (psia) Volgme, v (Bty/1b) (%) Enthalpy, Ab (ft~/1b) (Btu/Ib) Yapor to turbine 2060 66.33 3.892 498 100 417 Yapor to condenser 1460 2.66 384 114 (isentropic) Yapor to condenser 1460 2,66 60.5 418 88.1 79.8 (70% etficiency) Liquid to feed pump 1450 2.46 0.01275 81 0 79.8 Cycle efficiency = —= 19.1% 417 114 PERIOD ENDING SEPTEMBER 30, 1958 Table 3.5.3. Aluminum Chloride Cycle (isentropic) Cond Temperature Enthalpy, Enthalpy, Pressure, VSpIeCIflc- DProportior; Eha:gle " ondition o o \ olume, v issociated, nthalpy, ("R) h (Btu/Ib) S (Btu/"F) (psia) (Bro P (£13/1b) x Ab (Btu/Ib) Compressor inlet 1200 203.48 0.0828 5 10.346 0.0726 Compressor outlet 1356 228.6 0.0828 60 0.967 25.1 (isentropic) Compressor outlet 1378 234.9 0.0873 60 0.992 0.0889 31.2 (80% efficiency) Turbine inlet 2000 490.2 0.2388 60 2.516 0.8782 255.3 Turbine outlet 1653 427.5 0.2388 5 24.46 0.8410 62.7 (isentropic) Turbine outlet 1685 440.0 0.2458 5 25.40 0.8755 50,2 (80% efficiency) 37.6 ldeal Cycle Efficiency = —— = 14.37% 261.6 19 Actual Cycle Efficiency = —— = 7.43% 255.3 Table 3.5.4. Helium Cycle Temperature Pressure, Specific Condition (°R) b (psia) Yolume, v (#13/1b) Compressor inlet 1200 100 7.2 Compressor outlet 1582 200 4,75 (isentropic) Turbine inlet 2000 200 6 Turbine outlet 1517 100 9.1 483 — 382 101 ldeal Cycle Efficiency = ———— = — = 24.2% 2000 - 1582 418 386 - 478 =92 Actual Cycle Efficiency = = 418 - 96 322 115 Part 4 SHIELDING E. P. Blizard Neutron Physics Division 4.1, SHIELDING THEORY ANALYSIS OF NEUTRON PHYSICS DIVISION EXPERIMENTAL STUDY OF GAMMA RAYS PRODUCED BY NEUTRON INTERACTIONS IN AIR F. L. Keller 0. S. Merrill! The gamma-ray dose rate which results from secondary gamma rays produced by neutron inter- actions with the air can be neglected in most shield design problems because it is small com- pared with the dose rate which results from air- scattered source gamma rays. Situations may arise, however, in which it is desirable to allow a large amount of neutron leakage from the reactor shield and place part of the neutron shielding material around individual regions which must be further shielded. In many of these situations the air-scattered gamma rays, which are usually of rather low energy (that is, L = INELAST!IC-SCATTERING GAMMA RAYS / . o g -. PR JE— o c 0.0? D . T e ° l : i | o ® 0.0 .. 'Q. _ . ] - 8 i s b e ] 0.005 |—+ P . S | | _ o ] 0.002 -t : ! * * | 0.004 - e SO S SRS s s SR — T , S S ——— RS 0.0005 S - ; ; AN S - f T— ‘ . i s . | ] 0.0002 - | | 0.0004 L 0] 100 200 300 400 500 600 700 800 200 1000 1100 1200 PULSE HEIGHT Fig. 4,1.1, Pulse-Height Spectra of Capture and Inelastic Scattering Gamma-Rays Resulting from Neutron Inter- actions in Air. 120 spectra of gamma rays incident on the outside of the collimator. Therefore, it was considered ad- visable to make the comparison by first calculating the spectra incident on the outside of the water collimator from both capture and inelastic scat- tering gamma rays, then determining the attenuation and buildup associated with the passage of the radiation through the water collimator, and, finally, taking into account detector response functions to obtain calculated pulse-height distributions which could be compared directly with the experimental results presented in Fig. 4.1.1. Spectra of Gamma Rays Incident on the Outside of the Collimator A rather accurate estimate of the complete spec- trum of gamma rays from neutron captures in N4 was obtained by making use of the direct experi- mental measurements of Bartholomew and Campion, The report by Lustig, Goldstein, and Kalos on neutron cross sections of nitrogen3 supplied the needed theoretical values of the total inelastic scattering cross sections for neutrons with energies up to 18 Mev and also the results of Hauser- Feshbach calculations of the cross sections for exciting the first four individual nuclear levels with various energy neutrons up to 6 Mev, Since these individual level cross sections are relatively flat in the region of 6 Mev, they could be extrapo- lated to somewhat higher neutron energies. Then, for neutron energies above 6 Mev, the difference between the total inelastic cross section and the sum of the individual level cross sections repre- sents the cross section for exciting some level whose energy is greater than 6 Mev. Branching ratios obtained from a review article by Ajzenberg and Lauritsen® were used with the individual level cross-section curves to obtain cross-section curves 2 for producing gamma rays with discrete energies less than 6 Mev from these levels. The branching ratios also indicated that most of the levels in the region from 6 to 9 Mev decay directly to the ground state, Since these levels are rather closely spaced and since they decay predominantly to the ground state, it was assumed in calculating the gamma-ray spectrum that the difference cross section men- tioned above could be treated as a cross section for producing an inelastic scattering gamma ray 6F. Ajzenberg and T. Lauritsen, Revs., Modem Phys. 27, 77 (1955). PERIOD ENDING SEPTEMBER 30, 1958 with an energy approximately equal to the initial energy of the neutron. In the case of oxygen the first, and most important, level is at 6.09 Mev, and it was assumed that the entire inelastic oxygen cross section could be taken as the cross section for producing a 6.09-Mev gamma ray. With these cross sections and an assumed neu- tron distribution it was then possible to estimate the spectrum of the inelastic scattering gamma rays produced. The spectrum of fast neutrons above 3 to 4 Mev (which are the only neutrons of importance for inelastic scattering in air) was assumed to have a fission shape in these calculations. The spec- trum of inelastic scattering gamma rays obtained in this manner is shown in Fig. 4.1.2, where the discrete energy gamma rays have been represented ( 10—26) UNCLASSIFIED X 10 ORNL-LR-DWG 31793 5 2 +—nH s a s S S SN l O e S o [ 1] U S~ Y -7 e [ ] § / i I 1 - 1 3 ‘4 ’ 2 3 | H 1 iy ! S | o 4.4 /- I 1 04 ’ 7z x ZAZ A 7 s | duagd AN iKY [ \ D|0'2 .7 M?L’w_.;_fifi_.fié_fi’ AN k 7750 . — - ANNAN ANSANN:Y - 7E 2R - — - NN AN - _._‘fifi_ _/__?‘__fifi%_j ; \\\\\Q NN \\ --EZE700/ 7 . RN DRRNAY IZEN707007007 7 202 I SN\ ANWNY S THHUIH B 0 AR N I+ HH41—A />_ - 7 1 1 ‘N Ty o qadidg 7 a0 N\ N WU EBRZEEINN N AT EINN 2 I AN % g NN ‘U’ BRZ R HUNRBRZRY AR B A GH A il ’ YA HH 1 2 3 4 5 6 7 8 9 10 #H 12 13 £y, GAMMA RAY ENERGY (Mev] Fig. 4.1.2, Calculated Inelasfi'ic Scattering Gamma- Ray Number Flux Spectrum at the Detector (No Water in Collimator), 121 ANP PROJECT PROGRESS REPORT in histogram form with an energy width of 0.2 Mev. The fact that the calculated distribution above 6 Mev is continuous rather than discrete, is, of course, due to the lack of detailed knowledge con- cerning the excitation of individual levels in this region. Fortunately the number of gamma rays in the continuous portion of the curve is small com- pared with the number of gamma rays in the dis- crete lines, and hence the approximations used in the calculations concerning these gamma rays are probably sufficiently accurate. Spectrum After Attenuation and Buildup in the Woter Collimator Since all gamma rays which reached the detector without making a collision in the water collimator travelled through approximately the same water thickness (™4 ft), the uncollided spectra at the detector were determined by simply multiplying the incident spectra by e ““(E)T gt each energy point, where u(E) is the total linear absorption coefficient of water at energy E and T is the water thickness (~4 ft), Next, it was necessary to determine the distribution of scattered, or buildup, gamma rays at the detector for each case. Since, as mentioned above, all the uncollided gamma rays which reached the detector penetrated essentially the same water thickness, it was assumed that the distribution at the outside of the water coili- mator could be artificially represented by a single point isotropic source whose strength was nor- malized to give the same uncollided flux at the detector as that calculated above. In the case of gamma rays it appears that even for point isotropic sources in infinite media most of the scattered gamma rays which reach a detector are gamma rays which leave the source in almost the right direction to reach the detector without scattering. Goldstein and Wilkins? have calcu- lated the spectra of buildup gamma rays at a de- tector for cases of point isotropic sources in infinite media, and, since the assumed artificial point source was normalized so that the correct number of gamma rays left the source in the direc- tion toward the detector (although an incorrect number left at large angles to the source-detector 7H, Goldstein and J. E. Wilkins, Jr., Calculations of the Penetrations of Gamma Rays. Final Report, NYO- 3075 (June 30, 1954). 122 axis), it was further assumed that the Goldstein- Wilkins results could be applied directly to the artificial point source to determine the buildup spectrum at the detector in the actual problems. Plots of the total unnormalized gamma-ray number flux distributions at the detector (that is, collided, or buildup, plus uncollided) for capture and inelastic scattering gamma rays, which were obtained in the manner described above, are shown in Figs. 4.1.3 and 4.1.4, respectively. In these figures the histograms representing the uncollided discrete-energy gamma rays are superimposed on continuous distributions which are almost entirely due to scattered, or buildup, gamma rays. UNCLASSIFIED -2 ORNL-LR-DWG 34795 10 3 S B R },_ O w b s =) a N \ - ¥ - £ 9 w - i‘_) 2 : | = - - g e pa -3 S 10 @ o} ~ [aN) £ [#] ~ 4 5 = < [r - L= ot = = < U) x 2 > ’— a S L L _4 N O 10 o W o = e S ) =4 O 5 w ~N 3 < = or s 1 2 =z 2 E RS 40—5 1 2 3 4q 5 6 T 8 9 10 M 12 £, GAMMA RAY ENERGY (Mev) Fig. 4.1.3. Calculated Total Capture Gamma-Ray Number Flux Spectrum at the Detector (Water in Collimator). UNCLASSIFIED -26 (X140 ) ORNL-LR-DWG 31796 B | S y & M — GAMMA RAYS /cm?/sec/ Mev WHICH STRIKE A DETECTOR) 3 4 Tie (E'MUN-NORMALIZED NUMBER OF INELASTIC SCATTERING 1+ 2 3 4 5 6 7 8 9 10 M 42 57’,, GAMMA RAY ENERGY {Mev) Fig, 4.1.4, Calculated Total Inelastic Scattering Gamma-Ray Number Flux Spectrum at the Detector (Water in Collimator), Calculation of Pulse-Height Distributions from the Detector Finally, it was necessary to include the detector response so as to obtain pulse-height distribution curves whose shapes could then be compared directly with the shapes of the curves obtained from the experimental results, Experimental measure- ments of response functions of a 3- by 3-in. sodium iodide crystal are available for various mono- energetic gamma-ray sources, These response functions were replotted in histogram form, where the averaging was performed over 1-Mev intervals, and interpolations were then made to obtain esti- mates of the response functions at 0,2-Mev intervals over the entire energy range of interest. These PERIOD ENDING SEPTEMBER 30, 1958 response functions were then applied to the calcu- lated number flux distributions at the detector to give calculated pulse-height distributions from the detector for both capture and inelastic scattering gamma rays. A comparison of the shapes of the experimental and calculated pulse-height distri- butions from both capture and inelastic scattering gamma rays is shown in Fig. 4.1.5, The shapes of the measured and calculated curves are in excellent agreement, Since the spectrum of capture gamma rays which was taken to be incident on the outside of the water collimator was known to be good because most of it was ob- tained from the results of direct experimental measurements, the excellent agreement between the shapes of the calculated and measured pulse-height distributions from capture gamma rays indicated that the methods used to determine the effects of water buildup and detector response could be con- sidered to be reliable. As a corollary, the ex- cellent agreement which was also obtained between the shapes of the calculated and measured pulse- height distributions from inelastic scattering gamma rays gives confidence in calculations based on a spectrum of inelastic scattering gamma rays ob- tained in the manner outlined, This analysis has been published in more detail in a topical report.® MONTE CARLO CALCULATION OF THE DEPOSITION OF GAMMA-RAY HEATING IN STRATIFIED LEAD AND WATER SLABS L. A. Bowman® D. K. Trubey It was reported 10 previously that an Oracle Monte Carlo calculation of the penetration of monoener- getic, monodirectional gamma rays in a lead and water shield had been made that included a total of 512 problems. The problems resulted from all combinations of eight different lead and water con- figurations (see Fig. 4.1.6), four total slab thick- nesses (1, 2, 4, and 6 mfp), four energies for the incident gamma rays (1, 3, 6, and 10 Mev), and four angles of incidence (0, 60, 70.5, and 75,5 deg). The results obtained included the dose rate and energy flux throughout the slab and at the rear of the 8F. L. Keller and O. S. Merrill, Analysis of the Recent TSF Secondary Gamma Ray Experzmenz ORNL- 2586 (Avug. 25, 1958). %0n assignment from Wright Air Development Center, 195, Auslender and A. T. Futterer, ANP Quar, Prog. Rep. March 31, 1957, ORNL-2274, p 286. 123 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2—01—056—-22—596R{ 20— | r ] L9 ‘ ‘ L ' ’ ‘ L i : \ i Lo ¢ | N SN 100} 1+ — : P [ W H - ! — — - i i I ¢ . e _ _ i : e O ; ol — ‘ T s S e e B S R e S NS EEEE . — - B’ J ( — J ’ RO ' : R l ., i ‘ ‘ i — | RS R —_— ] " 1 | 1 i | . - .. 0% o } ! : ; ZITe e, T 7 | =T DIFFERENCE CURVE: . %, . | i -7 NO BORON PLEXIGLAS ON REACTOR . o°B o & . afe | | COLLIMATOR MINUS BORON PLEXIGLAS ‘ | A L ON REACTOR COLLIMATOR | e I e R I L ' o —— I l 1 SR ! i L ] | H R A~ o SR S SN TV SR | | — - ' . : ‘ ; ‘ i o i e A?..A . ‘OO,Q,_QARi,,, , | ‘ (Y : ooo} i i —_— e e S Lo - : ..'. ; A &bo 02}~ v | o e ——— ! Co 4 l | ! %, i A—f'\f ‘% ! _ | L e, e o 1 & ol s, [PcAlLcULATED PONTS © 5 Yt S S S ‘,/ (NORMALIZED) L T % - f’ - — ; FAYNNES S S e - ‘ R = L D , B A I A o o i L ! : (o] & i ST e ; ' o | Noos i | oy 1 W T = i ! ! P c | H . + i e L - | [ . 3 INELASTIC-SCATTERING GAMMA RAYS | ® o " ‘ © L -y N ; R R e oo 77777(..7777"71 PR - P e S T ‘ | i 3 ! o : '° | oo2 + -+ -t ! : : : B e A S : I — - s | C ‘ ’ 1 | . | o I o . ! _ .01 || o ‘ S ~ : S e R R 0.005 j:_ ; S — . - 4 . . . _1.__ .. . . R o i - . - i - 1 0.002 | 4 | et - S | | | . : [ ] 0.004 - S e . : . N e b o RN R : 0.0005 e B ; : - e b S - _ 4 i S __5 _— ? — e ! i | 0.0002 +—- e ‘ % ‘ - e et ‘ : ] ‘ | e 0.0004 —— i 1 i ‘ R [ : 0 100 200 300 400 500 600 700 8OO 900 4000 1100 1200 PULSE HEIGHT Fig. 4.1.5, Comparison of Experimental and Calculated Pulse-Height Spectra of Capture ond Inelastic Scottering Gamma Rays Resulting from Neutron Interactions in Air, 124 PERIOD ENDING SEPTEMBER 30, 1958 slab; dose-rate buildup factors; the heat deposited of the heating results with an empirical formula x throughout the slab; and the energy and angular is described here. distribution of the gamma rays reflected from and transmitted back into the slab, A comparison of . the dose-rate buildup factors for normal incidence obtained from this calculation with the factors obtained from an empirical formula was also given previously,! 1 and a recently completed comparison The heating results are given as the percentage of the total energy incident upon the slab that is absorbed in a specified region in the slab. Some typical plots of the results are shown in Figs. 4,1.7 through 4.1.10, which compare the Monte Carlo results averaged over a region of four in- V1L, A. Bowman ond D. K. Trubey, ANP Quar. Prog. tervals with values obtained by using the following Rep. Sept. 30, 1957, ORNL-2387, p 320. empirical formula: pa(EO,Mafx) X, +x, J (E ,0,x,Mat )= [sec 0 ————— exp = | ~—— X PRI X u,(E 5, Mat ) cos 0 ] X, X, . X, ] X+ %, . ] X, exp - + _ - exp — “1\ecos § © 2\cos @ O cos @ %2\ cos 9" 0 cos 6 ,ua(Eo,Mafl) X+ %, 2 exp —<4cos (1l —cos ) |1~ ’ ,(E g Mat,) (xy +x)4+ 1 VE, where J (Ey,0,x,Mat ) = percentage of total energy incident upon the slab that is absorbed in the slab at point x per mean free path, x, = number of mean free paths of the first material, x , = number of mean free paths of the second material, energy of the incident gamma ray, Il 0 ¢ = angle between the direction of the incident gamma ray and the normal to the slab, '“a(EO'me) energy absorption coefficient ,uz(E'o,Maf ) " total absorption coefficient (COS ) = NDA point isotropic energy absorption buildup factor for the first material,!’ B, ( 5 > NDA point isotropic energy absorption buildup factor for the second material, cos Xy +x, exp - ( 3 )— exponential attenuation to point heating is calculated, and cos ,ua(Eo,Maf]) X, +x # (g, Mat,) (x, +x2)4 +1] VE, is the empirical short-circuiting correction, . exp~<4cos (1l —cos @) |1~ 125 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2-01-059-238A p} = NORMAL THICKNESS IN MEAN FREE PATHS AT INITIAL ENERGY { / LEAD pd 2 / LEAD WATER e S / @ 7 LEAD b3S A WATER 2 - Hen ;s D = z WAT o 4 LEAD - AfER L . = HTh,0 a =] o [V Z 5 WATER o Q (s3] [=4 a / w6 WATER LEAD 7 WATER / LEAD 8 WATER LEAD r/4v'+fir/4t+— ’74—"’4‘ 7/g ——w= 7, TOTAL THICKNESS (mean free paths) ————m=— Fig. 4.1.6. Lead and Water Slab Configurations Used in Monte Carlo Calculations, The first bracketed factor represents the fraction of incident energy expected to be deposited per mean free path if the scattered gamma rays are neglected. The second bracketed term is the build- up factor. Near the boundary (x , small), where the spectrum is largely determined by the first material, the buildup is given by the first term. This term damps out as x, gets large, and the buildup factor is characteristic of the second material, The buildup factors used in the formula were the results of the weli-known NDA moments method calcula- tion. 12 The energy absorption buildup factors used were for a point isotropic source, since these were the only buildup factors presented in ref 12, 24, Goldstein and J. E. Wilkins, Jr., Calculations of the Penetrations of Gamma Rays. Final Report, Y0-3075 (June 30, 1954). 126 UNCLASSIFIED 2-04-059- 354 2.00 — C 3412 | REFLECTEC: ©.759 ‘ --i——— TRANSMITTES: G148 1.80 ~—— ABSORBED: 99.03 T — Eo= 3 Mev O FORMULA 5 1 | 9= 60deg ® MONTE CARLO " ( I 11 P | & 7 11 N = — — — : 3 | R .- [ = | L .._L_L__ B:J q ) ‘ ' 1 [ T a 140 —t — o j o | w 1 | [ 1 2 - —r i w 1 T,_.._,fi _ 2 120 r— W | | IS ‘ - E:" \ l 7 ! 1 > \ | w S — Z 1.00 1 T . s = \ <] I S S © \ ~ 0.80 g e R Q — T = \ 2 \ I 080 O . et w O g \ 2 0.40 a % - © A = 0.20 L Y iz 1 0 e B B O 0 { 2 3 4 Hor, NORMAL THICKNESS OF COMPOSITE SLA8 IN MEAN FREE PATHS AT INITIAL ENERGY Fig. 4.1.7. Gamma«Ray Energy Absorption in a Lead Shield as @ Function of the Shield Thickness, The third bracketed factor is the ‘‘short-circuit- ing’’ factor. An attempt was made to separate the effects of the various parameters in the exponent. The factors which depend on the angle reach a peak at 60 deg. It seemed reasonable that a peak might occur at about 60 deg when the combination of a decreasing path length and a decreasing cross section, as well as the final energy of a scattered gamma ray as the angle of scattering increases, were taken into account. The effect of distance from the initial boundary also shows a peak (near 1 mfp). There is little short circuiting at short distances, since the heating is due largely to first collisions. The short circuiting damps out at large distances since the buildup factor adequately accounts for the scattered gamma rays far from UNCLASSIFIED 2-04-059- 359 t 6431 T ; REFLECTED: 0,019 | TRANSMITTED: 3.21 1.80 ABSORBED: 96.8 = L E;= 6 Mev O FORMULA a | 8= O deg @ MONTE CARLOD w 1 : -t J “ 4160 LEAD % WATER g U B T 3 e _ : BN : @ 440 ‘ ] ] 1 @ I z : - ] + i [ [ < 420 1 - - > | I E | &) | 14 w —t z w . | N N— 1 I 100 T i a [ I ] : T N Q 14 w o o 2.80 w [14] @ (o] wn m < > 2.40 o x w 3 > e < (L] > 160 18 2] \ g o ¥ { 5 1.20 \ - w © — g \ ":;’ 0.80 \\ [0 w = \ 040 ’ \ o l \'.-'0—--.. - 0 1 2 3 4 Haor, NORMAL THICKNESS OF COMPOSITE SLAB IN MEAN FREE PATHS AT INITIAL ENERGY Fig. 4.1.9. Gomma-Ray Energy Absorption in a Water Shield as a Function of the Shield Thickness. described above, which had a very limited number of parameter values, and therefore it is possible that the fit would not be so good for other values of the parameters, particularly those outside the energy range examined. The worst fits were obtained for low energies, especially with lead following water, but even in these cases the error was less than 20%. In nearly all the cases examined, the error was less than 5%, This work has been described in more detail in a topical report. '3 13, A. Bowman and D. K. Trubey, Deposition of Gamma-Ray Heating in Stratified Lead and Water Slabs, ORNL CF-58-7-99 (July 28, 1958). 127 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2-01-059 - 3686 2.00 T 3472 T ] REFLECTED: 2.48 TRANSMITTED: 0.298 1.80 ABSORBED: 97.2 = Eo= 3 Mev O FORMULA g 8=60deq @ MONTE CARLO w : | A L 160 — WATER — LEAD — 2 — ] 1 = <1 s x ; . Lt | T 140 ’: 2 x } : I | ] | [4p] 2 20 | 3 | & [ i € \ | | g y a i w 1 > oo L a4 1.00 : z 1 d=225cm 1.60 MALLORY 1000 (FRONT SAMPLE) 7 = DISTANCE BETWEEN REACTOR AND SAMPLE 1.40 1.20 // d=4%7.5¢cm M, HEATING {w-em™3 kw1) MMM 0.60 \ LALLM 0.40 /. Be // 7 Be (FRONT SAMPLE ) 71/ (REAR SAMPLE) o _ 0.20 _ MALLORY 1000 4 | - T (REAR SAMPLE) g=27.0cm gd=33.4cm N 2 /////// LiH LiH {FRONT SAMPLE )——— {REAR SAMPLE) 3 4 5 6 SAMPLE THICKNESS (cm) Fig. 4.1.19. Calculated Total Heating Produced by the BSF Gammo-Ray Energy Specttum in All Samples in the Beryllium—Mallory 1000—Lithium Hydride Configuration, A MONTE CARLO CODE FOR THE CALCULATION OF DEEP PENETRATIONS OF GAMMA RAYS S. K. Penny A Monte Carlo code is being developed to calcu- late at a point detector the angular and energy dis- tributions of gamma rays emitted from a monoener- getic, point isotropic or point monodirectional source embedded in an infinite homogeneous iso- tropic medium of constant density. The word “‘monodirectional’’ is used here in a loose sense; the emission directions are actually in a half-cone specified by a polar angle, where the source- detector axis is the polar axis, and they are uni- formly distributed in the azimuthal angle. The code is now in the ‘‘debugging’’ stage on the IBM-704 electronic data processing machine. It is hoped that by using special techniques it will be useful for penetrations as deep as 20 mean free paths, In its final form the code can be used for neutron penetration also, provided a differential cross section for scattering is given, The purpose of this code is twofold, First, it is hoped that it will shed some light on the discrep- ancy between calculated and experimental energy distributions from a Co%® source in water, the chief difficulty being the magnitude of the distri- butions. The calculated values were obtained by the moments method,2%~22 and the experimental results were obtained by Peelle, Maienschein, and Love.?4 Second, there is merit in experimenting 20L. V. Spencer and U, Fano, Phys. Rev. 81, 464 (1951); see also J. Research Natl. Bur. Standards 46, 446 (1951). 2.lL. V. Spencer and F. Stinson, Phys. Rev. 85, 662 (1952). 22, Fano, J. Research Natl. Bur. Standards 51, 95 (1953). 23y, Goldstein and J. E. Wilkins, Jr., Calculations of the Penetrations of Gamma Rays. Final Report, NDA-15C-41 or NY0-3075 (June 30, 1954). 24R, W. Peelle, F. C. Maienschein, and T. A, Love, Energy and Angular Distribution of Gamma Radiation from a Co%0 Source After Diffusion Through Many Mean Free Paths of Water, ORNL-2196 (Aug. 12, 1957). 133 ANP PROJECT PROGRESS REPORT with specid®¥onte Carlo techniques in order to see when und where they can be applied. The special techniques referred to above are importance sampling on the first-collision distribu- tion coupled with double systematic sampling, statistical estimation, splitting and Russian roulette, and a special form of output, Other features in the code are an energy cutoff, a weight cutoff, and a cutoff for distance from the detector. Importance Sampling The importance sampling is discussed here from the viewpoint of a point isotropic source. The technique for the monodirectional source is simpler and straightforward. In the code, an attempt is made to estimate z= [dx fdy fdw ... fdt bx,yw,et) 2(x,y,w,0t) and it is assumed that the probability distribution function can be written as b(xfylwl"‘lt) = f(xf)’) g(w,...,t) . By definition z(:x,y) = fdw... fdt (W, eee,t) 2(%,¥,00,00e,t) 22(:x,y) = fdw... fdt glw, vee,t) zz(x,y,w,...,t) and f(x,y) is the probability distribution function for the first collision, Further, x is the cosine of the angle between the direction of emission and the source-detector axis, and y is the number of mean free paths to the first collision. If it is desired to sample from f*(x,y) instead of from f(x,y) (impor- tance sampling), then — [(x,y) dx Jdy f*(x,y) z(ix,y) ——— =% ; fax fdy f*(x,y) z(:x,y) *y) however, the variance is not usually the same when importance sampling is employed as when the normal sampling is used. The probability distri- bution function for the first collision that gives the minimum variance is of the form: *(x,y) = f(e,y) V22ex, ) N, where N is the normalizing constant. Kahn?® shows this for importance sampling with one variable, and 25y, Kahn, Applications of Monte Carlo, AECU-3259 (1954) (rev ed. April 27, 1956). 134 his argument can easily be extended to two vari- ables, The code will begin with the task of esti- mating z2(:x,y)by performing an auxiliary Monte Carlo calculation, |nstead of sampling from a first-coliision distribution, x and y will be fixed and z2(:x,y) will be estimated, This will be done for a number of points {x,y). This will then provide estimates for the functions f*(x,y), G(y:x), and F(x), which will be placed in tabular form in the machine. The functions G(y:x) and F(x) are defined as f_);o dn [*(x,n) Glyx) = S dn e and Flx) = f_: d«ff_: dn *(€m) . Once the auxiliary Monte Carlo problem has been completed and the importance distribution functions have been calculated, the main Monte Carlo routine is started. Double Systematic Sampling The importance distribution for the first collision is sampled with the use of a double systematic sampling technique, 2% as follows: 1+ ]/2 F(x1)= i=0,],2,o..N_] i+ % G()/l:xz) = ] = O, ]’ 2, see N - ] . in these expressions, N = 2%= number of histories; a is a positive integer 20. The point (x ,y )} is then found by interpolation in the tables. The 's must be randomized, since they should not be correlated to i, This can be done by using the periodicity of the pseudorandom numbers generated by the con- gruence method. Each binary bit in a pseudorandom number has a certain period, as illustrated below: 26 2524 23 22 21 20 20 erind x x x 0 1 bit eee X X X 26peyised by S. K. Penny and C, D, Zerby, ORNL. If it is desired to obtain all the even integers, 0, 2, . (2%%1 = 2), logical computer operations are used to obtain from a pseudorandom number the fol- lowing number: 20 20~1 232271790 000...0 x x period we x x x 0 bit Then, generating 2% random numbers in succession will provide the 2% integers desired in random order, and the random correspondence i €>7 can be obtained. Of course, a separate random number generator must be provided which has the same basic random number but an initial random number deep in the series. When a first-collision point is chosen by the sampling technique, the initial weight must be modified by the factor flx oy V(2 0y) o Statistical Estimation The technique of statistical estimation will be used at every collision; that is, the probability of the radiation reaching the detector after the colli- sion will be calculated and used to estimate the flux at the detector, Weighted Isotropic Scattering It is planned to sample the direction after collision from an isotropic distribution, rather than the Klein-Nishina differential cross section, and “‘weight’’ it with the Klein-Nishina distribu- tion, This procedure was derived primarily for simplicity and to save in calculational cost, [t is certainly biased sampling, but for small sepa- ration distances it is probably the best method of sampling. For large separation distances the variance of the estimated flux will increase, but the increase may be compensated for intuitively by the use of importance sampling and splitting. Splitting and Russian Roulette The techniques of splitting and Russian roulette are to be used, with the splitting occurring at the collision point rather than at a splitting boundary. In other words, if the particle crosses b boundaries with the boundaries successively closer to the detector, it will be split into 28 particles at the final collision point. If a particle crosses b boundaries with the boundaries suc- cessively farther away from the detector, it then PERIOD ENDING SEPTEMBER 30, 1958 has a probability of 279 of surviving., Hf it sur- vives, its weight is increased by 2%, The split- ting boundaries are concentric spheres with the detector as their center. Again the reason for this technique of splitting at the collision point is chiefly simplicity. In all cases the statistical estimation is performed before either the splitting or the Russian roulette techniques are used. Although it may be argued that splitting at the collision point will introduce biasing, it remains to be shown whether the increased cost associ- ated with an unbiased calculation would be warranted. Output The output will be of a special form; there will be no histograms. The usual process, which provides histograms, consists of estimating the flux at the detector and then determining the angular and energy distributions from the total flux arriving in a sector. The new code makes use of the expansion: k+127+1 2 2 Pk(w) Pl(T) 4,0 = [ do [ a7 Py(@) Py 91D,0,7) E-E c T=2——" -] E, - E, E = energy of radiation, ! E = cutoff energy, m I initial energy, cosine of angle between source de- tector axis and direction of radiation, € Il D = separation distance, P, P ;= Legendre polynomials, é(D,w,T) = flux at D per unit w per unit 7. At each collision statistical estimates are made of the A, (D) term?7 rather than simply the A . (D) term, which is easily seen to be the total flux. Single scattering is treated separately from multiple scattering. The code allows &,/ < 15 for 2755 sugyested by R, R, Coveyou, ORNL. 135 ANP PROJECT PROGRESS REPORT single scattering and &,/ < 7 for multiple scatter- ing. Of course, the coefficients for single scat- tering and multiple scattering are completely additive, A choice may be made of any combination of the following values in the output: number flux, energy flux, energy deposition in the medium (ergs/g), and energy deposition in carbon (ergs/g) at the detector, carbon being the standard material for dose-rate calculations or measurements, Also, the variance of the estimated coefficients (4, ,) will be automatically calculated. Testing Procedure The code will be tested against three standards. First, results for no energy degradation and iso- tropic scattering will be checked against the exact solutions given by Case, deHoffman, and Placzek. 28 Second, results for a medium other than water will be checked against the energy distributions of energy flux fumished by the moments method,2%~23 Third, results for water as a medium will be checked against the experimental work of Peelle, Maien- schein, and Love?? and values obtained by the moments method, 2923 A MONTE CARLO CALCULATION OF THE NEUTRON PENETRATION OF FINITE WATER SLABS J. B. Hilgeman?? C. D. Zerby Monte Carlo calculations? of the neutron dose- rate distribution beyond finite water slabs were performed for various parameter combinations, In the idealized probiem a plane monodirectional, monoenergetic beam of neutrons was assumed to be incident on a water slab of infinite area and finite thickness. The angle between the incident neutron direction and the normal to the slab was denoted F. L. Keller as 6, and the number of neutrons which penetrated the siab was assumed to be recorded by a spherical 28y M. Case, F. de Hoffman, and G, Placzek, In- troduction to the Theory of Neutron Diffusion, Los Alamos Scientific Laboratory, Los Alamos, 1953, 29On assignment from the U.S. Air Force. 30The Monte Carlo machine program which was used in this study is a revision of the neutron air-scattering program which has been described by C, D, Zerby, A Monte Carlo Calculation of Air-Scattered Neutrons, ORNL-2277 (April 23, 1957). This program employs the method of statistical estimation. 136 detector. The space around the detector was divided into a number of solid angle intervals with the apex at the detector point, and the number of neutrons which entered into each of these solid angle intervals was recorded. The energy spec- trum of the radiation which entered each of the solid angle intervals was determined by dividing the energy range from some cutoff energy, E_ the source energy, E . into an arbitrary number of equal intervals and recording each contribution in to its appropriate energy interval, This energy spec- trum was then used to determine the tissue dose- rate contribution from each energy and solid angle interval, The scattering probability was assumed to be isotropic in the center-of-mass system. This is essentially exact for the scattering by hydrogen (the major scatterer) and is fairly satisfactory for the scattering by oxygen over most of the energy range of interest in these calculations. The density of water was taken to be 1 g/cm3, All the results were normalized to one incident neutron per second per square centimeter of slab surface. A parameter study was carried out in which plane monodirectional beams of neutrons with energies, Eqy of 0.55, 1.2, 2, 4, 6, and 8 Mev were incident on the water slabs at angles, Oy of 0, 30, 60, and 75 deg. The thicknesses of the slabs ranged from I to 8 mfp (mean free paths). The number of case histories used for a particular problem varied from 5,000 to 10,000, depending upon the slab thickness and the angle of incidence. The solid angle inter- vals at the detector were defined by 90-deg azimuthal angle intervals and 15-deg polar angle intervals with respect to a polar axis which was normal to the slab surface., A cutoff energy, E_ of 0.1 Mev was used. This value was chosen to approximate the low-energy cutoff of most of the present dosimeters. Fifteen equal energy intervals were used to determine the spectrum, Figures 4.1.20 through 4.1.23 show a set of representative curves which were generated from data obtained from this study. In these figures dose-rate buildup factors, B , are plotted as a func- tion of the finite slab thicknesses for various energy neutrons incident at angles, 60 of 0, 30, 60, and 75 deg. The dose-rate buildup factor may be defined as the ratio of the total dose rate at the detector to the dose rate which would result if every collision were equivaient to an absorption, It should be noted that the results are plotted as a function of UNCLASSIFIED 2-04-059-259 1000 500 ol ' 200 o 100 1 / % 50 /l / “ 20 / © [ o2 o 10 ] // / { / / ; / / / ‘/ N/ e / / // I 8o=0 deg | 2 %/ P i 1 0 i 2 3 4 5 6 7 8 NORMAL THICKNESS IN MEAN FREE PATHS Fig. 4.1,20, Dose-Rate Buildup Factors for 0.55-Mev Neutrons Incident at Yarious Angles on Finite Water Slabs. PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED 2—01—059-260 1000 500 *— 8,= 75 deg ~~—— 200 100 50 ""'--....____ BUILD-UP FACTOR \ 20 / o Ob //Q’/ R0 £, ] // L~ 2 %l/ Q { 2 3 4 5 6 7 8 NORMAL THICKNESS iIN MEAN FREE PATHS Fig. 4.1.21. Dose-Rate Buildup Factors for 2-Mev Neutrons Incident at Yarious Angles on Finite Water SIUbS. 137 ANP PROJECT PROGRESS REPORT UNCLASSIFIED UNCLASSIFIED 2—-01-052-261 2-01-059-262 1000 1000 . 500 Jo P ‘T R I 500 T /90= 75 deg [ o 9] M 200 f=— S R 200 y QO 100 R RS s +— — __ o 100 . / _ f 50 — P S S 50 / el 5 o] o [ a i o a = N ? f 3 / 3 = = @ o o 20 , R — - 2o / / ® [ bQJQ/ e 10 l // 90= 60 dEQ QO// / 7 ]l // — / / / / / 5 / [ / ] 5 / / / // 8q= 30deg V % // | / o 8y = Odeg /_,-——" 8,=0deg / %/ —] > V . / v | LA = | 4 4 | 0 1 2 3 4 5 6 7 8 f NORMAL THICKNESS IN MEAN FREE PATHS 0 1 2 3 4 5 8 7 8 NORMAL THICKNESS IN MEAN FREE PATHS Fig. 4.1.22, Dose-Rate Buildup Factors for 4-Mev Neutrons Incident at Various Angles on Finite Water Fig. 4.1.23. Dose-Rate Buildup Factors for 8-Mev Slabs. Neutrons Incident at Various Angles on Finite Water Slub5| 138 the normal thickness of the slab in mean free paths. When plotted in this manner it is obvious that the buildup factor for a given energy and slab thickness should increase with increasing values of 4. be- cause of the ‘‘short-circuiting’’ effect, These re- sults were compared with the results of a similar calculation which was performed by QObenshain, Eddy, and Kuehn,3! and the buildup factors from the present calculation were found to be consider- ably smaller, in general, than their values, The cause of this apparent discrepancy is not yet known. Figures 4.1.24 through 4,1.27 show the angular distribution of the scattered neutron dose rate at UNCLASSIFIED 2-01-059-255 2 N | \DETECTOR' 40-7 \ \ | | JO S | - \ AN [ i N\ 3-mfp-THICK SLAB : \\ N, _ - |® 5 | N ol HE N .:‘:’ [ \ \ \ 5|8 e 2 N aly et 4 mfp (=] -._,m 10—3 \ N ) w T =y a & 5 Q (] = Q o - @ =4 2 a A (92 S 9 . 10” ~ B \\ 5 T~ — 2 S i : 10710 E -4 ) Cos a Fig. 4.1.24, Fast-Neutron Dose Rates at Rear of Finite Water Slabs Resulting from Multiply Scattered Neutrons (EO = 0,55 Mev), PERIOD ENDING SEPTEMBER 30, 1958 the detector for neutron beams of various energies normally incident (6, = 0) on water slabs of various thicknesses, Since there is azimuthal symmetry at the detector for these cases, the curves are plotted as dose rate per steradian versus cos @, where a is the polar angle at the detector. These plots were 31, Obenshain, A, Eddy, and H. Kuehn, Polyphemus. A Monte Carlo Study of Neutron Penetrations Through Finite Water Slabs, WAPD-TM-54 (1957). UNCLASSIFIED 2—01—059-2564A -6 \ 10 N . 5 \\ ) \ \\{fp_THICK sLaB /—T: . \ \ - | %10 \ 5| HE N\ s s T Tg, \ AN o g \ — 2 2 o wl 468 \\ \imfp W b AN e (o] 5 \ . = N o z AN 1 \ w a 2 N w 15° % - % DETECTOR Cos a Fig. 4.1.25. Fast-Neutron Dose Rates at Rear of Finite Water Slabs Resulting from Multiply Scattered Neutrons (E'0 = 2 Mev). 139 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 6 2—01—-059-257 10 : : ‘ i ! 5 o0 S I Z \"'kDETECTOR ] —— 21 ] S e N i — 3§ | 34g7 | o . AR Z-mfp-THICK SLAB | :J 5 A ek ——— “. P | 3 S 5 (1) — < @ w w O ) 2 = (@] [ia 5 o 108 T - w a = 5 2 |- 107° ' : Cos a Fig. 4.1.26, Fast-Neutron Dose Rates at Rear of Finite Water Slabs Resulting from Multiply Scattered Neutrons (EO = 4 Mev), generated by drawing smooth curves through the histogram output of the machine calculations. For cases of normal incidence (65 = 0), a very large fraction of the total scattered dose rate at the detector is contributed by neutrons which have undergone only one scattering event in the slab, The fraction of the total scattered dose rate which was contributed by singly scattered neutrons for each of these cases is given in Table 4.1.1. From this table it is seen that single scattering calcu- lations may be expected to yield fairly accurate results for dose rates from neutrons which are normally incident on thick water slabs. The higher orders of scattering become more important, how- ever, as the angle of incidence, 90, is increased. A detailed description of ail of the results of this parameter study will be given in a separate report.3? 32, B, Hilgeman, F. L. Keller, and C. D. Zerby, Neutron Penetration of Finite Water Slabs, ORNL-2463 (unpublished). 140 UNCLASSIFIED 2-01-059-258 SLAB 0 — SN DETECTOR __ | \ | ‘ f-mfp-THICK SLAB 107° I ~ - _ 1 —1 rep-hr ' steredian source neutron-cm—2. FAST-NEUTRON DOSE RATE ( CcOs @ Fig. 4.1,27, Fast-Neutron Dose Rates at Rear of Finite Water Slabs Resulting from Multiply Scattered Neutrons (E0 = 8 Mev), Table 4.1.1. Fraction of Total Scattered Dose Rate at the Detector Contributed by Neutrons Which Have Undergone Only One Scattering Event (90 = () Fraction of Dose Contributed by E,(Mev) Singly Scattered Neutrons 1 mfp 3 mfp 4 mfp 6 mfp 0.55 0.827 0.864 0.902 2.0 0.727 0.86 0.846 4.0 0.791 0.832 0.877 8.0 0.707 0.829 0.762 PERIOD ENDING SEPTEMBER 30, 1958 4.2. LID TANK SHIELDING FACILITY * W. Zobel STUDY OF ADVANCED SHIEL DING The configurations for which the fast-neutron and " MATERIALS (GE SERIES) gamma-ray measurements are reported here are 4-J, L. Jung 4-K, and 4-L, in which the thickness of a stainless steel layer adjacent to the source plate was in- creased in steps from 1 to 3in. In Fig. 4.2.1 the gamma-ray dose rates beyond these configurations are compared with the gamma-ray dose rates in oil only. For purposes of comparison, the dose rates beyond configuration 4-A, which consisted of 4 in. The experiment designed by GE-ANP for studying the production of secondary gamma rays in con- figurations containing advanced shielding materials has been completed. Most of the results of the fast- neutron and gamma-ray dose rate measurements bey'ond the various conflgurc'mons lisgd in this ex- of stainless steel in oil, are also presented. periment were reported previously. < The remain- Corresponding fast-neutron dose rates are shown in ing data on thermal-neutron flux measurements and Fig. 4.2.2. on fast-neutron and gamma-ray dose rate measure- ments are presented here, The configurations for which the data were obtained are described in Table . . 4,2,1. The dimensions and compositions of the ma- Dec?‘;;"‘ 5:9“5‘]7’,' ?grffi.%;fl;',”;'éflfvp Quar. Prog. Rep. terials used were the same as those given in Table 2|, Jung, ANP Quar. Prog. Rep. March 31, 1958, 4,2.2 of ref 2. ORNL-2517, p 90. Table 4.2 1. Description of Latest Configurations Tested in GE Series” . Confi " zgr Location of Unavoidable Oil Gap nfi ° N gut: ren Description of Configuration Probe” Within Configuration® umbe ) e (cm) (em) 4-A 4 in. of stainless steel + oil 16,2 2.8 4-) 1 in. of stainless steel + oil 7.0 1.3 4-K 2 in. of stainfess steel + oil 10.0 1.8 4-L 3 in. of stainless steel + oil 12.8 2.0 4-N 4 in. of stainless steel + ]/2 in. of boral + oil 17.6 3.0 4-Q ]/2 in., of boral + 4 in. of stainless steel + ]/2 19.6 3.6 in, of boral + oil 4-T 4 in. of stainless steel + ]4 in. of boral 4 oil 17.6 3.7 5-A 4 in, of Be + 12 in. of LiH + ail 45.8 1.1 41-A 4 in. of stainless steel (dry) + oil in an Al tank 15.4 1.9 (Jrine-thick woll) (oir gap) Oil Qil only in canfiguratian tenk 3.2d %Includes some configurations described in ANP Quar. Prog. Rep, Dec. 31, 1957, ORNL-2440, p 263, ond in ANP Quar. Prog. Rep. March 31, 1958, ORNL-2517, p 90. Actual distance from source plate assembly to back of salid configuration. “Includes 1-cm recession of the aluminum windaw in the configuration tank. Pasition of inside surface of the aluminum window in the configuration tonk, 141 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2-01-057-71-435 104 T T LT T T4 T s RN =R SERSERE : o Ee I B A | s 7_,,_T50:~cc IPN C?AMBER_” _jr__ —_ ,—) ‘ ] GAMMA-RAY TISSUE DOSE RATE (ergs-g ' - hr ! S0 . ~ANTHRACENE ‘- } SCINTILLATION ] P ‘ DOSIMETER 2 ‘ - ! ‘ o b CONFIGURATION STAINLESS | | NUMBER STEEL (in) 5 - g ——4-J 1 a - 4-K 2 v 4-L 3 2 1 -0 4-A 4 X OIL ONLY L S PO S 2c 40 60 80 100 120 140 160 75, DISTANCE FROM SOURCE PLATE {em) Fig., 4,2.1. Gammao=-Roy Tissve Dose Rates in Oil and Beyond Configurations 4-A, 4-J, 4-K, and 4-L: Effect of Increasing the Thickness of a Stainless Steel Configu- ration, The thermal-neutron fluxes beyond these same configurations are presented in Fig. 4.2,3. Intro- ducing 1 in, of stainless steel to the oil medium re- duced the flux by 11.5%. Stepwise additions of 2, 3, and 4 in, reduced the flux by 24.6%, 32.2%, and 42.1%, respectively, at a distance of 100 ¢m from the source, compared with the flux measured in oil alone, The curves in Fig. 4.2.4 indicate that placing ]4 or I/2 in. of boral behind the 4 in, of stainless steel had only a negligible effect on the flux 100 cm from the source. It also appears that ]/2 in. of boral in front of the stainless steel in addition to the 142 UNCLASSIFIED 2-01-0587 ~71-436 R - " R I N CONFIGURATION NO. STAINLESS STEEL (in} — : 1 | — a- a- 4. a4- . oieonty o : ,,,Z[ H : | t ) W FAST NEUTRON TISSUE DQOSE RATE (en;s-g_“hr_I o] 20 30 40 50 60 70 80 20 100 Ho Z,, DISTANCE FROM SOURCE PLATE (cm} Fig. 4.2.2, Fast-Neuvtron Tissue Dose Rate in Oil and Beyond Configurations 4-A, 4-J, 4-K, and 4-L: Effect of Increosing the Thickness of o Stainless Steel Configuration, PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED 2-01-057-74-432 Sx 107 — ] T Il 2 4 (.‘; 100 3 LY N - - W T 5 LAY S LAY - — - I\ eacd A\, ¥ 5 \ 6 \\\\\ CONFIGURATION NO. STAINLESS STEEL (in.) 40 e . - JE— 1 [— p—— - —— " SecC THERMAL NEUTRON FLUX (neutrons-cm™ 2 n -2 10 0 10 20 30 40 50 60 70 80 20 100 10 120 130 140 150 2y, DISTANCE FROM SOURCE PLATE (cm) Fig. 4.2.3. Thermal-Neutron Fluxes in Oil and Beyond Configurations 4-A, 4-J, 4-K, and 4-L: Effect of Increasing the Thickness of a Stainless Steel Configuration. 143 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2--057-71-433 . | i e — e e e ] CONFIGURATION — BORAL STAINLESS BORAL ) NUMBER (in) STEEL (in) (in) 0 4 0 +0IL 0 4 14 +0IL 0 4 o +0IL VE 4 1o +0IL — 0 4 (DRY) +OIL IN Al TANK | THERMAL NEUTRON FLUX (neutrons-cm 2-sec”w™!) 0 10 20 30 40 50 60 70 80 20 100 110 i20 130 14Q 150 Zey DISTANCE FROM SOURCE PLATE {cm!} Fig. 4.2,4, Thermal-Neutron Fluxes Beyond Configurations 4-A, 4-N, 4-Q, 4-T, and 41-A: Effect of Inserting Boral Within Configurations Containing Stainless Steel. 144 2 in. of hgral behind the stainless steel did not affect the flux at 100 cm appreciably. The boral did, however, cause a decrease in the thermal- neutron fluxes at distances greater than 120 c¢m from the source, possibly as a result of the re- duction in the number of photoneutrons arising from high-energy gamma rays. High-energy gamma rays originate in the stainless steel following the capture of thermal neutrons, and the boral reduces the number of captures by absorbing most of the neu- trons which are thermalized in oil and then re- flected back into the stainless steel, 374 Thermal-neutron flux measurements beyond con- figuration 41-A are also plotted in Fig. 4.2.4 for comparison with the curve for configuration 4-A. At 100 ¢m from the source the flux beyond con- figuration 41-A was 16.5% higher than that beyond configuration 4-A. The two configurations differed only in that configuration 41-A was dry; that is, the 1 cm of oil between the aluminum window of the con- taining tank and the first slab of stainless steel, as well as the oil between the slabs, was removed. As described previously,? several types of gamma- ray shields were used in conjunction with 4 in. of beryllium and 12 in. of lithium hydride. Measure- ments beyond a configuration consisting of only the beryllium and the lithium hydride (configuration 5-A) were made to assist in evaluating the data. In order to complete the set of measurements, thermal- neutron fluxes beyond this configuration are shown in Fig. 4.2.5. STUDY OF SECONDARY GAMMA-RAY PRODUCTION IN LEAD J. M. Miller A comprehensive program for measurement of the production of secondary gamma rays in lead was initiated at the Lid Tank Shielding Facility (L. TSF) for comparison with the calculation being made by the Nuclear Development Corporation of America (NDA). In the tests run thus far, thicknesses of lead varying in l/z-in. steps from 1 to 6 in. and additional thicknesses of 7]/2 and 9 in. have been 3p. K. Trubey, An Estimation of Photoneutrons from Carbon-13 in an Oil Shield, ORNL-2200 (Aug. 13, 1958). 4. T. Chapman et al., Measurement of an Effective Neutron Removal Cross Section of Lithium at the Lid Tank Shielding Facility, ORNL CF-54-11-3, p 3 (Nov. 2, 1954). PERIOD ENDING SEPTEMBER 30, 1958 utilized. In each case the lead was kept dry by placing it in a steel tank (5/ in.~thick walls) posmoned against the source plate. An aluminum tank (Y-in.-thick walls) filled either with oil or with borc:fec‘3 water was always placed immediately behind the lead. The steel tank which holds the entire con- figuration has a ¥-in.-thick aluminum window on the source side, and the recession in the tank wall between the window and the first slab of lead in- troduces a l-cm-thick air gap at this poeint in each configuration. Gamma-ray tissve dose-rate measure- ments were made in the oil or borated water, and in some cases fast-neutron tissue-dose-rate and thermal-neutron flux measurements were made in the oil. The gamma-ray dose-rate measurements in oil and borated water beyond the various thicknesses of lead are plotted as a function of the distance from the source plate in Fig. 4.2.6. The measurements in borated water were made primarily to determine the effect of suppressing the thermal-neutron flux beyond the lead. The gamma-ray dose rates in the borated water were a factor of 3 lower than the dose rates in oil in the region close to the lead and a factor of 13 lower approximately 120 cm beyond the lead. Cross plots of the data from Fig. 4.2.6 are shown in Figs. 4.2.7 and 4.2.8. The plots in Fig. 4.2.7 present the gamma-ray dose rates at points 100 cm beyond the configurations, corrected for the inverse 72 attenuation, and show that most of the primary gamma rays are attenuated by the first 3 in. of lead. The gamma rays observed beyond greater thick- nesses of lead are practically all secondary gamma rays. The plots in Fig. 4.2.8 present the gamma- ray dose rates beyond the various configurations at distances of 80 and 100 cm from the source. This figure shows that a lead thickness of 3 in. gives the maximum effectiveness in reducing the gamma- ray dose rate at a fixed distance from the source and that increasing the thickness beyond 3 in. does not result in a further reduction. The thermal- and fast-neutron measurements in oil beyond the various configurations are plotted as functions of the distance from the source in Figs. 4.2.9 and 4.2.10, respectively. This study will be continued with configurations in which beryllium and lithium hydride will be used in combination with the lead., 145 ANP PROJECT PROGRESS REPORT a 2-01-057-0-74-437 n 2 T ','U CONFIGURATION 5-A: 4 in. OF Be +1{2 in. OF LiH + OIL o2 10 N N\ AN . N\ d | M THERMAL-NEUTRON FLUX (neutrons-cm™ 2 N 5 - \\ — — \‘k\ — 2 \ 1 \\\ AN . N\ N\ N N {0 40 50 60 70 80 90 100 {10 120 130 t40 z,, DISTANCE FROM SOURCE PLATE (cm) Fig. 4.2.5. Thermal-Neutron Fluxes Beyond Configuration 5-A, 146 PERIOD ENDING SEPTEMBER 30, 1958 UNCL ASSIFIED 2-04-057-73-428 5 x 10° I\ CONFIGURATION NO. COUNTER TYPE LEAD (in. 2 N\\t tin- _ o PM({A) \ i-a { b 50 cc } ! o PM(A) 10° A\ “\\\ f-b { . 50 ce } e _ ' PM{A) < 5 \\\\\\\‘\\‘ Ol 3 4-c { a 50 (cc) } 2 N A, _ A g PM{A ° AN AV - {. 50 cc } 4Y, PM (A \\. ) \\1 =D { ; 50 (cc) } ° » ™ \ \\\X - {! PwA) ab, \ ~ B-H.0 v 50 cc NS e 0B \ \ ¢ ¢ 50 cc 102 N\ O \‘ ~ X NS T \\\\ AN\ ] * 5 g N\ TS £ AN \\j&\ I 2 \\\\ P\\\\ u'.;J 10 \\ AN \\\L ~ g ANAN ~ ~ X AN AN L AN AR NN 8 5 AN RN N N 8 N\ AN . A\ : NG w (o] S \\\/ \\\ ' g : \\ \\ I R : 3 < O N SN \ bl 5 2 N PM{A) = ANTHRACENE SCINTILLATION DOSIMETER Q/ 50 cc = 50-cc ION CHAMBER / 10~ A Y S 5 it} ) 10”2 10 20 30 40 50 60 70 80 90 100 1o 120 130 140 {50 160 Zy, DISTANCE FROM SOURCE PLATE (cm) Fig. 4.2,6. Gamma-Ray Dose Rates in Oil and Borated Water Beyond Various Thicknesses of Lead as a Function of the Distance from the Source Plate. ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2 2-01-057-73-427 {10 5 Iy \ IN OlL BEHIND LEAD I; ‘ \ ,_:" \ U\Kfi 5 s [\ & \ F———— @ : \ '._ < (r \ wi \ § 2 2 a—N BORATED WATER BEHIND LEAD T s s ! \ 3 \\ N S5 \ ~—] -\_A 2 o 0 2 4 6 8 10 LEAD THICKNESS (in.} Fig. 4.2.7. Gamma-Ray Dose Rates in Oil and Borated Water 100 cm Beyond Various Thicknesses of Lead {Measurements Corrected for Inverse r2 Attenuation), THERMAL-NEUTRON FLUXES MEASURED IN OIL AT THE LTSF: AN ERRATUM L. Jung Results of measurements of the thermal-neutron fluxes in oil in the usual LTSF configuration tank were presented in an earlier report.> During a periodic check of measurements, a computational error was discovered in the normalization of the instruments to gold-foil measurements in the oil which had caused the reported curve to be 16% too low; Fig. 4.2.11 shows the corrected curve. Similarly, all other thermal-neutron measurements 148 UNCLASSIFIED 2 2-01-057-73-429 //,.\fi T | | _BORATED WATER y 1 —] BORATED WATER WA R S i GAMMA-RAY DOSE RATE (ergs-g~'-hr™l-w™!) © 80c¢m FROM SOURCE PLATE 2 A 100 cm FROM SOURCE PLATE 0 2 4 6 8 10 12 LEAD THICKNESS (in.} Fig. 4.2.8. Gamma-Ray Dose Rotes in Oil and Borated Water at Points 80 and 100 cm from the Source Plate as a Function of the Thickness of Lead Adjacent to the Source, in oil which have been reported in two of the pre- vious progress reports 145 should be increased by 16%. RADIATION TRANSMISSION THROUGH BORAL AND SIMILAR HETEROGENEOUS MATERIALS CONSISTING OF RANDOMLY DISTRIBUTED ABSORBING CHUNKS W. R. Burrus® One material commonly used as a thermal-neutron suppressor is boral, a heterogeneous mixture of 5D. W. Cady and E., A, Warman, ANP Quar. Prog. Rep. Sept. 30, 1957, ORNL-2387, p 297. *Now at QOhio State University, Department of Physics, Columbus, Ohio. PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED 2-01-057-73-430 5x107 2 _.f.[\\ A P ? 107 \“ Efi\\g{*‘ iy AED A Y YUY \\ 1\\ ‘A\‘\\ 5 X LN 1 \ A ) NEEARRY \ \\ \ s 4% in. OF LEAD 108 o 6 in. OF LEAD 4 7 in. OF LEAD © 9 in. OF LEAD N R B W \ L\ Vi T R ! S L VA W =05 . Lo\ | T T, . N WY j ) g, AR _ ; A | £ 4 ('J 10 \\\\‘\‘ w0 \ AVA Y s RN Y £ ° — AR \\\ 2 ANEANNY O B R B AN ] ~ ] T - 10° ] — \ S - - e S— )\ Y 14 , RiY \ 2 ° : — = A \Y _ %J 2 — \ ] I 102 - N LY % 5 \K\ . A\ -~ 10 fl\f . — o\ \N\ 5 N\ 2 \//;D. 1 I . A\ ¥ A Y b, U 5 h WS W |/ 0 w . 2 10—4 [ — 0 10 20 30 40 50 60 70 80 90 100 10 120 130 140 Zz, DISTANCE FROM SOURCE PLATE (cm) Fig. 4.2,9. Thermal-Neutron Fluxes in Oil Beyond Various Thicknesses of Lead as a Function of Distance from the Source Plate. 149 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2—-01-057-73-43 103 - T - T A - . S . \ I I _ o N )Y_ A 1 —i 1 k— N { \ A \ 1‘ - T ‘ P \ \ \\ | e 4% inoOF LEAD | \ o 6 in OF LEAD 2 | 4 7Y% in. OF LEAD Y AW E— L W © 9 in OF LEAD [~ ] — \ \‘}\\“\ \\\ — A OIL — - 4; —] ° AN AW _ . — — AW Bl g \ oS | o NN (. _ Lo A NN\ — — = R\ e S [ ANEAN _. ] : NN 5 @ - o ‘ : i @ { \ \\ N g N/ - 8 5 - - - 77\%\7 - . ——t z NN £ } AN\ : AN 10! e S F )4 - ] A\Y Y N s N N\ . N\ 1072 A\ NN AN VI NGEx ’ \N AN 2 AN 107 0 10 20 30 40 50 60 70 80 90 100 1o 120 Fig. 4.2.]0- Distance from the Source Plate. 150 Z, DISTANCE FROM SOURCE PLATE {cm) Fast-Neutron Dose Rates in Oil, Beyond Various Thicknesses of Lead as A Function of the PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED 2-01-057-0-71-434 5X107 2 ¥ 10 \- X z\ ‘\ 5 \ L] P /"f / rd EU THERMAL NEUTRON FLUX (neutrons - ¢m 2 -sec—1-w™1) 2 ,l © GOLD FOILS 0 f——id——1 v Y—in FISSION CHAMBER A o v 3—in. FISSION CHAMBER N 12%—in. BF; PROPORTIONAL COUNTER X 0 10 20 30 40 50 60 70 80 90 100 Ho 120 130 140 150 160 z, DISTANCE FROM SOURCE PLATE (cm} Fig. 4.2,11, Thermal-Neutron Flux in Oil in the LTSF Configuration Tank. 151 ANP PROJECT PROGRESS REPORT commercial-grade boron carbide and aluminum sandwiched between aluminum sheets. The sand- wich is usually rolled to a thickness of 7 or 1/8 in. Since the mixture is nonuniform, considerable space exists between the chunks of boron carbide attenuat- ing material, and, consequently, radiation can pene- trate the shield by passing unattenuated between the chunks. Because of this “‘channeling effect,”’ which is a statistical effect caused by the increased transmission along paths which pass through less than the average material thickness, the masses of boral and similar heterogeneous shields must be from a few per cent to several hundred per cent greater than the masses of homogeneous shields which would give the same amount of attenuation. R. R. Coveyou’ has suggested a mode! with which to calculate the approximate transmission of radi- ation through materials that consist of randomly dis- tributed chunks. The material is considered to be divided into layers which have a thickness charac- teristic of the size of the chunks. The holes in the layers are assumed to be located in a manner statistically independent of holes in adjacent layers, so that the over-all transmission is the product of the transmissions of all the layers, As the chunks are made more attenuative, the radiation passing through the holes between the chunks becomes more important, A method which is based on the Coveyou model and which has been extended to include a distri- bution of various chunk sizes and shapes has been developed for calculating the fransmission of radi- ation through heterogeneous shields. The method has been used to compute the transmission of neu- trons through boral. For the calculation it was assumed that the boral sandwich was rolled to a thickness of ]/8 in. and that the thickness of the B,C-Al mixture was 0.085 in. with 40 vol % boron carbide. This resulted in an over-all volume fraction of approximately 25% for the absorbing chunks, which were assumed to be spherical in shape. The chunks were first considered to be of 11 different sizes between 20 and 100 mesh; how- ever, it was found that assuming only four sizes gave approximately the same results, and only four groups were used thereafter, The transmission calculated by this method for normally incident 2200-meter/sec (0.0253-ev) neu- trons through ]/s—in.-'rhick boral was 0.076. This is to be compared with a transmission of 0.0015 70 . . . Private communication, 152 calculated for normally incident 2200-meter/sec neutrons by the homogeneous approximation. However, the homogeneous approximation is an inappropriate first approximation for this type of shield. The transmission of nomally incident neutrons through a ]/B-in.-thick boral shield as a function of energy is shown in Fig. 4,2,12, along with the limit as the chunks become opaque (low energies). The average transmission over the neutron distribution shown (Maxwell-Boltzmann distribution at room temperature) is 0.096 for a constant efficiency de- tector and 0.084 for o 1/v detector. The transmission of isotropically incident neutrons through %-in.-thick boral as a function of energy is shown in Fig. 4.2,13, For this case the average transmissions are 0.024 for a constant efficiency flux detector, 0.021 for a 1/v flux detector, 0.041 for a constant efficiency current detector, and 0.034 for a 1/v current detector. The calculated results can be compared with the results of two experiments which have been per- formed at ORNL to determine the transmission through '/B-in. thicknesses of boral as measured by 1/v detectors. In the first experiment® the radi- ation consisted of thermal neutrons escaping from a thermal column on top of the ORNL Graphite Reactor with an angular distribution of the (1 + /3 cos 6) type,? which is more forwardly peaked than an iso- tropic flux. Consequently, the experimental values should be between the computed values for normal incidence and those for isotropic incidence. The transmission obtained for a Brooks and Perkins boral sample was 0.070, while the transmission for a Carbide sample was 0.094. periment '? the radiation was a collimated beam of normally incident neutrons from a beam hole at the ORNL Graphite Reactor. The transmissions ob- tained for two different Alcoa samples were 0.065 and 0.070, respectively. This method will be described in detail in a separate report, 1 In the second ex- BR. O. Maak, B, E. Prince, and P. C, Rekemeyer, Boral Radiation Attenuation Characteristics, MIT Engineering Practice School, KT-251 {(Nov. 27, 1956). 9R. F. Christy, Lecture Series in Nuclear Physics, MDDC-1175, p 115 (Dec. 1947). ]OG. deSaussure, private communication. W. R. Burrus, Neutron Transmission Through Boral and Similar Heterogeneous Materials Consisting of Randomly Distributed Absorbing Chunks, ORNL-2528 {to be published). PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED ORNL-LR-DWG 25040 1.0 0.5 /’/ v // 0.2 2 P e L1 / z 7 / (.T’.n / /l @ z 0 / E: / /l = -~ g // A e / — HOMOGENEOUS 0.05 P / APPROXIMATION \ LIMIT AS £ —= 0O {OPAQUE CHUNKS) 002 474 =0.0253 ev TYPICAL 20-100 MESH B,C DISTRIBUTION 007 o0d 0.2 0.5 1.0 2.0 50 10 20 40 E/kTy Fig. 4.2.12. Neutron Transmission Through l/a-in.-Thick Boral as a Function of Energy; Normally Incident Flux,. UNCLASSIFIED ORNL-LR-DWG 25044 1.0 #75=0.0253 ev 0.5 | TYPICAL 20-100 MESH B,C DISTRIBUTION "/ A NORMAL INCIDENCE>/ P - 02 —~ // / \\ ‘//' / o z » v / 3 THERMAL NEUTRON 7 - 0 SPECTRUM (PER UNIT LETHARGY) | il Z O FOR 4T =0.0253ev / — A\ —7 z - 3 LA~ P P 4 L \ |~ e [ L ” — (SOTROPIC =Sy 0.05 N / CURRENT DETECTOR —-am, \ L+ ISOTROPIC - p— / ¥ FLUX DETECTOR [ ,/ \ / /,/ // / 1] / /r // / 0.02 7 /‘/ / \\ e / // ‘ faarmaeent"] L1 \ / AT \ O'O-‘ / 007 0d 0.2 05 1.0 20 50 10 20 40 VAT, Fig. 4.2.13. Neutron Transmission Through 1/8-in.-TI1i¢:k Boral as a Function of Energy; Isotropically incident Flux. 153 ANP PROJECT PROGRESS REPORT 4,3, BULK SHIELDING FACILITY F. C. Maienschein Neutron Physics Division THE GE<«BSF STUDY OF HEATING IN SHIELDS K. M. Henry Most shielding research programs, as was pointed out previously, ' are devoted to optimizing shields with respect to radiation attenuvation and to minimiz- ing the size and weight of the shield. This effort tends to increase the power densities, energy deposition rates, and temperatures of the shield. If energy deposition rates could be determined, then temperatures could be calculated and cooling systems designed accordingly. Errors in the size and location of the cooling channels which result either in less than maximum power operation or in a shield that is not optimized in weight or size thus could be avoided. Until the study of radiation heating in shields’ was initiated recently at the BSF in cooperation with the General Electric Company, energy deposition (heating) tests had been performed only on a small scale with samples in geometries that neglected scattering and interface conditions. For example, the energy absorption in small samples of lead, iren, and aluminum, with the calorimeter com- pletely surrounded by water, were reported by Binford.2 In the experiment now in progress an attempt is being made to compare the energy deposition rates found in calorimetric samples in relatively large laminated shields with calculated results. If the calculations can be verified ex- perimentally, then shield designs can be optimized with respect to heating as well as attenuation. The Laboratory participation in the experiment includes providing the facility for the tests and personnel to perform the experiment and to tabulate the raw data. In addition, the BSF is supplying neufron foils and counting channels for gamma-ray and fast-neutron dosimetry. The General Electric Company is furnishing all the shield slabs, the heating samples, and the containing tank with its associated parts. They are also supplying the U233 fission chamber and its electronic counter- parts. In addition, GE personnel will be responsible ‘K. M. Henry, ANP Quar, Prog. Rep. March 31, 1958, ORNL-2517, p 113, 2, T. Binford, Nucleonics 15(3), 93 (1957). 154 for all data analysis, point kemel calculations, data correlation, and preparation of the final report. The necessary reactor flux and power calibration data for the BSR loading being used, as well as the radiation attenuation in water at various distances from the reactor, have been supplied to GE by ORNL. The shield mockup being investigated consists of slabs of beryllium and lithium hydride separated by a gamma-ray shielding section which consists of one of three materials: iron, lead, or Mallory 1000. Measurements of the radiation intensities and temperatures at the slab interfaces in the Be- Fe-LiH configuration were reported previously.’ Subsequently, the entire series of measurements for all three of the configurations has been completed, but the results, including those published pre- viously, ! are of questionable value since it has been discovered that all the heating samples leaked and oil entered the air region surrounding the sample. Consequently, further publication of temperature measurements will be postponed until the series of measurements is repeated with new samples. A detailed description of the experiment and the results of some measurements of radiation intensities are given below. The BSR loading being used in the experiment consists of an array of 28 elements, including two shim-safety rod elements and one regulating rod element. The face adjacent to the shield is 15 in. wide (five elements wide), and 24 in. high; the thickness of the core is either 15 or 18 in., depend- ing on whether the row of elements contains five or six elements. More detailed information on the structure was presented in a previous reporT,3 along with the power and flux distributions, Data on the attenuation of the flux and dose rates from the face of the core through various thicknesses of water were also published in previous reports. 43 3E. B, Johnson, Power Calibration for BSR Loading 33, ORNL CF-57-11-30 (Nov. 28, 1957). 4E. C. Maienschein et al., Attenuation by Water of Radiations from a Swimming Pool Type Reactor, ORNL- 1891 (Sept. 7, 1955). SAttenuation in Water of Radiation from the Bulk Shielding Reactor: Measurements of the Gamma-Ray Dose Rate, Fast-Neutron Dose Rate, and Thermal- Neutron Flux, ORNL.2518 (July 8, 1958). A plan view of the reactor, the shield, and the con- falnmg tank was presented in the previous repor'r and is repeated in this report as Fig. 4,1.11 in Chap. 4.1. The shielding slabs are submerged in oil in a containing tank which is positioned 4.75 in. from the reactor face. (In Fig. 4.1.11 the tank is shown to be only 3.5 in. from the reactor. It had been hoped that it could be positioned this close to the reactor, but for the tests reported here, as well as those reported previously, it was not possible. The recent installation of a new reactor support will allow the closer positioning for future runs.) The containing tank, which was designed to withstand a 5-ft differential water-cil head, was fabricated from ¥-in.-thick aluminum plates rein- forced with 0.25 by 3 by 3 in. channels. It is 4.5 ft wide, 8.5 ft long, and 29 ft deep. After it was positioned in the pool, it was filled with concrete blocks to a height of 5.5 ft. An aluminum platform, which rests on the concrete and is bolted to the tank walls, provides locating pins with which a removable aluminum tray holding the shielding con- figuration can be positioned. Each shielding configuration consists of a 4-in.- thick slab of beryllium, followed by a gamma-ray shield and 16 in. of aluminum-encased lithium hydride. The lithium hydride section consists of two slcbs of 4 and 12 in., respectively, separated by a /8 -in.-thick aluminum spacer which is fitted to the smaller slab in an attempt to reduce the thickness of the ** in the assembled shield. The larger lithium hydride section was supplied by the LTSF and is used in this experiment only to provide an effectively infinite lithium hydride medium behind the smaller section. The gamma-ray shield consists of either 3 in. of lead, 3 in. of iron, or 2 in, of Mallory 1000. A slab of oil extends be- yond the lithium hydride to the tank wall. In ad- dition there is a %-in.-thick layer of oil between the tank wall and the beryllium. (The desired thickness of this oil layer was ¥% in., as indicated in Fig. 4.1.11 in Chap. 4.1). All shields are as- sembled on the pool room floor and installed in the tank as a unit. This simplifies the positioning of the slabs and provides an opportunity for measuring the thickness of cracks between the slabs. The cracks do not tend to change size with handling, as indicated by measurements made before and after the tests, The beryllium slab and each of the gamma-ray shielding slabs have stepped holes along the crack’’ PERIOD ENDING SEPTEMBER 30, 1958 horizontal shield axis in which plugs holding the heating samples are inserted. These slabs alse have l/z-in. deep grooves on one side that extend from the horizontal holes upward to the tops of the slabs. When matched, these grooves form instru- ment wells for vertical flux and dose-rate traverses at the interface of the beryllium and the gamma-ray shield, The small lithium hydride slab has a straight-through hole along its axis, and two instru- ment wells, one near each side, extending from the hole upward to the top of the slab. When vertical traverses are not being made, the instrument wells are filled with plugs of the same material as the slab, the metallic plugs being bare half cylinders and the lithium hydride plugs being full cylinders encased in aluminum. As has been described previously, ' the containers that hold the beryllium, iren, and lead heating samples are each 3 in. in diameter and 2 in. long with an intemal void 2 in. in diameter and 1 in. thick. The design of the sample cases, which were fabricated from the same materials as the samples they hold, is shown in Fig. 4.1.12 of Chap. 4.1. The samples themselves are 1Y in. in diameter and of various thicknesses (the lead is 0.1 in. thick, the iron 0.2 in. thick, and the beryllium 0.4 in. thick); they are suspended by three small insulating support tubes in the center of the void in the container, The case for the Mallory 1000 sample dlffers in thot it is stepped (3 i |n. in diameter for 1Y% in., then 2/ in. in diameter for / in.). The Mallory 1000 sample thickness is 0.1 in. Because of fabrication diffi- culties, the case for the lithium hydride sample was made from 0.25-in.«thick aluminum with external dimensions measuring 2.96 in. in diameter and 1Y, in, in thickness, The lithium hydride sample is 1.25 in. in diameter and 0.5 in, thick. Calibrated iron-constantan thermocouples are attached to the center of each side of the samples and to each in- side face of the cases. The leads are brought through holes in the cases (sealed with Devcon B plastic steel} out to the surface of the oil. Neutron foil activation measurements were made at the interfaces of the Be-Pb-LiH configuration. Cobalt foils were positioned 12 in. apart along the horizontal centerline and 8 in. apart along other radii that extended from the center on 45-deg angles. All the foils were placed in recesses in an aluminum foil holder 4 ft by 4 ft by 0.09 in. thick. The results for both bare and cadmium-covered foils are shown in Figs. 4.3.1 and 4.3,2. 155 ANP PROJECT PROGRESS REPORT ol 2-01-058-0-444 793 x 10? “ 2.02 x 108 (Ca) s 4 “ 177 x 108 (ca) 8.63 X 107 7.24 % 10°? 2.38 x 108 (ca) w \ 8.23 x 108 {Cd) 326 x 10'° Y [ ! Vs 3.02 x 10'° \‘ 845 x 108 (Cd) 9,53 x 10% (ca) ™ 3.28 % 10'° HORIZONTAL CENTER L LINE OI!-_" SHIELD 130 x 102 (cd) 10 8 542 X 10 10 i B ¥ x1o'0 cay > b _ 1.B2 X140 S 1( 153 X 10 (N 5.78 x 108 (cq) AT s \\ = (VA / - P . V- 3.70 x 100 Ly 108 X 10% (Cd) 1.07 x10% (Ca) / 3.90 x 10" v T - / 10 - / ae0x10 W & 927 x10°% (ca) AN \" o) 241x108 (cay /S 1.23 x 100 1.04 x10%0 - 3.46 x 10° (Cd) w L 367 x 108 (Ca) 1.56 % 10'C |« — VERTICAL CENTER LINE POWER: 49.3 kw OF SHIELD (100 kw NOMINAL POWER) Fig. 4'3.10 ments at the Be«Pb Interface (Facing Reactor). Cobalt Foil Neutron Activation Measures The 2- o NN, 1 \ | | . | 7 " + 713 X 107 (Cd) S 1.34 x 10° y ' Ve 1.49 x 10° 792 X107 Cd) 7.38 X 407 (Cd) ‘ 158 X 10° e g o / 4.94 X 10 /\\ . 2.67 x 10% {Ca) / s, / e ,2.54 X 10° (Ca) , 634X 102 /_,- ) 599 x 10° sz 108 (e e / G,,;\ /ON PLUG . 2, NS e 1.61 X 108 [Cd) 327 %107 e / 4.08 X +0° —— P a — - —— e e ) \ 1.88 X 10° (Cd) “ HORIZONTAL CENTER \‘\ LINE OF SHIELD - ~, 2.34 X 10% (Cd) 9.66 X 10° 9.75 x 10° 4.49 x 108 (Cd) 3.80 X 16° (Ca) ) 1.32 x 10'0 | 2.23 x10° ; 140 x 108 (Cd) 144 x 108 (cd ' 279 X 102 = ye . 426 x 0% 1,50 X 109 (Cd) ‘ ee— VERTI CENTER LI POWER: 49.3 kw - oF SH::E'T_"D ENTER LINE {100 kw NOMINAL POWER) 1 Fig. 4.3.2, Cobalt Foil Neutron Activation Meosure- top number refers first run, and the run. The two are of the uncertainty to the observed neutron flux in the lower number to that in the second not necessarily comparable because in positioning between the two runs The in the ments at the PbeLiH Interface (Facing Reactor). top number refers to the observed neutron flux first run, and the lower number to that in the second run, The two are not necessarily comparable because and the steep flux gradient at each interface. Vertical thermal-neutron traverses were made in each instrument well in the Be-Pb-LiH configuration with a 3/4-in.-0D U235 fission chamber. Each L . - counter was attached to a %-in.-dia by 3-ft-long tube which formed part of a four-tube preamplifier housing. A conventional amplifier, scaler, and power supply completed the electronic setup. Difficulty was en- countered during these measurements as a result of the photoneutron background produced by reactor fission-product decay gamma rays. |n addition, the reactor had to be operated at such a low power (to prevent overloading of the counting channel) that 156 of the uncertainty in positioning between the two runs and the steep flux gradient at each interface. these fission-product gamma rays made a linear power response versus demand difficult to achieve. The total error introduced in the results, which are plotted in Fig. 4.3.3, is probably £10%. Fast-neutron measurements were made in the vertical instrument wells with a *‘scaled-up”’ 7gmin.- OD version of the Hurst-type phantom fast-neutron dosimeter. The detector was attached to its pre- amplifier in a manner similar to that of the thermal- neutron detector. Binary-type integrating scalers were used as the summing device. Some pulse- height distributions were measured simultaneously. 6 2-01-058-0-443 10 T T T I — | & O O [¢] 0. T_ 5 ) T 10 3 S HOLE 2 —(Be-Pb INTERFACE)— g Ve v [ = .. -’ o] » - HOLE 2 —--}—e bt (Pb-LiH INTERFACE) ® o 4 S (o) . £ 5 (.J L c . = [e] E L £ 5 3 T 10 % 2 - ) o '_ o] W s I fo o = 102 Co HOLE 3 ] —O {LiH~-LiH INTERFACE) T O O |- ) ~ B @] 0 o a L 10 I Q9 0 4 8 12 16 20 24 28 DISTANCE ABOVE SHIELD AXIS (in.) Fige 4.3.3. Thermal-Neutron Fluxes at the Shield Interfaces as a Function of Distance Above the Shield Axis (GE-BSF Shield Heating Experiment). The effects caused by the fission-product decay gamma rays are believed to have influenced the fast- neutron measurements 5%, These measurements are plotted in Fig. 4.3.4. The ionization chamber used for the gamma-ray vertical traverses was a ‘'scaled-down’’ version of the conventional graphite-walled CO-filled ionization chamber. lts output was read on a direct-reading micromicroammeter. The results are presented in Fig. 4.3.5. Additional information concerning this experiment can be found in several reports published by GE.6=9 6). G. Carver et al., Proceedings of the Fifth Semi- annual ANP Shielding Information Meeting, Atlanta, Ga., May 14~15, 1958, C/25801, vol. |, paper 10. 7). G. Carver, Description and Preanalysis of Proposed BSF Nuclear Heating Measurements, XDC-58-1-51 (Dec. 23, 1957). 8R. 4. Maier, Report of Data from First Nuclear Heating Measurements in Bulk Shielding Facility, XDC-58-8-61 {July 28, 1958). YA, W, Casper, Point Kernel Calculations of Core Gamma Heating in the BSF Nuclear Heating Experiment, XDC-58-8-234 (1958). PERIOD ENDING SEPTEMBER 30, 1958 2-04-058-0-444 10 ° HOLE 1 o (Be-Pb INTERFACE) 5 | £ i T 10 ; U El R= E , =] c W o T HOLE. 2 = (Pb-LiH INTERFACE) $ i (=8 Q = ] w - d T Q g 1 e ¥ 2 ® ' HOLE 3 H @ [ (LiH-LiH INTERFACE) 1 - ? =z o 14 — jun] J ] 3 , b 407 7] CONFIRMING POINTS = E ‘ o 1073 0 4 8 12 16 20 24 28 DISTANCE ABOVE SHIELD AX!S (in.) Fig. 4.3.4. Fast-Neutron Dose Rates at the Shield Interfaces as o Function of Distance Above the Shield Axis (GE ] & [ ] | ® BF; COUNTER + FAST-NEUTRON DOSIMETER 4 4 ANTHRACENE SCINTILLATION DOSIMETER + + .8 27 ft : 33 ft BOTTOM VIEW wtill Fig. 4.4.3. Experimental Arrangement for Total Radiation Measurements in Horizontal Midplane Around the ASTR. 2—01-056-23—-1-758 R{ 2 x 101 : : [ : _—ASTR SIDE TANKS EMPTY ; 10~ A~ KL (CONF 5) N yd V4 Y - y, NG - - |; 5 / \-"\ /’ \ T / / ~ e AN N\ J.: "4 —” N : / N\ c 10_2 y d N E / ASTR SIDE TANKS FILLED N\ W . / WITH PLAIN WATER (CONF 3) \ = | ae \ L 2 L/ \ 5 / UNSHIELDED DETECTOR KEPT AT CONSTANT \ & 1073 |4 DISTANCE OF 59 ft FROM ASTR \ é 4 ALTITUDE = 186 ft ~ = 5 6= 0 deg ALONG REACTOR AXIS (@] 2 1074 0 30 60 90 120 150 180 210 240 270 300 330 360 8, REACTOR ORIENTATION ANGLE (degq) Fig. 4.4.4. Total Gammo-Ray Dose Rates in Horizontal Midplane Around the ASTR. 161 ANP PROJECT PROGRESS REPORT 3 and 5). (The reactor-to-detector separation dis- Fig. 4.4.3, the fast-neutron measurements were tance is measured from the center of the reactor to also made at a point 59 ft from the reactor, but * the center of detection of the counter.) The dose the thermal-neutron flux measurements were made rates at @ = 90 and 270 deg increased approxi- at a point 53 ft away, mately a factor of 4 when the water side shield The angular radiation mappings were supple- . was removed. This increase is only slightly mented by foil exposures, most of which were made greater than would be expected on the basis of on the reactor shield surface. However, foils were the thickness of water removed; however, at 6 =0 also exposed at various positions in space, some deg the dose rate increased a factor of 10. This as far away as 33 ft. The foils used were gold indicates the importance of the side neutron shield- (bare, gold-covered, and cadmium-covered), indium ing in reducing the gamma rays at this point. Cor- {bare and cadmium-covered), copper, aluminum, responding fast-neutron dose rates (Fig. 4.4.5) and sulfur, and magnesium. [n addition, long pieces thermal-neutron fluxes (Fig. 4.4.6) consistently in- of copper wire were exposed near the reactor shield. creased by a factor of 100 at all points when the The results of three such exposures are shown in water side shield was removed. As shown in Fig. 4.4.7. The wires were positioned parallel to FAST-NEUTRON DOSE RATE (mrem-hr-"'-w™) 162 2-01-056-23-776 # ASTR SIDE TANKS EMPTY (CONF 5)——/ N ! t y 4 N 7 AN J N \ - AN / N\ | UNSHIELDED DETECTOR KEPT AT CONSTANT o'l .. DISTANCE OF 59 ft FROM ASTR ALTITUDE =186 ft 8=0deg ALONG REACTOR AXIS | T ASTR SIDE TANKS FILLED WITH PLAIN WATER (CONF 3) L ) 2 i & /fl yd NG 7 A 7 NG y 4 N\ / N pd N ,/ \\\_ / AN // \\\ L N 103 y 0 30 60 90 120 150 180 210 240 270 300 330 360 8, REACTOR ORIENTATION ANGLE (deg) Fig. 4.4.5. Total Fast-Neutron Dose Rates in Horizontal Midpliane Around the ASTR. PERIOD ENDING SEPTEMBER 30, 1958 oy 3 2-01-056—23-777 10 5 [~ ASTR SIDE TANKS EMPTY (CONF 5} . 2 -\ 102 \ T; 5 UNSHIELDED DETECTOR KEPT AT CONSTANT T DISTANCE OF 53 ft FROM ASTR E ALTITUDE = 186 ft » 8 = O deg ALONG REACTOR AXIS [ a 2 £ e o o 10 = . g ASTR SIDE TANKS FILLED WITH PLAIN WATER (CONF 3) 5 5 7 Py =z é / \Z""--. \\ 3 1 1 R I 5 A, 2 - 16" 0 30 60 90 120 150 180 210 240 270 300 330 360 8, REACTOR ORIENTATION ANGLE (deg) Fig. 4.4.6. Total Thermal-Neutron Fluxes in Horizontal Midplane Around the ASTR. the reactor centerline, in the horizontal plane con- taining the reactor axis, at distances of 0, 20, and 40 cm from the shield surface. The results plotted in Fig. 4.4.7 are given as a function of the dis- tance from the front end of the shield. For these exposures the two ASTR side shield tanks were filled with water {(configuration 3). Direct-Beam Doses and Fluxes. — In addition to the total dose-rate and flux measurements made around the reactor, direct-beam neutron measure- ments were made as a function of the angle 6. For the fast-neutron measurements, the dosimeter was placed in a cylindrical water collimator located 12.5 ft above the ground. In this position the center of detection of the dosimeter was in the same horizontal plane as the reactor axis and 33 ft from the center of the reactor. The collimator hole was 3'/4 in. in diameter and approximately 24 in. long and was surrounded by an 18-in. minimum thickness of water. Typical plots of the direct- beam fast-neutron measurements taken while the ASTR was rotated around its vertical axis are shown in Fig. 4.4.8. These measurements were token both with and without the water side shield on the reactor (ASTR configurations 3 and 5). Direct-beam thermal-neutron flux measurements were made at a reactor-detector separation dis- tance of 15 ft and an altitude of 186 ft. For these measurements the BF3 counter was positioned in a cylindrical, paraffin collimator which was lined 163 ANP P 5 10 - : e e _ «tn | : : \ r o ; » : ‘ | J 5 9 I _ c : 4 g ; - 20 cm FROM w2 ON SURFACE SURFACE | 5 Q L = 10?4 > i : : ; i ~ N Ny = : - . 8 5 -/- 1 \.\5\ H '_ - . : .\\\ - — 2 40 ¢m FROM , z / SURFACE : R — ) | ) . 1\,_, . = /. | \ N\ = ‘ ; ! i AN, I 3 i ' : i = 10 . i ! ! \ L — = 4 1 . | | \ - A7 I | L. y : ! 1 fi 5 - ',/" | i i [¥E] @ - ASTR SIDE TANKS FILLED WITH PLAIN | WATER; BORAL PLATE REMOVED FROM ---- SHIELLD SURFACE (CONF. 3NB) | ! ’ , ] {0 I ‘ | 0 12 24 36 48 60 72 84 96 Fig. 4.4.7. Relative Thermal-Neutron Fluxes Alang Side of ASTR Shield: Copper Wire Activation Measure- ments on Shield Surface and 20 and 40 c¢cm from Surface. ROJECT PROGRESS REPORT " 2-0t-056-23-778 with cadmium. This collimator hole was 2 in. in diameter and was surrounded by a 6-in. minimum thickness of paraffin; the distance from the out- side end of the collimator hole to the center of detection was approximately 12 in. The collimator was suspended from the support truss, and a re- motely positioned cadmium shutter allowed both bare and cadmium-covered measurements to be made. Examples of these measurements are given in Fig. 4.4.9 for ASTR configurations 3 and 5. Scattered Doses. — Measurements of the scattered fast-neutron and gamma-ray dose rates were also made as a function of 8 for various ASTR shield configurations at distances of 33, 59, and 100 ft, although none are reported in this paper. For these measurements, which were made both at an altitude of 12.5 ft and at an altitude of 186 ft, the detectors DISTANCE FROM FRONT END OF ASTR SHIELD (in.) were suspended from the support truss in a line which, when 6 = 0 deg, was coincident with the reactor axis. The arrangement is shown in Fig. 4.4.10. The details of the direct-beam shields are shown in Fig. 4.4.11, These shields were remotely positioned to vary the direct-beam shield angle a from 8 deg to approximately 30 deg and thus to eliminate various solid angles of radiation while et 2-01-056-23-1-779 2 it \\ — — . t — — -~ Y 7 4 N - _ - -~ b 1 5 P il A / N/ U — / \‘ ‘ l RS, L AN L ol - “—F— ASTR SIDE TANKS \ / \ s o EMPTY (CONF 5) — N = / fiomne Ao . | - ¥ ) T . T T A F" \ i - ‘\ I T < I S I \ ] L R ; — \ / <\ E / \ ~ o \ } N \ I b i — 7 NN '<_[ ~ Y 4 l[ J N B .S D o i /< / 4 _ 7~7\7Wi T A\fi " . p— 8 / / / % h o T (] 2 - 4 ; r \ i ] 2 / K ‘ ; S 41073 |- , .1 ™ASTR SIDE TANKS FILLED J \ = 7 WITH PLAIN WATER (CONF 3) h Y o yd - A Y w3 7 I I I l N il . | l \ | AN g 2 ~{ DETECTOR IN COLLIMATOR KEPT AT CONSTANT CISTANCE OF 33 ft FROM ASTR N\ 104 / ALTITUDE = 12.5 ft \\ ———1 § =0 deg ALONG REACTOR AXIS - — . — 2 — — 1073 0 20 40 60 80O 100 120 140 160 180 200 220 240 260 280 300 320 340 360 164 8, REACTOR ORIENTATION ANGLE {(deq) Fig. 4.4.8. Direct-Beam Fast-Neutron Dose Rates in the Horizontal Midplane Around the ASTR. PERIOD ENDING SEPTEMBER 30, 1958 2—O|—056—2!-|-—757 T - I - I | yd S~ ASTR SIDE TANKS EMPTY ___/ N\ N (CONF 5) \ / ,,a ] T ..F L Y A \ Y . \ / \ \ \ / \ o/ A \/ \ THERMAL-NEUTRON FLUX (counts-min~t-w™1) 7 i A\ \ v 4 7 N \ L | y i 7 N\ | _——ASTR SIDE TANKS FILLED \ \ 5 LA V4 " WITH PLAIN WATER CONF 3} A \ A // C 7 \ N\ , / \\ 1 1 1/ \v/\ /\/ \ — ¥ 4 \fv N 7 LY y 4 AY 5 P N— DETECTOR IN COLLIMATOR LI S / KEPT AT CONSTANT DISTANCE \ / OF 15 ft FROM REACTOR » ALTITUDE = 186 ft — / 6 = O deg ALONG REACTOR AXIS | | | | 0 30 60 30 120 150 180 210 240 270 300 330 360 8, REACTOR ORIENTATION ANGLE {deg) Fig. 4.4.9. Direct-Beam Thermal-Neutron Fluxes in the Horizontal Midplane Around the ASTR. UNCLASSIFIED 2-01-056-23-783 DIRECT BEAM SHIELDS ANTHRACENE SCINTILLATION DOSIMETER FAST-NEUTRON DOSIMETER BFy COUNTER o+t H A4 m9!> { Fig. 4.4.10. Experimental Arrangement for Scattered Radiation Measurements in Horizontal Midplane Around the ASTR. 165 ANP PROJECT PROGRESS REPORT the reactor-to-deteator separation distances re- mained fixed. By taking many measurements in this manner, it was possible to extrapolate the plots of dose rate versus a to obtain the scattered dose rates for a = 0. By subtracting the resulting values from the total dose rates obtained at the same separation distances, a second method for determining direct-beam dose rates was available, UNCLASSIFIED 2-01-056-23-784 + FAST-NEUTRON DOSIMETER A ANTHRACENE SCINTILLATION DOSIMETER DISTANCE VARIABLE DISTANCE | FIXED a &) Pb 8in. f ‘\ BORATED RUBBER Fig. 4.4.11. Direct-Beam Shield Used Between the ASTR and Fast-Neutron and Gamma-Ray Detectors. Measurements in Crew Shield Mockups In the NTA experiments the ASTR and its shadow shield were located approximately midway between the nose and the tail of the airplane. A cylindrical crew shield mockup was located 33 ft in front of the reactor in a position that was midway between the reactor and the crew compartment in the nose of the airplane. The total distance between the reactor and the crew compartment was 65 ft, In the TSF experiments, measurements were made both in and around the crew compartment and the crew shield mockup, both of which were suspended from the support truss, as shown in Fig. 4.4.12; however, the typical data presented in this paper for this set of measurements were all obtained in the crew shield mockup. The walls of the crew shield mockup were fabricated from borated rubber; the thickness at the rear was 21 in., and the side and front thicknesses were varied from 0 to 11 in. (see Fig. 4.4.13). The 30-in.-dia by 70-in.-long cavity in the mockup was lined with 2.5 in. of lead on the rear and with 0.093 in. of lead on the side and front. In addition to the measurements in the Convair shields, measurements were also made in an ORNL UNCLASSIFIED 2-04-056-23-790 @® LOCATIONS FOR DETECTORS QUTSIDE CREW SHIELDS {(ONE FAST-NEUTRON DOSIMETER, ONE ANTHRACENE SCINTILLATION DOSIMETER, AND TWO BFy DETECTORS) CONVAIR CREW COMPARTMENT i OR ORNL-TSF COMPART- MENTALIZED TANK CONVAIR CREW SHIELD MOCKUP |l bl i | il o JiE l ’lH“ »“:F ’1[‘ ‘ - i7" N-REACTOR ‘ ROTATOR LEAD J SHADOW SHIELD\L . ™asTR r 33 ft =‘ Fig. 4.4.12. Configuration of ASTR and Crew Shields. 166 comparLTIe‘Q!t’c:Iized tank that had been used for earlier experiments at the TSF. The compartments in this tank, shown in Fig. 4.4.14, were filled either with plain or with borated water. The rear shield was 37 in, thick, and the side shield was varied from O to 20 in. When this tank was used it replaced the Convair crew compartment. 2-01?6—23-789 INSIDE DIMENSIONS 30-in. DIA x 70-in.LONG V/A BORATED RUBBER (PERMANENT)}, 21-in. THICK BORATED RUBBER, VARIABLE THICKNESS FROM O TO il in. B Lc20 (PERMANENT), 0.093-in. THICK ON SIDE AND FRONT, 2.5 in. ON REAR DETECTOR STATIONS 251, 252 DETECTOR STATIONS 243A, 2438 DETECTOR STATIONS 24¢, 242 FAST-NEUTRON DOSIMETER ANTHRACENE SCINTILLATION DOSIMETER CADMIUM-COVERED BF, COUNTER BARE BF ; COUNTER oepp Fig. 4.4.13. Convair Cylindrical Crew Shield Mockup. PERIOD ENDING SEPTEMBER 30, 1958 Effect of_Reiecior Side Shield Thickness. — The oo measurements of radiation in and around the Convair crew shield mockup 33 ft from the ASTR were taken as a function of the thickness of the borated rubber side shield on the crew shield (de- tector) for most of the ASTR shield configurations at altitudes of 12.5 and 186 ft. Typical gamma-ray dose rates, fast-neutron dose rates, and thermal- neutron fluxes at the 186-ft altitude are shown in Figs. 4.4.15, 4.4.16, and 4.4.17, respectively, for configurations 3 and 5. The measurements in the ORNL compartmental- ized tank 65 ft from the ASTR as a function of the water side shield thickness were made only at the 186-ft altitude. Three reactor shield=detector shield combinations were used: (1) Plain water shields both on the reactor (configuration 3) and on the detector tank; (2} a plain water shield on the reactor and a borated water shield on the de- tector tank; and (3) borated water shields both on the reactor (configuration 3B) and on the detector tank. Typical measurements in this tank are also plotted in Figs. 4,4,15 through 4.4,17. Effect of the Ground. - In order to determine the effect of the ground on the measurements made in and around the Convair crew compartment and UNCLASSIFIED 2—-01-056—15—P-490 e - 131.60 in. OUTSIDE e 1 ~—-—37.05 in. QUTSIDE —-—‘ o~ 73.68in OUTSIDE—- -~ -——— —=Ji= 20,49 in OUTSIDE = i I 12.0in. (TYP) g 1] 8k . t = = |2.|2f|"l. { | =~ £ - ML 2.0 in T F g-— < [ ot o = e -73.48 in: e o @ (f\(p). t m l - | 8 " | } py Lol e L R _ = ) 1.90 in. - = = 10 413211 36.0in.DIA (TYP) —s le S D D D D oo ‘ NI f o~ = S— * ? N 5 } J ’ { 17.74 in. DIA— o 6 : 18,0 in. OD 7 2.0 in. (Typ)-A 1" 8 —={ 11.90in. =— 9 -Hz‘sean‘—-—t —= 24.06in. (TYP) |- - 7.96 in.—=] - ) ' 0.375in. Fig. 4.4.14. ORNL-TSF Cylindrical Compartmentalized Tank. (@immgt with caption) 167 ANP PROJECT PROGRESS REPORT —1 2-01-056~23-1-756 10 e e R R ALTITUDE =186 ft B — " | CONVAIR MOCKUP: 33 ft FROM REACTOR | ORNL TANK: 65 ft FROM REACTOR —2—~ASTR SIDE TANKS EMPTY (CONF 5) | L N N N A S S — 10 ' A\ i + — | = - ’IN CONVAIR MOCKUP WITH — ] 7/‘ BORATED-RUBBER SIDE SHIELD ] | _ASTR SIDE TANKS | 7 FILLED WITH PLAIN 3 |- WATER (CONF 3} ] GAMMA-RAY DOSE RATE (mrem-hr™!-w \Q " IN ORNL TANK WITH | §§\ 107 E=——— P AIN WATER SIDE SHIELD =~ +— S "IN ORNL TANK WITH —— i = BORATED-WATER SIDE SHIELD=""T T * ] | ASTR SIDE TANKS FILLED | A WITH BORATED-WATER (CONF 3B) 0 4 8 12 16 20 DETECTOR SHIELD SIDE THICKNESS (in.) Fig. 4.4.15. Gamma-Ray Dose Rates in the Convair Crew Shield Mockup and in the ORNL-TSF Compartmen- talized Tank as a Function of the Detector Shield Side Thickness: ASTR Experiment. crew-shield mockup, many of the measurements were repeated as a function of altitude. Typical gamma-ray measurements made in air and in the crew-shield mockup 33 ft from the reactor are shown in Fig. 4.4.18 for two ASTR shield con- figurations: configuration 5, in which there was no side shielding on the reactor, and configuration 5.5, in which only the lower halves of the two ASTR side tanks were filled, Corresponding fast- neutron measurements are given in Figs. 4.4.19 and 4.4.20. Thermal-neutron flux measurements made in air for ASTR configuration 5 and three reactor-to-detector separation distances, 27, 53, and 94 ft, are presented in Fig. 4,4.21. e EAEY 168 T 2-01-056-23-1-755R4 =1 2 X140 1 T [ T T T T 1 40_‘ \\ 9=Ode9 — e ALTITUDE =186 ft — S CONVAIR MOCKUP 33 ft FROM REACTOR— \\ ORNL TANK 65 ft FROM REACTOR ] N N S N SR B " \\ASTR SIDE TANKS EMPTY (CONE 5) 10 = k. N i T N o = £ 103 N { \ [}} N 1 £ S +; = =N CONVAIR MOCKUP WITH = \< BORATED RUBBER SIDE SHIELD o A | | 1 i | l w HEEEE w404 9 X ASTR SIDE TANKS FILLED —— - N WITH PLAIN WATER (CONF. 3) —] S AN = N D s N Z 10 |_ [75] & AN ™ IN ORNL TANK WITH N ~ BORATED WATER SIDE AN 156 | SHIELD: SAME CURVE WAS N = OBTAINED WHEN ASTR SIDE TANKS = — =~ WERE FILLED WITH PLAIN WATER N ~ (CONF. 3) OR WITH BORATED WATER {CONF. 3B) g7 0 4 8 12 16 20 24 DETECTOR SHIELD SIDE THICKNESS {in.) Fig. 4.4.16. Fast-Neutron Dose Rates in the Convair Crew Shield Mockup and in the ORNL-TSF Compartmen- talized Tank as o Function of the Detector Shield Side Thickness: ASTR Experiment. Gamma-Ray and Neutron Spectral Measurements In addition to the measurements discussed above, which essentially repeated those taken in the NTA experiments at Convair, the TSF experiments in- cluded a series of measurements at an altitude of 186 ft to determine the gamma-ray and neutron energy spectra at various locations. Gamma-ray spectral data were obtained in the Convair crew shield mockup for the 11-in.-thick side shield and in the ORNL compartmentalized tank for several thicknesses of borated water on the side. For these measurements two sodium iodide crystals were used with an RCL 256-channel pulse-height 2-01-056-23-1-7 10 o | I 1 | I 1 | 1 Al 8=0deg p— ALTITUDE =186 ft — CONVAIR MOCKUP 33 ft FROM REACTOR | ORNL TANK 65 ff FROM REACTOR 2 ¥\ 10 A\ LY \ \ = \ASTR SIDE TANKS EMPTY (CONF. 5) i 10 \ * N\ 'c A E @ X 2 1 r\ A\ RE] 3 2N LY 7 LY § /1' A Y | \ / - \ ( N = e \ |\ IN CONVAIR MOCKUP WITH 5 10 \——= BORATED RUBBER SIDE SHIELD -— w N ¢ : ’ — W \ ASTR SIDE TANKS N X7 FILLED WITH PLAIN g N\_| || WATER (CONF 3) £, \\ \\ IN ORNL TANK WITH T 10 \. PLAIN WATER SIDE = . N S SHIELD - ™ ~ ™ IN ORNL TANK WITH N "~ BORATED WATER SIDE AN 10~ | SHIELD: SAME CURVE = WAS OBTAINED WHEN = = ASTR SIDE TANKS WERE —X " FILLED WITH PLAIN WATER _ (CONF. 3) OR WITH BORATED \\ _4 | WATER (CONF. 38) 40 4 I 1 | 1 1 1 N 0 4 8 12 16 20 24 DETECTOR SHIELD SIDE THICKNESS (in.) Fig. 4.4.17. Thermal-Neutron Fluxes in the Convair Crew Shield Mockup and in the ORNL-TSF Compartmen- talized Tank as o Function of the Detector Shield Side Thickness: ASTR Experiment. analyzer. Typical pulse-height spectra for four side shield thicknesses on the crew shield (de- tector) and ASTR configuration 5 (reactor side tanks empty) are shown in Fig. 4.4.22. A pulse- height spectrum for a 20-in.-thick detector side shield is shown in Fig. 4.4.23 for ASTR configura- tion 3 (both reactor side tanks filled with plain water), A spectrum for the same detector shield thickness but with only one of the two ASTR side shields filled with water (configuration 15) is pre- sented in Fig. 4.4.24. Neutron spectral measurements were made in a 6-ft-dia, 6-ft-long cylindrical water tank which was suspended from the support truss 65 ft from the reactor and had eight conically shaped collimator PERIOD ENDING SEPTEMBER 30, 1958 - 2-04-056-23-1-780 10 5 oy /‘\\ \ T \\dEASUREMENTS IN AIR 2 e — £ \ — ® ' 162 < ¥ . Ty | ] £ I MEASUREMENTS IN CONVAIR ] s i MOCKUP WITH NO SIDE SHIELD.] m > [ ASTRSIDE TANKS/ —~— > — EMPTY (CONF. 5) 3 A i ‘—-.N S ~ ) > ~~ MEASUREMENTS IN CONVAIR P MOCKUP WITH 11—in.~THICK ! BORATED RUBBER SIDE \\ SHIELD = 7 Ry / g / 7/ — © [ 5 |_“LOWER HALVES OF | ASTR SIDE TANKS FILLED WITH PLAIN WATER (CONF. 5.5) 2 REACTOR-DETECTOR SEPARATION DISTANGE = 33 fi 8= 0deg 0 25 50 75 100 125 150 {75 ALTITUDE (ft) Fig. 4.4.18. Gamma-Ray Dose Rates in Air and in the Convair Crew Shield Mockup as a Function of Altitude: ASTR Experiment. ot 2-01-056-23-4-7864 i MEASUREMENTS IN AIR — ASTR SIDE TANKS EMPTY (CONF. 5) ; I MEASUREMENTS IN CONVAIR /NOCKUF’ WITH NO SIDE SHIELD \ < FAST-NEUTRCN DOSE RATE (mrem-hr '-w '~ REACTOR-DETECTOR SEPARATION " DISTANCE =33 ft T 8=0deg I} 5X40 2 ' ! ‘ 0 20 60 S0 120 150 180 ALTITUDE (ft) Fig. 4.4.19. Fast-Neutron Dose Rates in the Convair Crew Shield Mockup with No Side Shielding as a Function of Altitude: ASTR Experiment. 169 ANP PROJECT PROGRESS REPORT 2 — 8=0deg ASTR-CREW SHIELD SEPARATION DISTANCE = 33 ¢ FYZRE L] -3 2-01-056-23—-1-754Ri 10 T I . / ASTR SIDE TANKS EMPTY {CONF 5} 'z / \ T ]/ \ £ ] 5 E ) 2 o Ll " 8 --_-—'_'—-_~ [m] =z LOWER HALVES OF ASTR SIDE TANKS 8 FILLED WITH PLAIN WATER {CONF 5.5) 5 V8] T — w & 0 30 60 90 120 150 180 ALTITUDE (f1) Fig. 4.4.20. Fast-Neutron Dose Rate in the Convair Crew Shield Mockup with 11ein.-Thick Borated Rubber Side Shield as a Function of Altitude: ASTR Experiment. holes radiating from near its center. The collima- tor axes were spaced at 45-deg angles, each col- limator half-angle covering 5 deg. Neutron-sensi- tive plgtes were placed at the inner ends of the collimator holes, which were sealed off from one another. The results of all of these measurements will be reported by Convair, Miscellaneous Measurements During the course of the experiment it became apparent that several measurements were needed which were not originally scheduled. These in- cluded resonance and threshold detector measure- ments as a function of distance from the reactor and sulfur-foil measurements on the ASTR shield 170 2-01-056-23-1-782 10" | ! \ 6=0 deg ] = N\ ASTR SIDE TANKS EMPTY = N\ (CONF. 5) , g T ! ES T > \ c c AN » 27 ft FROM ASTR £ \ N g \ T > = | w 2 ™, = S ~ Ns ft FROM ASTR i— - ,_éJ \ ] J 402 N\94 ftFROMASTR | | 2—, where D = gamma-ray or fast-neutron dose rate (mrehr=tow=1 or mrep-hr"]-w"), C = arbitrary constant, r = reactor-detector separation distance (ft), A =relaxation length = 785 ft for gamma rays = 630 ft for fast neutrons. UNCLASSIFIED 2-01-056-10-3-268R{ 10 I ) T ) | I T ! ] I — | | | & —— a2 HURST-TYPE DOSIMETER; REACTOR % |L VISIBLE | PORTABLE DOSIMETER (DIRECTIONAL); | 2 4 REACTOR VISIBLE _l,_¥ ‘A PORTABLE DOSIMETER {DIRECTIONAL); 3 ’ REACTOR NOT VISIBLE 10 A - i _— P N Y I N i | L N } 5 T f 7 N\ A=630 ft N T e “ i T A2 ; = 10 T ] \\\ ] . i o N - E S N N\Q A\ S N 2 LN \.\ I N\ I _ AN AN 5 \\ 2 q 0 800 1600 2400 3200 4000 4800 r, DISTANCE FROM REACTOR TO DETECTOR (ft) Fig. 4.4.27. Measured Fast-Neutron Dose Rates as a Function of the Distance from the TSR-l with a 15-cm- Thick Water Shield. PERIOD ENDING SEPTEMBER 30, 1958 Predicted Dose Rates 4200 ft+ from the TSR-ii Gamma-Ray Dose Rates for Bare TSR-Il. — The gamma-ray data presented in Fig. 4,4.26 were used to predict the gamma-ray dose rates from the bare TSR-{l simply by accounting for the differences between the thicknesses of the shielding materials on the two reactors along a straight line from the reactor to the detector. While the TSR-| had a 15-cm-thick water shield, the TSR-Il shield will consist of 2.5 cm of water, 3.5 cm of aluminum, and 5.1 cm of lead. The modified expression is therefore (2 Doy(TsR-1t,bare) = -1/ A k4 ”(EH2O"H2O+2AI"AI+EP b*P b) C - P e e 1 r2 —r/A (3) ¢, -2 T2 f where D., = gamma-ray dose rate (mr-hr™1), C,=172x 104 me-hr= 't 20w =1 (a constant taken from ref 3), P = reactor power level (w), x = thickness of a shield material on the TSR-I! minus the thickness of the same material on the TSR-1{, =2.5¢cm = 15 em = =12.5 cm, xH2O cm cm cm Xpp =35 cm, xp = 5.1 cm, 2 = macroscopic absorption cross section for gamma rays in the specified material, -1 E,0 =0.0396 cm™", 2, =0.0953 cm=!, 26y, =0.467 cm™ . With these values of x and 3 substituted in Eq. 2 and with the assumed power of 5 Mw, the value of C, in Eq. 3 becomes 9.35 x 10° mrehe=1-ft2. The complete expression for the gamma-ray dose rates at various distances from the bare TSR-|l operating at a power of 5 Mw.is thus -r/785 9 e () DoirsR-it,bare) = 935 % 10" — r mrehr=1 This expression is plotted as D,yr2 versus rin Fig. 4.4,.28 175 ANP PROJECT PROGRESS REPORT Fast-Neutron Dose Rates for Bare TSR-11 (Based UNCLASSIFIED on ASTR Data). — Rather than attempt to predict 10° £ p1-0s6-2o- A TeR fast-neutron dose rates for the bare TSR-Il from 5 the available TSR-| data, it was considered pre- ferable to obtain measurements from a reactor \ shield that more nearly resembled that to be used R . \ on the bare TSR-1l. For this reason, measurements " 0 \\ were taken at a distance of 4200 ft from the Convair -,T \\ — — Aircraft Shield Test Reactor (ASTR) (see preceding £ 3 \\ L section of this chapter) with its thinnest shield 5 N ] configuration (see configuration § in Fig. 4.4.2). £ 2 \(’FAST NEUTRONS This shield consisted of 15 ¢m of water, 4.45 cm qu‘o - of aluminum, and 7.3 ¢m of lead. It should be : e e e pointed out, however, that the ASTR is approxi- 2 > \ AN mately cylindrical, and the shield thickness given ; \\\ \\ is in the perpendicular plane bisecting the longi- 2 2 \ \ ’ tudinal axis of the reactor, X 10 . \ = N\ AN The point at which the fast-neutron dose rate B N, was measured was 4200 ft from the side of the " S \\\ \\\ ASTR along a line perpendicular to the longitudinal 2 N axis of the reactor and originating at the reactor N N center. The angle between this line to the detec- ;» 108 < tor and the reactor axis is identified in the discus- B sion below as the angle 6. With the reactor operat- > GAMMA RAYS - \\\ ing at a power of 1 Mw, the measured dose rate was AN 0.78 mrem/hr (ref 4). As a check on this measure- e N ment, a comparison was made using data taken at o N various distances from the Nuclear Test Airplane 0 {000 2000 3000 4000 5000 6000 while it was in flight at altitudes greater than r, DISTANCE FROM REACTOR TO DETECTOR (ft) 10,000 ft (ref 5). An extrapolation of these data to ground level resulted in a fast-neutron dose rate Fig, 4.4.28. Predicted Maximum Values of D, r2 and of 0.77 mrem-hr"l, which seemed to confirm the Dn r2 as a Function of the Distance from the Bare TSR-II measurement at the TSF. Operating at a Power of 5 Mw. The TSF measurement 4200 ft from the ASTR was then used to calculate the dose rate from the bare TSR-1l by normalizing Eq. 1 to the measured value, accounting for the differences in the shielding on the two reactors, accounting for the differences in the neutron fluxes on the surfaces of the two reactors, and assuming that the fast-neutron dose rates 4200 ft from the reactors would be proportional to their total fast- neutron leakages. The expression for the dose rate is thus - (EHzotzo"'zAlel"'EP b*P b) G 22 p, , Gl ® © (3) D (TSR-11,bare) = Ca 3 3 1 where D, = fast-neutron dose rate (mremehr=1), C3 = a constant with units of mrem-hr=1ft2.w™1, At the same point the gamma-ray dose rate was 0.55 mr/hr. 55. C. Dominey and J. B. Moore, Airborne Radiation Mapping Data, NARF-58-14T (March 14, 1958). 176 PERIOD ENDING SEPTEMBER 30, 1958 (,‘152/(;51 =ratio of the neutron flux at the surface of the TSR-II to that at the surface of the ASTR, tzo =2,5em ~15em =~12.5 ¢m, %, =3.5¢cm -4.45cm =-0.95 cm, Xpp =91 ecm =73 ecm =-2.2 cm, 2 = macroscopic cross section for neutrons in the specified material, Iy o =0.1538 em™!, EAI = 0,079 em™T, S5, =0117 em!, G, = geometric factor for ASTR, which is the ratio of the average fast-neutron dose rate at a given distance to that at 6 = 90 deg for the same distance D(6) " D(6 =90 deg) ] .[;" D(6) sin 0 6 D(6 = 90 deg) - f sin 0 d0 0 ] n - D(6) sin 6 dO, 2D(6 = 90 deg) -l; () sin D(6) = dose rate at a constant distance from the reactor as a function of the angle 8 (taken from the top curve in Fig. 4.4.5 of this chapter), G, = geometry factor for the TSR-I, which, because of the spherical symmetry of the reactor, is equal to unity, The final value of Gz/G] = 1/0.599. If it is assumed that the power distributions in the ASTR and the TSR-Il are flat, the flux is inversely proportional to the volumes of the cores and the term v, /V2 can be substituted for (,'152/(;5], where vV and V, represent the volumes of the ASTR and the TSR-Il, respectively. Equation 5 then becomes —7/A e n v, G, ‘(szoxH20+EAIxAI+ZPbeb) (6) Dn(TSR-lI,bare) = C3 T2 _V_ G_ Pe 2 Uy —r/)\n (=4 (7) =0y Assuming a reactor power of 5Mw and reducingEq.6 to Eq. 7 gives a value of 6,40 x 10'! mremhr= 112 for the constant C4+ The complete expression for the fast-neutron dose rate from the bare TSR-II operating at a power of 5 Mw is thus ~7/630 1 -1 (8) Do (TSR-11,barey = 6-40 < 10 —z mrem-hr 177 ANP PROJECT PROGRESS REPORT This is plotted as Dflr2 versus r in Fig. 4.4.28, When the value of ris 4200 ft, D_ is 46 mremshr~ ', FastsNeutron and GammasRay Dose Rates for TSRall in a Beam Shield. — No experimental information was available on which to base the calculation for the highly collimated beam; therefore, it became neces- sary to perform an experiment with the TSR-1 in which a 15-in.-dia air-filled collimator was placed adja- cent to the reactor in its 12-ft-dia water tank, Measurements of fast-neutron and gamma-ray dose rates and thermal-neutron fluxes were made at a distance of 4200 ft as a function of the beam orientation angle, 6, with the detectors positioned at § = 0 deg (6 deg west of due south). These measurements are plotted in Fig. 4.4.29. The fast-neutron and gamma-ray measurements had a very high background, so the statistical accuracy is poor. At the best point, 6 = 0 deg, the probable errors total 32% for the gamma-ray measure- ments and 8% for the fast-neutron measurements. The statistical accuracy of the thermal-neutron data is much better, however; the probable errors range from 7.4% at 6 = 130 deg to 2.4% at 0 = 0 deg. As a result, the thermal-neutron data were used to obtain an assumed shape for the other measurements, The measurements obtained for fast-neutron and gamma-ray dose rates at 0 = 0 deg were used to predict dose rates from the TSR-ll with a similar collimator by accounting for the differences in the shielding ma- terials on the two reactors along the beam axes. The shield thicknesses used for the TSR-1l were the same as those given above for the bare reactor. The shield thicknesses on the TSR-l were reduced to 1.9 cm of water and 3.0 cm of aluminum, The fast-neutron dose rate at 4200 ft from the TSR-Il operating at 5 Mw in a beam shield thus becomes o) 5 _p , e'(EHZOxH20+EAIxAI+EPbeb) n{TSR-1l,beam) n(TSR-l,beam) where, for this case, xH2O =25em-19cm=0.6cm, xpp=35¢em - 3.0ecm=0.5¢cm, Xp) = 5.1 em, and the values of X are the same as those given following Eq. 5. The resulting value of D (TSR-11, beam) 4200 ft from the reactor for a power of 5 Mw is 1,07 mrem+hr=1. A similar procedure gave a value of 0.23 mrshr=! for the gamma-ray dose rate, Do (1sR-11, beam) The total dose rate (neutrons and gamma rays) from the highly collimated beam shield was thus 1.3 mrem«hr—7, 82% of which was due to fast neutrons. Figure 4.4.30 shows these predicted dose rates as a function of distance from the reactor. These dose rates apply only for a point directly in line with the beam, and they can be reduced by about a factor of 6 by turning the beam 90 deg. FasteNeutron Dose Rates for the Bare TSR«ll (Based on TSRe«l Data). — The fast-neutron collimator dose rates presented in Fig. 4.4.29 were used to make an additional prediction of the fast-neutron dose rates at large distances from the bare TSR-Il. The calculational procedure was as follows. If ¢, represents the neutron flux on the surface of the TSR-| and it is assumed that the flux has a cosine current distribu- tion, the net number of neutrons crossing 1 cm? of the TSR-| face per second per steradian in the direction a{a being the angle from the normal) is /277) cos a neutronsscm™2.sec” ' .steradian~'. If it is further g 1 178 PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED UNCLASSIFIED _ 2-04-056-25-1-752 04 - —26-5-794 9o 2X|O3 2-04-056-25 2x10? AR o \X 10? s AR RN W . ,__,_/, A N\ N\ 5 X\ L> FAST NEUTRONS 5 / N ‘.\‘ \ \ d \\—D— <\THERMAL-NEUTRON FLUX 2 2 A\ 1 ~n \ ] N -4 i GAMMA-RAY DOSE RATE (mr-hr w ' sec w o ~6 P 3 'g 10 - - 10 L » 07 A 7 _ = \ P ; G - = \ £ 3 s 0 E 3 \ > € < [ f/f"\\ w \ E ST T Y N B /AR 5 \ - g 1 ! ® 2 2 W R e 7N\ . \ s £ 5 / 9 4o \ oF 3 £ = > A 3 - % = —T < Q £ 2 /| \ £ i\ 3 £ 30’ y i e ! A\Y z g & - +— e D W a 5 \L="CAMMA RAYS 5 & w g a7 AN 2 @ 2 = 7 X = A\N 5 = oo . | - . i \ ) : s hd o \\ E: g /\ 2 2 — i = 7 FAST-NEUTRON DOSE RATE \ \ = W . / Lo | \ b : . { AN\ o & 5 // | | A b 2 2 15+in.~dic COLLIMATOR ADJACENT TO REACTOR —— \\ ~ IN TANK 8 = HORIZONTAL ANGLE BETWEEN COLLIMATOR 5 \\ \\ 5 AXIS AND REACTOR-DETECTOR AXIS N -8 | | | i | \ 1) -150 -120 -30 =60 -30 0 30 6C \ 8, BEAM ORIENTATION ANGLE {deg) 2 N \\ 2 \ \ 10! Ay Fig. 4.4.29. Measured Fast-Neutron Dose Rates, 0 1000 2000 3000 4000 5000 Gamma-Roy Dose Rates, and Thermal-Neutron Fluxes DISTANCE FROM REACTOR (ft) 4200 ft from the TSR-1 as o Function of the Beam Angle & Beam Emitted Through Air-Filled Collimator Pene- Fig. 4.4.30. Predicted Gamma-Ray and Fast-Neutron trating Reactar Shield. Dose Rates as a Function of the Distance from the TSR-II Operating at a Power af 5 Mw in a Beam Shield. assumed that there is no wall scattering, then the total number of neutrons leaving the collimator of length L and radius a is (10) ? cosal A AQ = (;;cos a) (7a?) where A = area of the collimator, AQ) = solid angle subtended by the end of the collimator, as seen from the reactor, cos a =approximately 1, since a is small for L >> a. The dose rate at an angle 6 from the collimator due to 1 neutron per second emitted from the collimator is: 2712 " P 179 ANP PROJECT PROGRESS REPORT where D(60) is the measured dose rate 4200 ft from the TSR-l as a function of the angle 6. Then the total dose rate from a unit isotropic source (a source emitting 1/47 neutrons+sec™!.steradian=) is: D(6) / 2mL? L2 - (12) K == ) 27 sin 6 d6 = — D(6) sin 6 40, y 411\ (17a”) ma'p, J8=0 where 277 sin 6 40 is the differential solid angle. It ¢, represents the neutron flux on the surface of the TSR-Il and it is assumed that it also has a cosine current distribution, the total number of neutrons emitted per second from the TSR-I| would be v 13) wr [ (22 2 si aniig, [ i z ( T Tcosa (27 sin a da) = 77”]9/)2 fa:() Cos a sin ada=277r]q.‘>2 ' a=0 where 7, is the outer radius of the core; the other terms have been defined previously. {t is now possible to calculate the dose rate at 4200 ft from the bare TSR-II by multiplying Eq. 12 by Eq. 13 and introducing an attenuation term: -(ZH,0%H,0" A1 A" EP L P e L2 T (14) D, 1l bare) = 2t {—— f D(6) sin 0 46 (TSR-11l,bare) 172 7Ta4q')] o 2.2 ~ 5 2L /Y, m ‘(\ZHzotzOJ"EAIxAIJrEPbeb) . (15) S (s f D(6) sin 048 | e : \V2 6=0 where the values of £ and x are the same as those given following Eq. 9, and, as in Eq. 6, it is assumed that the flux is inversely proportional to the volumes of the cores. (In this case, however, V, is equal to the volume of the TSR-l.) The resulting dose rate for a 5-Mw operating power is = (szoxH20+zAlel+zP b*P b) e e b 2(48)2(14.75)2 (6,480) n(TSR-1l,bare) — (75)4 (10,640) f'" D(6) sin 0 d6 F==0 =193 [(1.79 x 10-7 mrem-hr‘l-w"])(S x 108 w)] ]— 2.07 = 83 mrem-hr ™! This value compares reasonably well with the value of 46 mrem:hr=! calculated above. As a check on the validity of this type of calculation, a similar procedure was used to predict the fast- neutron dose rates 4200 ft from the ASTR which could be compared with the measured dose rate reported above for the same reactor shield configuration. The calculated result was 2.78 mrem+hr=! and differed from the measured value by a factor of 3.56, the latter value being 0.78 mrem+hr=1, As a further check, a calculation of the fast-neutron dose rate 4200 ft from the TSR-I in its 12-ft-dia tank gave a predicted value . which was a factor of 1.8 higher than the measurement. These two calculations indicate that the value of 83 mremshr~! given above is probably a factor of 2 or 3 too high. The discrepancies probably arise partly - because the number of neutrons which leave the collimator is greater than Eq. 9 predicts and partly because 180 PERIOD ENDING SEPTEMBER 30, 1958 the beam results for fast neutrons may not vary with angle in exactly the same manner as the results for thermal neutrons (which were used to give the curve shape for these calculations). In any event, it appears that the dose rate at 4200 ft which was calculated by this method should be conservative in the sense that the calculated dose rate should be larger than the actual dose rate. Calculation of Uncollided Fast-Neutron Flux 4200 ft from the TSR-| A calculation of the uncollided neutron flux at a distance of 4200 ft from the TSR-| was performed to obtain a fast-neutron buildup factor for the case in which a beam is pointed at the detector. For this calculation the measured neutron spectrum, N(E), at the face of the Bulk Shielding Reactor® was used for neutron energies up to 2.7 Mev and a normalized Watt's spectrum was used for higher energies, where N(E) is given in units of neutrons.cm™%.sec™'.w™='.Mev~.steradian='. If a cosine distribution on the face of the reactor is assumed, the uncollided fast-neutron dose rate at a distance ! is: o ~2.(E)I 17 N(E) cos a C(E) AQ e (17) D, U(TSR-LI) = f y dA_ dE , E=0 AS det where A_ = area of the source (the reactor face), dA_ = differential area on the reactor face, A, ,, = area of detector, ET(E) = macroscopic total cross section for air, C(E) = flux-to-dose rate conversion factor, AQ) = solid angle subtended by the detector at dA_, ! = distance from dA_ to the detector. In order to base the calculation of D, u(tsr.y ©n @ measured fast-neutron dose rate at the surface of the reactor, it was necessary to calculate the dose rate at the surface of the reactor as well as at the distance /. For this case, the distance ¢, defined in Fig. 4.4.31, approaches zero and e—ET = L (18) D Im | f N(E) C(E) <—A"e*> 2mx dx | dE TSR-1,surface) ~ 7 n(TSR-lsurface) E=0 A_ Ager \! 12 4] [ (¢ 0] tx dx =2 I N(E) C(E) f - E=0 | Yx= ! 2 Im N(E) C(E) — fm | e = 27T — o | Jo (x2 + t2)3/2 e 4] —t 03] = 27 N(E) C(E dE o (E) C( (x2 + t2)]/2 =27 f N(E) C(E) 4dE , E=0 ®R. G. Cochran et al., ANP Quar. Prog. Rep. Dec. 10, 1953, ORNL-1649, p 117. 181 ANP PROJECT PROGRESS REPORT where x is the distance from the horizontal centerline of the reactor to the area dA. When the detector is at large distances from the reactor, / >> x, cos ¢ is equal to 1, and the integral over the source simply gives the area of the collimator: I Ao [ ~Z (E) ! ‘(2H2o"H20+2A|"A|) (19) D uctsra,n = 5 N(E) C(E) e dE e [ E=0 &Y -(“'H2o"H20+EA1"A|) where the exponential e accounts for the attenuation of the 1.9 cm of water and 3.0 cm of aluminum between the reactor and collimator. The buildup factor, B, is: 2 fE zo N(E) C(E) dE A ® “3. (E) coll f N(E) C(E) e T dE 12 E=0 (20) B = Dmecs,l=4200 ft meas,surface by < ‘(“Hzo"H2oT~A|"A|) e (442 x 10-7) (27)(3.65 x 10°) ] (2,05 x 10%) (6.83 x 10~8)(16.62) (0.59) =7.38 < —2..(E)I Figure 4.4.32 is a plot of N(E) C(E) e T as a function of the neutron energy E for two separate calculations. In the first calculation the averaged cross sections were taken every 0.5-Mev energy interval. These cross sections are given in Table 4.4.1. For the second calculation the cross sections used were taken from BNL-325 (ref 7) at 20-kev intervals below 2 Mev, at 100-kev intervals from 2.0 to 6.0 Mev, at 200-kev intervals from 6.0 to 8.0 Mev, and at 0.5-Mev intervals from 8.0 to 12.0 Mev. The resulting inte- grations under these two curves differed by less than 10%. ’D. 4 Hughes and J. A, Harvey, Neuiron Cross Sections, BNL-325 (July 1, 1955). UNCLASSIFIED 2-01-056-25-D-795 - | DETECTOR - # - | Fige 4.4.31. Geometry Used for the Calculation of the Uncollided Flux 4200 ft from the TSR-l, 182 UNCLASSIFIED 2-01-056-25-A-753R 2 o' _a—— USING AVERAGE TOTAL CROSS ____| ] SECTIONS AT O.f-Mev INTERVALS ——] I ® \ 1 _ i '\ w 5, % \ N o r ‘ Q s 1 \ 2 0 } 8 40 i Wi/ N\ o I X f \\ W _L 5 l \ W 1 Ay A\ »E | IS N | v \\ S, ] . g ! ; NN\ = | \ \ - 10 I —t \\ LY I | LU T v N AY Y 5 AW USING AVERAGE TOTAL CROSS LN | SECTIONS AT 0.5-Mev INTERVALS/ \ , )\ 10 0 2 4 6 8 10 i2 ENERGY (Mev) Fig. 4.4.32. Calculated Uncollided Fast-Neutron Spec- trum 4200 ft from the TSR-l for the Case in Which an Air Filled Collimator Penetrates the Reactor Shield and Points at the Detector, Table 4.4.1. Average Cross Sections Used for Calculation of Uncollided Flux PERIOD ENDING SEPTEMBER 30, 1958 E, Neutron S Air E, Neutron S Air Energy T'_] Energy T'_] (Mev) (em™") (Mev) {em™") 0.5 13.7 x 10~3 6.5 6.59 x 10~ 1.0 10.6 x 10~3 7.0 6.79 x 1073 1.5 9.23 x 103 7.5 6.79 x 10~3 2.0 8.19 x 10~ 8.0 6.59 x 10~ 2.5 7.18 x 10~2 8.5 6.59 x 10~ 3.0 8.06 x 1073 9.0 6.59 x 10> 3.5 8.50 x 10™° 9.5 6.59 x 10> 4.0 8.50 x 1077 10.0 6.59 x 10~ 4.5 7.09 x 10> 10.5 6.59 x 10~° 5.0 7.04x 1077 11.0 6.59 x 10~° 5.5 7.24 x 1073 1.5 6.59 x 10~° 6.0 6.74 x 1073 12.0 6.59 x 107> 183 ANP PROJECT PROGRESS REPORT 4,5. TOWER SHIELDING REACTOR Il C. E. Clifford Design work on the Tower Shielding Reactor || (TSR-Il) is now essentially complete. The several design changes which have been incorporated since the previous report! and recent studies of operating conditions and experimental facilities are described below. The two shields which will be used initially with the TSR-|| are also discussed. MECHANICAL DESIGN? The latest concept of the TSR-Il is shown in Fig. 4.5.1. The design and much of the fabrication of the major components inside the reactor tank have been completed, although the fuel elements will be modified following full-scale flow tests, The present design differs from the one last reported in several respects. The inside dioameter of the upper portion of the reactor tank has been increased from 36.75 to 39.75 in. to facilitate removal and in- sertion of the fuel elements and the shielding above the core. In addition, the ionization chambers have been moved outward in the upper region of the inner cylinder to provide more room for the control mechanism positioning and actuating devices which will be housed in a turret atop the central cylinder region, Another modification in the design has been the elimination of one of the boral hemispherical shells just outside the lead region. This will ensure adequate clearances in the annular fuel region. A 2-ft-thick shield of lead shot and water has been added above the central fuel elements to reduce the gamma-ray leakage through the central cylinder. This will be especially important when the reactor is enclosed in a shield mockup. This lead-and- water shield, which can be seen in Fig. 4.5.1, ]C. E. Clifford and L. B. Holland, ANP Quar. Prog, Rep. Dec. 31, 1957, ORNL-2440, p 275. 2The mechanical design work on the TSR-ll is being done by the Engineering Department, 184 L. B. Holland will be penetrated by approximately 120 helical ducts through which the cooling water will flow. The gamma-ray leakage through this shield will never exceed the leakage through any shield used with the TSR-ll. The shield shown above the core in the region between the reactor tank and the inner cylinder will be removable. The design of the TSR-Il is such that it can be encased in various shields. The first shield with which the reactor will be used will be a spherical shield having a beam hole both fore and aft (see discussion of shield designs in subsequent section of this chapter). The entire assembly will be sus- pended from a platform, as shown in Fig. 4.5.2, and it can be rotated around the vertical axis so that the beam holes can sweep a horizontal plane. When the reactor is suspended in this manner, the cooling water will pass through a coaxial swivel joint be- fore it enters and after it leaves the reactor tank. The reactor and beam shield may also be sus- pended so that the beam holes will sweep a vertical plane, For this case, two 12-ft girder-like members will be attached to the side of the shield through an axle and bearings, and the cooling water will enter through a swivel fitting and the hollow axle of one support member and leave through a similar arrange- ment on the other support member. By alternately using the two beam holes, a beam can sweep the whole 360 deg in the vertical plane without tilting the reactor more than 90 deg from the vertical position, Calculations have indicated that, with a 30-deg conical section removed from the shield, a 6775 ft-lb torque will be required to rotate the reactor and shield so that the removed section will sweep a vertical arc. A much smaller torque will be re- quired to sweep the horizontal arc. A Char-Lynn torque booster working through a worm and 180- tooth gear will be used to rotate the reactor as- sembly. A torque booster under full load has per- formed satisfactorily in laboratory tests, CONTROL TURRET ~—a DRIVE GEAR N ROTATION BEARING ‘i‘ 7 o s I! 7 NE CENTRAL CYLINDER #i§ ASSEMBLY 95} a1 oI SHIELD SUPPORT A SLEEVE 0 FISSION CHAMBER —-3 HELICAL COOLING-WATER TUBES —&f{ INCHES LEAD-WATER SHIELDING — WATER INLET : - I_"_ nl 1 —— N IRt A - ] - 5% 4 Pt i Zfi N — o o o © - il ou. e? © a% o g0 2 PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED ORNL-LR-DWG 32708R1{ WATER OUTLET PLATFORM o= 7 L 5 s s, SN | I, ¥— IONIZATION 8f1 4in. M CHAMBER 6ft 7in. y e~ ALUMINUM REACTOR TANK " |8 PERMANENT N LEAD-WATER SHIELDING CORE REGION TP T TT LT - ,. [ 22777700 CONTROL MECHANISM AND SUPPORT ASSEMBLY Fig. 4.5.1. TSR-ll Design (Vertical Section). 185 ANP PROJECT PROGRESS REPORT o i v TR N A ‘ ’W _, ! { i [ i, CONTROL AND INSTRUMENT LINES DRIVE MECHANISM AND SAFETY MECHANISM OPERATORS BEAM HOLE - COAXIAL SWIVEL JOINT UNCLASSIFIED 2-01-060-368 # ——— SUPPORTING PLATFORM FOR HORIZONTAL ROTATION SUPPORTING FLANGE FOR VERTICAL RCTATION Fig. 4.5.2. TSR-il in Beam Shield. CONTROL SYSTEMS? Control Mechanism The experimental mockup of the control mechanism for the TSR-II, which was described previously, ' was tested and found to have a short life owing to excessive wear of the traveling nut of the position- ing device, Several modifications of the design failed to increase the life. Design changes were 186 then mode so that the pressure on the traveling nut was reduced by nearly a factor of 10. This was accomplished by reducing the size of the seal that restricts the flow out of the chamber and increasing the area on which pressure is applied to seal the control plate assembly against the traveling nut. 3The controls system for the TSR-ll is being develop- ed by the Reactor Controls Group. The force on the traveling nut was reduced to 10 |b, the difference between the spring force and the hy- draulic force. Proper guidance and longer life were ensured by using stellite bearings instead of minia- ture roller bearings on the control plate assembly and by using piston rings to provide a good seal between the assembly and the cylinder wall. With these changes, a model of the control mechanism (see Fig. 4,5.3) was still operating successfully when a test was terminated after 5000 cycles. The model is shown in the test stand in Fig. 4.5.4. The model will be tested later under flow conditions. A total of five control mechanisms will be used to operate neutron absorbing grids for shim-safety pur- poses in the TSR-Il. A sixth grid, which does not scram, will be used for fine control. The neutron- absorbing umbrella-shaped grids will be made of Inconel-clad cadmium ribbons. It had previously been planned? to use perforated boron-loaded plates as the neutron absorbers byt the fabrication of the Inconel-clad cadmium was more feasible. 4C. E. Clifford and L. B. Holland, ANP Quar. Prog. Rep. June 30, 1957, ORNL-2340, p 323. %“ CONTROL PLATE PERIOD ENDING SEPTEMBER 30, 1958 Electrohydraulic Transducer for Control of Water Pressure in the Control Mechanism J. E. Marks? An electrohydraulic transducer will be used to actuate the control mechanism. The purpose of the transducer is to allow electrical control of the water pressure in the TSR-Il control mechanism with the standard ORNL magnet circuitry. The transducer is a bypass pressure regulator with electrical rather than mechanical adjustment. As a force-balancing device it compares the control mechanism water pressure acting on a diaphram with the thrust of a solenoid and applies the difference between the two forces to a spring. Any differential force causes the spring to deflect proportionately, and the deflection is used to displace a valve plug away from or toward its seat, depending upon whether the water pressure effect is greater or less than the thrust of the solenoid. This valve is in a shunt relation to the control mechanism. Slnstrumenfuficm and Controls Division. UNCLASSIFIED 2-01-060-378 TSN SRR RSN SIRSRNLX 7777777%é%/// 7T 77l // | 00 WATER FLOW PN g oo ] T \ S N S \ \ 1IN OVERTRAVEL 58 LTI LL VAV AP AN NN RN N CLEANING RING/ \DRIVE NUT Fig. 4.5.3. TSR.ll Control Mechanism. 187 ANP PROJECT PROGRESS REPORT UNCLASSIFIED Bl P0T0 43086 Fig. 4.5.4. TSR-1l Control Mechanism in Test Stand. Between the source of water and the control mechanism~transducer load, there is a restriction of known magnitude. The load pressure is princi- pally a sum of the source pressure less the pressure drop across this restriction, and the pressure drop is in turn proportional to the square of the total flow through both devices. Hence the transducer can affect the pressure seen by the control mech- anism. 188 The electrical input to the solenoid is from the circuitry of the standard magnet amplifier and sigma bus. One design model of the transducer, shown in Fig. 4.5.5, has been built and tested. With an in- put of 100 to 0 ma, the control range was from the operating pressure P to (P — 20) psig. This range was sufficient to regulate the load pressure between the scram condition and the stable operat- ing condition of all the control mechanisms tried. The load pressure followed a scram signal input after a delay of 20 msec. This was adjudged as satisfactory; however, a later design incorporating several improvements is to be tested soon. The advantages of the transducer over an “‘on- off’" type of device, such as a solenoid valve, are the same as those of magnet amplifiers in com- parison with relays in conventional safety magnet confrol circuits. Furthermore, the fransducer will supply a regulated source of water at any required pressure to improve stability and reliability of the control mechanism operation. TSR-Il Block Diagram L. C. Oakes® The logic of the TSR-II block diagram (shown in Fig. 4.5.6) falls into the general pattern of that of other ORNL reactors and therefore it is not dis- cussed in detail here. However, many small dif- ferences brought about by special features of the TSR-1I have altered the specific modes of operation, and these differences are described below. Grid Motion. ~ Since the TSR-I| will have one drive unit common to all five of the shim-safety grids, it will not be possible to move one grid in- dependently of the others. Thus any operation of the drive is analogous to ‘‘group operation” in the conventional reactor. For orderly operation, a means for producing more modest changes in reactivity must be available to the operator. For this operation, a slow insertion and slow withdrawal mode is provided wherein the drive runs at approxi- mately 20% of full speed. All grid withdrawal is stopped by 10-sec periods, and the grid drive switching is so arranged that a 25-sec period will stop fast withdrawal and automatically limit shim motion to the slow-speed mode. Avutomatic Shim Insertion. ~ Although automatic shim insertion has been used previously, the scheme presented here is aslight departure from the normal concept. As shown in the block diagram, when the servo is on and the regulating grid reaches PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED 2-01-060-38 FEEDBACK ZERO ADJUSTMENT, DIAPHRAGM SPRING AND — CENTERING DIAPHRAGM SOLENQID ORIFICE A A A 4 N N 4 N N 000G N N % . A — D)ISCHARGE % \'§ AN ) 3 in 1} — A Wi HHl” —— INLET N AN - / i__ — Fig. 4.5.5. Electrohydraulic Transducer for Control of Water Pressure in TSR-Il Control Mechanism. the lower limit, the shims will be inserted at the slower speed until the regulating grid withdraws to the intermediate limit switch. This action takes place only if the period is more positive than -100 sec. There are two normal conditions which will cause the regulating grid to drive to the lower limit during operation. One will occur when, for some reason, such as the reduction of the water inlet temperature, the reactivity is increasing and the regulating grid drives to the lower limit while holding a constant power level. Under this con- dition, since the reactor period is infinite, the shims would be expected to drive in to re-establish the regulating grid in the center of its stroke. The other condition will occur when a demand for re- duced power causes the regulating grid to drive to the lower {imit. In the latter case, however, the period is likely to be shorter than ~100 sec, and automatic shim insertion will be inhibited. In this condition a substantial shim insertion would drive the reactor far subcritical, and upon reaching the desired lower power level the regulating grid would be driven to the upper limit with the reactor still subcritical. The 100-sec period interlock eliminates the need for reshiming after a reduction in power level. Fission Chamber Drive. —~ An automatic mode of operation is provided for the fission chamber which, at request, will hold the counting rate between 2 counts/sec and 10° counts/sec. This mode of operation relieves the operator of the task of keep- ing the counting channel on scale and contributes to an orderly operating system. |n addition, since the pulse preamplifier is contained in the same can as that used to contain the fission chamber and sees the same flux, the automatic mode of operation en- sures that the electronic components will not be unduly irradiated. When the automatic mode of operation has not been requested, an annunciator warns the operator when the counting rate has ex- ceeded 10° counts/sec. Test and Automatic Rundown Inhibit (TARDI). - TARDI permits periodic testing of shim drive mechanism performance without the encumberance of certain operating interlocks. |t may be used only so long as all seat switches indicate that the grids are seated, the shim pump pressure is zero, and the key switch is off. Door Interlocks. = At the request of the operator, the front door interlock may be temporarily blocked to permit personnel to enter and leave the control house at shift change without shutting down the reactor. Permission to bypass the front door in- terlock constitutes a mode of operation and is 189 ANP PROJECT PROGRESS REPORT UNCLASSIFIED ORHL-LR-QOWG 32707R{ | | INSERT - ’{ OFF }‘ ‘F|SSIONCH.I CONTROL POWER WITHORAW FISSION CH. AUTOMATIC | MAIN PUMP FLOW WATER FLOW SEAL SEAL >200 gpm BLOCK OUT >100 cps l < 100 ¢ps NO REVERSE log # OR CRM CONFIDENCE MAIN PUMP FLOW J WATER FLOW SERVO "ON" TARD! "OFF" T MORE POS. THAN =100 sec TARDI "ON" REG. GRID IN OR [PS:"D /| ABOVE INT. LIMIT // >200gpm BLOCK OUT rd SEAL l_—., Lnu. GRIDS SEATED | | "NO" SAFETY TROUBLE i START-UP PERMIT STOP SEAL PR R > AST | RTJ Wl RTJ REG. GRID " TARDI "ON | “F ST INSE | [ SLOW INSE IN LOWER LIMIT _PERIOD > 10 sec | DOCRS ANDO HOIST SEAL NO FRONT DOOR BY-PASS | | ND INSERT NO WITHDRAW INTERLOCK PERMIT I FRONT DOOR MON!TRON—‘ REAR ODOOR MONITRON } ORIVE NOT AT DRIVE NOT AT UPPER LimIT LOWER LtMIT I INSERT NO CLUTCH WATER FLOW Pee20ib BLOCK OUT FRONT DOOR BY-PASS SAFETY TROUBLE INTERLOCK PERMIT PERIOD < 5 sec I ’ 2 CHANNELS FAILED SAFETY LEVEL] To>180°F * < {Omr/hr < {0 mr/hr FISSION CHAMBER NOT MOVING { SERVO ERROR | M WDR. REG. GRID | | I WOR. GRID | PUMP FLOW l LOW EMERGENCY I SERVO ON J log &/ LEVEL REVERSE ] >{0 X SERVO DEMAND >4.2 Wy { PERIOD > 25 sec | I PERIOD < 25 sec_l SCRAM RESET s NO WATER FLOW log # > 10 BLOCK OUT ISERVO 0FF1 | SERVO ON } ’ L NO INSERT] I NO INSERT j GRID DRIVE NOT AT UPRER LIMIT | GRID DRIVE NOT AT UPPER LIMIT GRID DRIVE NOT GRID DRIVE NOT AT UPPER LIMIT AT LOWER LIMIT S GRIO DRIVE NOT AT LOWER LIMIT REG. GRID INSERT GRID DRIVE NOT AT LOWER LIMIT FISSION FISSION CHAMBER REG. GRID WITHDRAW SLOW GRID WITHDRAW FAST GRID WITHDRAW CHAMBER INSERT WOR. INSERT INSERT HOIST PCWER WATER FLOW BLOGK OUT CONTINUOUS BELL Sw. rcHANGEOVER l L PUMP DELAY s EXP. OR OPER. DISABLE FRONT BY-PASS TARDI = TEST AND AUTOMATIC RUN DOWN INHIBITOR HOIST MOTION REQUEST DOOR BY-PASS "ON" REACTOR REACTOR I 5-sec DELAY }-—— IN POCL OUT OF POOL DEADMAN SW. — 3-min DELAY log # <107 ANY GRID OUT OF SEAT | ; ‘—i 5-sec DELAY START-UP DELAY FRONT NC FRONT DOOR REACTOR OCOR HOIST CONTINUOUS MOTION START-UP BvePASS BY-PASS B PERMIT BELL PERMIT PERMIT INTERLOCK INTERLOCK PERMIT PERMIT Fig. 4-506- Block Dingram of fhe TSR'IIO 190 described as follows: When the log N indicates that the power level is below 500 w the operator may request that the front door interlock be by- passed. While the bypass is in effect the following conditions prevail: 1. Personnel may enter and leave the front door without producing a scram. 2. No shim grid withdrawal is permissible. 3. The servo will remain in operation to counter- act reactivity changes resulting from such things as reduction in the temperature of the cooling water. 4. Any rise in power above 500 w will produce a reverse. At the request of the operator, normal operating conditions may be restored. Before permission is granted to withdraw shims, however, the horn which is always sounded prior to reactor operation at the TSF must again be sounded for 3 min to warn per- sonnel who may have been left outside the building. Water System Interlocks. — There are two modes of operation of the reactor from the standpoint of the water system: the power mode and the test mode. With the power mode the reactor is operated at or near full power, and the flow is varied between 200 and 1000 gpm over the power range. With the test mode the reactor is operated at a power level sufficiently low to prevent overheating with no water flow. This is brought about by a test switch designated throughout the block diagram as ‘‘water test.”’ Prior to a startup for a power run the main pump must be operated at the full flow of 1000 gpm for 3 min in order to produce sufficient mixing of the water to remove any severe temperature gradients from the system. This pump delay may run con- currently with the startup delay. Failure of the main pump will produce a scram rather than a re- verse, The more drastic measure is preferred in order to lessen the afterheat probiem, The small volume of water in the reactor vessel and the exposure of the vessel and interconnecting hoses to the cold weather bring about special re- quirements with respect to afterheat and winter freezes. A highly reliable turbine pump which will deliver between 10 and 40 gpm to the reactor through the main pump lines will be operated 24 hr a day. Since this emergency pump must remove the after- heat in the event of main pump failure, it will be operated from an a-¢ to d-¢ converter and a bank of 60 station batteries capable of running the pump at full delivery for 8 hr or more. If a failure of the emergency pump should occur, the reactor will be PERIOD ENDING SEPTEMBER 30, 1958 reversed to shutdown level. An electric heater in the emergency pump line will maintain a safe temper- ature when the system is unattended in the winter. Reactor Controls Cubicle The reactor controls cubicle has been delivered and set up at the Tower Shielding Facility. The cubicle, which is 14 ft long, 7 ft 5 in. high, and 7 ft deep, is shown in Fig. 4.5.7. It is designed for unattended operation and is patterned after the cubicles for the Bulk Shielding Reactor and the Pool Critical Facility. The cubicle will be partially wired and then moved into its final position when the con- trol console and vertical board for the Tower Shielding Reactor | are removed. HEAT REMOVAL EQUIPMENT The portion of the 5-Mw heat removal system which is being constructed by an outside contractor should be completed in October., The only holdup at present is the delivery of one pump to the contractor. The system will be subjected to a hydrostatic test with the 6-in. lines to the tower legs capped off. A flow test of the system will be made by short circuiting the 6-in. pipes to the tower legs with a 6-in. hose. In order to provide for the event that the studies by Gambill and Hoffman (see below) indicate that a Reynolds number of 7000 is necessary, an in- vestigation is being made of the cost of raising the cooling water flow rate or raising the system static pressure or both to achieve the desired Reynolds number for turbulent flow. NUCLEAR CALCULATIONS M. E. LaVerne Nuclear calculations which were performed for the TSR-1l with the 3G3R Oracle code were reported previously. '+6=7 A multigroup, multiregion reactor code® has now become available, and an adaptation of this code to a Goertzel-Selengut type has been used to perform a number of further calculations for the TSR-|l, as discussed below. 8C. E. Clifford and L. B. Holland, ANP Quar, Prog, Rep. March 31, 1957, ORNL-2274, p 294. 7C. E. Clifford and L. B. Holland, ANP Quar. Prog. Rep. Sept, 30, 1957, ORNL.-2387, p 304. Bw. E. Kinney, R. R, Coveyou, and J. G. Sullivan, Neutron Phbys, Ann. Prog. Rep. Sept, 1, 1958, ORNL.- 2609, p 84, 191 ANP PROJECT PROGRESS REPORT T cLassiFieD | PHOTO 44sds 2 STSS Fig. 4.5.7. TSR-1l Reactor Controls Cubicle. Comparison of Calculations and Critical Experiments In order to confirm the validity of this calcu- lational method and also to determine the effect of a lead-boral layer outside the fuel region, calcu- lations were first performed for three critical ex- periments which were set up in finite slab geometry as shown in Fig. 4.5.8 (the figure shows only one- half of each configuration). Each fuel region con- sisted of a 4 by 6 array of standard BSR fuel ele- ments, and boral sheets simulated the control plates. Clean, or nearly clean, experiments® were obtained for configurations A and B by heating the water in the assembly tank and simultaneously 9These critical experiments were performed by D. F. Cronin, J. K. Fox, and L. W. Gilley. 192 withdrawing a cadmium rod to maintain the assembly critical. On configuration B, a portion of the rod (2.75¢ worth) was still inserted when the rod travel was used up. Configuration C was set up to evaluate the effect on the reactivity of removing the lead and boral layers. Because estimates of excess reactivity at room temperature for both configurations B and C were available from rod calibrations, configuration C was not heated to attain the ‘‘clean’’ state. A comparison between the results of experiments and the calculations is presented in Table 4.5.1. The error for the three cases averages about 1%. The experimental & is the multiplication constant that would have been observed with the rod com- pletely withdrawn, i.e., in the “‘clean’’ condition. Previously computed temperature coefficients were PERIOD ENDING SEPTEMBER 30, 1958 Table 4.5.1. Comparison Between Calculated and Experimental Values of Multiplication Constant Water Correction to T Multiplication Constant, k Calculated k& Corrected Error Confi ti emperature nrguration o Experimental Calculated . Thermal Calculated & (%) ("F) Density Base A 108 1.0000 0.9897 0.0014 0.9911 0.9 B 118.9 1.0002 0.9872 0.0018 0.989%90 1.1 C 73.4 1.0081 0.9921 0.0035 0.0002 0.9958 1.2 soamre o ez -10 | k = I J zm™ ! o [ Jwarer w E 15 N & L3 Q| y. I 8 b - =20 8[ Y T w 8 g al \ / I _1 | 7 o% -es| &l — E 0 q w | o > | I CONFIGURATION A 30 | | CONFIGURATION B i CONFIGURATION C [ ] . 1 ; | L 1 1 ] L | J 0 10 20 30 40 50 60 {cm) Fig. 4.5.8. Configuration for TSR-1l Critical Experi- ments in Slab Geometry. used to correct the calculated & for differences be- tween the experimental and the calculational thermal base and dens ity. Reactivity Coefficients Void coefficients for the core were computed from the change in the multiplication constant produced by deleting the water (but not the aluminum or uranium) from thin spherical annuli in the core. The results are shown as the points in Fig. 4.5.9. A “teasonable’’ curve was drawn through the calcu- lated points and volume-averaged to obtain a mean void coefficient of ~2.15 x 1074 Ak/k (in per cent) 20 25 30 35 40 REACTOR RADIUS (cm) Fig. 4.5.9. Core Void Coefficient as a Function of the Reactor Radius. per cubic centimeter of void, This is to be com- pared with a value of =3.5 x 10=# which was reported for SPERT |.19 Temperature coefficients were calculated both for density and for thermal base changes. The net temperature coefficient was =0,77 x 102 Ak/k (in per cent) per °C, which can be compared with a valve of -0.9 x 10~ 2 for SPERT 1.0 For various reasons (for example, cumulative tolerances in fabrication), core boundaries may de- viate from the desired locations, An estimate of the effect of such deviations on the reactivity was made by deleting uranium from a thin spherical annulus at the outer edge of the core while the total amount of uranium was kept constant. The radial coefficient was found to be 0.23 Ak/k (in per cent) per centimeter, These reactivity coefficients are summarized in Table 4.5.2, in which a mass coefficient is also given. 10F . Schroeder et al., Nuclear Sci, and Eng, 2, 96 (1957). 193 ANP PROJECT PROGRESS REPORT Table 4.5.2, Summary of Computed Reactivity Coefficients for the TSR-!l Average void coefficient for core, -2.15 % 10-4 Ak/k (in per cent) per cubic centimeter of vaid Temperature coefficients, Ak/k (in per cent) per °C ~1.39 x 10=2 Density coefficient +0.62 x 10=2 Thermal base coefficient Net coefficient -0.77 x 10-2 Radial coefficient, Ak/k (in per +0.23 cent) per centimeter Mass coefficient, (Ak/k)/(Am/m) 0.30 Control *‘Shell’’ Effectiveness The control grid effectiveness was recalculated as a function of the separation distance from the fuel, with the contral grids being treated simply as thin regions or shells. The results are shown in Fig. 4.5.10. The shape of the curve is in good agreement with the results of the earlier 3G3R Oracle calculations, ! although it is a factor of 5.2 higher. The apparent discrepancy is due primarily to the difference in absorption cross sections used for the shells. The experimental data shown in the figure are from the Bulk Shielding Reactor test with eight solid stainless steel plates, which was described previously.!! The BSR data are normalized to the calculated values at the first experimental point. CONTROL GRID EFFECTIVENESS The nuclear model used for the calculations de- scribed above is a complete shell which is assumed to move uniformly away from the fuel region, and therefore it does not duplicate the control grids which will be used in the TSR-ll. The actual con- trol and safety grids will cover less than one-half of the core-reflector interface, and, because of the fixed radius of the plates, the edges of the plates e E. Clifford and L. B. Holland, ANP Quar. Prog, Rep. Junme 30, 1957, ORNL-2340, Fig. 5.3.6, p 327. 194 UNCLASSIFIED 100 2-01-060-54 ] ] I ] ——— | ——= ORACLE CALCULATIONS: 46-cm-dio INTERNAL WATER REFLECTOR, —] 50 14 - ¢cm - thick CORE, 2-cm-~thick LEAD, L 0635-cm-thick BORAL, 20 - cm-thick EXTERNAL|WATER REFLECTOR | | T e I B _~ EIGHT 77-mil-thick STAINLESS STEEL PLATES X 20 IN LOADING 628, FOUR CENTER ELEMENTS, x VALUES NORMALIZED AT FIRST EXPERIMENTAL < \ POINT. . . > 4 [ \ o> i 7o) " o b N < *o\ g ] [ — . ”" L T 5 . \\ N 2 0 q 2 3 4 5 6 7 SEPARATION DISTANCE (cm) Fig. 4.5.10, Calculated Effect of a Boral Shell in the Internal Water Reflector on the Reactivity as a Function of the Distance Between the Shell and the Core: Com- parison with BSR Data. will move away from the core at a slower rate than will the center of the plate. In order to determine the worth of an actual grid, the grid area was divided into rings with their common center at the center of the grid, and the worth of each ring was determined as a function of its normal distance from the core. The total worth of the grid was then defermined by summing the worth of all areas of each grid for a given separation distance between the core and the center of the plate. The results are presented in Fig. 4.5.11. The Inconel-clad cadmium ribbons in the regulating grid and in the lower shim-safety grid will be spaced ]’46 in. apart to permit water to flow through the internal reflector region; therefore, these plates are worth only one-half as much in reactivity as the remaining shim-safety grids. After the control grid worth was determined, the thermal-neufron flux distribution throughout the reactor (see Fig. 4.5.12) and the source distribution in the core (see Fig. 4.5.13) were determined from the multigroup calculations with the control grid po- sitioned for operation at 5 Mw. UNCLASSIFIED 2-01-060-50 5 ; I [ I TEMPERATURE 66 °F~COLD CLEAN 4 \\ ATION IN SEPARATION DISTANCE REACTIVITY CORRECTED FOR VARI- x OVER PLATE AREA, S | 2 3 P : g \\ALL SAFETY GRIDS > E E 2 \\ O > c {2 — = o s 1 - ) = 4 0.8 \ I = x ul I - 0] 0 {0 20 30 40 50 60 70 RADIUS {cm) Fig. 4.5.12, Thermal-Neutron Flux in the TSR-1! as a Function of Radius: 5-Mw Power Level. 195 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2-0{-060-44 150 — Coe e e e SQURCE INTEGRAL 4T N - —] = 150.326 neutrons /sec 140 SOURCE INTEGRAL AT 5 Mw, 3.84 x 10'7 130 120 "o ———- [ SOURCE DISTRIBUTION (neutronsscm 7sec) {00 90 | : ! 80 i ! | 1 i 23 25 27 29 L1 33 35 37 RADIUS IN FUEL REGION (cm) Fig. 4.5.13, Source Distribution in the TSR-Il Fuel Region as a Function of Radius: Boral Shell 4 cm from the Fuel. Of the 6.8% increase in the reactivity, 1.5% is to cover the worth of the control grids in the fully withdrawn position, 0.3% is allowed for structural material in the core region, 1.4% is to cover the worth of fixed neutron absorber plates in the in- ternal reflector region, 2.0% is to allow for an error in the calculations, and 1.6% is the amount that is controlled by the control grids. The 1.6% that is controlled by the grids represents approximately one-half the reactivity worth of the shim-safety grids and covers the following: 0.6% to counteract the negative temperature coefficient of reactivity from 40 to 180°F, 0.3% excess re- activity necessary to permit servo control, 0.6% to counteract xenon buildup for 8 hr of operation at 5 Mw, and 0.1% for short-time burnup of U233, The 0.3% allowed for the structural material in the core region was determined by calculating the re- activity change caused by deleting the U235 from shells in the core region. An integral value was obtained by summing, as a function of radius, the areas of these shells which were intersected by the cylindrical support members. 196 The 1.4% is to cover the fixed neutron absorbing plates which will be mounted in the internal re- flector region near areas of the core-reflector in- terface not covered by the movable plates. These plates may be adjusted when the reactor is unloaded « so that some shimming may be done in the initial loading, as long as sufficient excess reactivity is left to provide 0.8% for U233 burnup. A 2.0% excess was allowed for calculational errors because the comparison of the calculation with critical experiments showed an average error of this amount, The TSR-I| elements (see Fig. 4.5.14) have been fabricated to contain the 8.1 kg of U2 established as reported above, However, since the geometry of the core permits only a limited adjustment of excess reactivity by control plates and no means of adding excess reactivity by the insertion of additional fuel elements, the fuel elements have been assembled in a temporary fashion so that the core loading can be altered by changing a limited number of fuel plates if a critical experiment should indicate that it is necessary. Such a critical experiment will be . performed at the Critical Experiments Facility. When the fuel elements are assembled in their final form the worth of the control plates will be deteri- mined and some of the reactivity coefficients will be measured. CALCULATION OF HEATING IN THE FUEL PLATES A hand calculation was performed to estimate the heating per fuel plate in each element. The source distribution given in Fig. 4.5.13 was used to de- termine the power distribution in each fuel plate as a function of the length. The power generation for some of the plates in the central fuel elements is shown in Fig. 4.5.15 and for the annular elements in Fig. 4.5.16. There are twelve annular elements and eight central elements. The total power per fuel plate was found by integrating the area under the distribution plots for each fuel plate. The re- sults are plotted as a function of the fuel plate number for the central fuel elements in Fig. 4.5.17 . and for the annular fuel elements in Fig. 4.5.18. These values will be used in the flow distribution studies to determine the required flow per fuel . channel (see below). PERIOD ENDING SEPTEMBER 30, 1958 |UNCLASSIFIED PHOTO 45083 Fig. 45.14. Photo of TSR-Il Fuel Elements. 197 861 UNCLASSIFIED 2-0i-060-46 i \ | 76 -~ |-&— EDGE DISTANCE (VARIES FROM 0.250 in. FOR PLATE NO. \\ TO 0.486 in. FOR PLATE NO. 40). 72 i’/_‘*“‘_—"__'—_"'_‘"\{ 1 ] | | | ( . 68 e /T FUEL REGION \ \ FUEL PLATE 64 \ \ PLATE AVERAGE POWER GENERATION —---rd NO. (Btu /hr/t2) PLATE NO. 40 4 53,980 ™ \ 10 54,270 ] N 20 54,650 \ 30 30 55,340 40 56,980 56 SN20 N U 52 10’ s \\ N\ N 60 ] POWER GENERATION (BTu/hr/ffz) 48 N A\ NN 40 10 12 14 16 18 20 22 24 26 FUEL LENGTH {cm) O no D 2} @ Fig. 4.5.15. Power Generation in the Fuel Plates of a Central Fuel Element as a Function of the Distance from the Inside Edge of the Fuel. LId0d3IY §SSIH00dd LI23F0¥d ANV 661 POWER GENERATION (Btu/hr/ft2) UN CLASSIFIED 2-01-060-47 80 I ' 76 \\ \ | HORIZONTAL MIDPL ANE ; PLATE AVERAGE POWER GENERATION S M \ u/hrsfte 7 \\ q {_““—I ______ )i--— FUEL REGION Z?' (8;91:6;” ) ! 42 58,190 \ | 44 55,520 68 \ N FUEL PLATE :S 5;‘31:?323 \ \ 53 51,350 ca N~ . 56 50,010 \ PLATE NO. 58 48,510 T d o 60 \\44 — 7 41,350 \\\:wEE\ 56 e . = \ 53 ‘\ Q o '\_‘\5856 \\\s\\\ 44 I— 66 \‘\\ X ™~ \ N 74 \ \\ \I w0 — 7 T 1T o) 2 4 6 8 10 12 14 16 18 20 2 24 26 28 30 FUEL LENGTH {(cm) Fig. 4.5.16. Power Generation in the Fuel Plates of an Annular Fuel Element as a Function of the Distance from the Horizontal Midplane to the Edge of the Fuel. 8561 ‘0 ¥IEWILJ3IS ONIONT a0l¥3d 00C POWER PER FUEL PLATE (kw) 10 UNCLASSIFIED 2-01-060-42 INNER RADIUS OF PLATE{4=1.600 in. PLATE THICKNESS = 60 mils A CENTER-TO-CENTER SPACING OF FUEL PLATES =180 mils NUMBER OF CENTRAL FUEL ELEMENTS = 8 (4 ABOVE MIDPLANE, 4 BELOW) /// 5 10 15 20 25 30 35 40 CENTRAL FUEL ELEMENT PLATE NUMBER Fig. 4.5.17, Power Generation in the Fuel Plates of a Central Fuel Element as a Function of the Fuel Plate Number. 1Y0d3IY SSIYO0¥d LD3rodd dNV 102 POWER PER FUEL PLATE (kw) UNCLASSIFIED 2-01-060-58 14 2 10 \_\\\ 8 ™. 5 N INNER EDGE OF PLATE 41 IS AT 8.945 in. 4 ————— PLATE THICKNESS = 60 mils \ CENTER-TO-CENTER SPACING OF FUEL PLATES = {80 mils NUMBER OF ANNULAR FUEL ELEMENTS = 42 40 45 50 55 60 65 70 ANNULAR FUEL PLATE NUMBER Fig. 4.5.18. Power Generation in the Fuel Plates of an Annular Fuel Element as a Function of the Fuel Plate Number. 75 8561 ‘0 YIHWILHIS ONIANT QOiy¥3d ANP PROJECT PROGRESS REPORT FLOW DISTRIBUTION STUDIES Flow Studies of Central Fuel Elements The hydraulic flow test stand which has been as- sembled to test the confrol mechanisms under 1000- gpm flow conditions is shown in Fig. 4.5.19, and the complete piping arrangement and holdup tank is shown in Fig. 4.5.20. This test stand will also be used to study the flow distribution in the central fuel elements and in the internal reflector region. The stand will be extended to accommodate not only the central fuel elements but also the lead-and- water shield above the core region. Flow Studies of Annular Fuel Elements W. R. Gambill'? H. W. Hoffman'?2 The full-scale, quarter-sphere model constructed for study of the flow features of the annular fuel elements in the TSR-I| core is shown in two stages of assembly in Figs. 4.5.21 and 4.5.22. Figure 4.5,21 shows the containment fixture, the ends of an inner fuel plate assembly, and the separator plate. In Fig. 4,5.22 the second-pass *‘‘orange ' or outside fuel plates, have been placed in slices,’ position and covered with a Lucite shell and back- up cage. The groove in the head-end plate (shown in the left lower corner of Fig. 4.5.22) receives a pass rib which separates inlet and exit water fiows. A movable mockup of the control plate assembly is located in the first pass cavity separating the two The entire model is The separation between inner fuel element sections. fabricated from aluminum. the inner surface of the Lucite shell and the outer circumference of the outside fuel plates is approxi- mately 0.5 in. The static pressure drop across the full length of every other flow channel is measured with manom- eters connected to pressure taps located at channel inlets and exits. Taps for the outer (second pass) flow channels are visible in Fig. 4.5.21. The manometer fluid is a CCl -1, solution. The flow model, horizontal for the present tests, will later be placed in a vertical position. The flow rate is measured with a calibrated orifice and a differential pressure (DP) cell, the nozzle inlet and exit pressure with Bourdon-type gages, and the water temperature with a standard Weston pipe- line thermometer. The watertemperature may be 12Reactor Projects Division. 202 varied from 58 to 102°F with a steam-heated heat exchanger. Tests to date have been made at a con- stant flow rate of 250 gpm. The measured static pressure drops for the model in the ‘“‘as-received’’ state are shown in Fig. 4.5.23. Variation of control plate position had very little effect on the pressure drops for the second-pass flow channels, as was expected; since all the first- pass flow must reverse through the 0.5-in. gap around the separator plate. Extreme radial asymmetry of flow is evident. A very large portion of the first-pass pressure drop is apparently due to flow complications in the central control cavity. It is probable that some sort of coupled vortex flow pattern, induced by the geometry and the control assembly mockup, exists in the central cavity. In the outer region of the core it was found that most of the water entered the first seven channels and that flow in the remaining outersregion channels was either negligible or slightly reversed. Visual observation of air bubbles, injected upstream of the first pass, confirmed the right portion of Fig. 4.5.23. Thus, it was observed that the bubbles streamed through the first few channels but were in reverse flow in the outermost channels. Velocities have been calculated from the pressure- drop data of Fig. 4.5,23. These are relatively un- informative for the first pass because of the large extraneous flow resistance associated with the central cavity. For the uninterrupted channels of the second pass, however, the results indicating velocities from —2 to + 11 ft/sec are considered to be valid. Calculations were based on the recent parallel-plate relations given by Rothfus er al., 13 which indicate a Reynolds number of ~ 7000 as the upper extension of the transition region or lowest level of fully established turbulence. Preliminary calculations indicate that the core hydrodynamics will be more of a problem than the heat transfer as such. |f velocities are sufficient for full turbulence to exist, fuel element surface temperatures will not exceed saturation temperatures even at low static pressure levels, To maintain full turbulence in all channels, however, a larger total flow rate may be required. For a Reynolds number of 7000 in each first-pass flow channel, a ]3R. R. Rothfus et al., J. Am. Inst, Chem, Eng. 3, 208 (1957). PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED PHOTO 44878 Fig 4.5.19. Hydraulic Flow Test Unit. 203 ANP PROJECT PROGRESS REPORT Fig. 4.5.21. Full-Scale Quarter-Sphere Model for Flow Test Studies of Annular Fuel Elements (Before Instal- lation of Outer Pass). 204 AP(psi) PERIOD ENDING SEPTEMBER 30, 1958 Fig. 4.5.22. Full-Scale Quarter-Sphere Model for Flow Test Studies of Annular Fuel Elements (Outer Pass and Plexiglas Shell Installed). o8 06 oz UNCLASSIFIED 2-01-060-48 INSIDE ZONE (FIRST PASS) T OUTSIDE ZONE (SECOND PASS) ! | e I | T T | | CONTROL PLATE AT MAXIMUM RADIUS [ e 5 0| \/] ‘ | 1| FLOW RATE: 250 gom e — WATER TEMPERATURE: 58°F | MODEL POSITION: HORIZONTAL o | | | CONTROL PLATE HALF-EXTENDED \‘[;// 1 f / CONTROL PLATE AT MINIMUM RADIUS —/ | T A / T RS SN ~— SEPARATOR PLATE | | 6 20 24 28 32 36 40 44 48 52 56 60 FLOW CHANNEL NUMBER Fig. 4.5.23. Radial Variation of Pressure Drop in the Fuel Element Flow Test. 64 &8 205 ANP PROJECT PROGRESS REPORT total flow of 910 gpm is necessary; for the outer region, a flow of 1380 gpm is required. These tlow rates are based on an equivalent diameter that is equal to twice the water gap between the plates and an evaluation of viscesity at mean pass temperatures (131°F for the first pass and 148°F for the second pass). The fuel element surface-to- centerline temperature rise appears to be negligible for the worst operating condition. The next test will be made with hot water and with sealing strips around the equator of the outside re- gion to stop a small bypass stream observed earlier, Later tests will involve placing flow resistances (perforated plates or screens) in the inlet ends of the smaller-radius outer-region flow channels. It is hoped that a redistribution of flow-satisfying heat- transfer requirements can be attained in this fashion. Flow redistribution in the first pass is potentially a considerably more difficult problem, and studies of first-pass flow have barely begun. A flow test in which cold water will pass through a single channel of an outside fuel-element section is being prepared. The flow rate and static pressure drop will be measured, and a calculated friction factor vs Reynolds number plot should reveal the extent of the transition region for the particular entry conditions and curved fuel plates of the TSR-]| reactor. REACTOR CORE KINETIC STUDIES The results of the initial kinetic studies of the reactor core, which were made on the ORNL Re- actor Controls Analog Facility, 4 are presented here without analysis. The first two cases were run to determine how well the control mechanism could protect the reactor without the inherent safety features of a heterogeneous reactor. Three other cases were run to partially investigate the self- limiting features of the reactor based on the temper- ature coefficient of reactivity. With the reactor system simulator operating at an initial power of 5 Mw and in a prompt critical con- dition, excess reactivity was added and the simu- lator delayed neutron circuits were closed. The re- sulting power excursion was recorded and the safety rods were allowed to drop when a level of 7.5 Mw was reached. Concurrent with the reactor power M4 This investigation was made by R, K. Adams, F. P, Green, and E. R. Mann of the Instrumentation and Con- trols Division. 206 excursion the mean temperature of the exit fuel section was recorded. Both the power excursion data and the mean temperature are plotted on Fig. 4,5.24a. The net negative temperature coefficient of the reactor was disconnected; hence no limiting effects of temperature coefficient were present. For purposes of comparison, the time vs reactivity profile of the safety grids is plotted on the same figure (see Fig. 4.5,246). The same type of in- formation for an initial power of 5 Mw and a level trip at 6 Mw is presented in Fig. 4.5.25. The safety grid profile shows that the insertion time for the grids was longer in this case. With the reactor operating at an initial power of 0.5 Mw and with the net temperature coefficient of the reactor “‘active,’’ the water temperature leaving the air cooler was reduced from 145.8 to 32°F. (There is an approximately 130-sec delay time be- fore the water reaches the reactor.) The resultant power excursion and system temperature changes are plotted as functions of time in Fig. 4.5.26. The value which was used for the temperature coefficient of reactivity (Ak/k) was ~6.5%. Later calculations and critical experiments show the value to be near that of the BSF (see above). These cases will be re-run. Figure 4.5.27 pre- sents the same type of data for an initial power of S Mw, With the reactor operating at 5 Mw, an abrupt stop in cooling water flow was also simulated, The re- sulting reactor power and system temperature changes are plotted in Fig. 4.5.28. SAFETY SYSTEM MEASUREMENTS J. E. Marks The evaluation of the TSR-I[ Safety System was based upon a value of delay time and a time function of control plate position that were determined ex- perimentally with a test assembly incorporating the best designs of the proposed system components. The measuring techniques described below were considered capable of yielding the accuracy re- quired for a reliable reactor analysis. The test assembly included a simulated sigma bus circuit, a standard magnet amplifier, an electro- hydraulic transducer, and a control mechanism with connections between these components similar to those expected in the final reactor assembly. A differential transformer type of linear motion trans- ducer was mechanically connected to the control plate to translate the position of the plate into a 1000 900 — 800 700 600 500 REACTOR POWER (Mw) 400 300 200 100 | N d REACTIVITY (% Ak/k) i W Fig. 4.5.24. (a) Reactor Power Level and Fuel Temperature Rise as a Function of Time After Insertion of PERIOD ENDING SEPTEMBER 30, 1958 UNCLASSIFIED 2-01-060-57 N [ POWER EXCURSION - POWER EXCURSION | PERIOD = 070 ¢ 102 1 | 970 x 10° Mw 8.6 msec 2 300 1O x40 w‘""\/- | w o« 2 |~ 9.0msec g s 8.70 x (0™ Mw = g 8.70 x10% Mw o E \ w ¥ 200 B s ’ = o d 5.55 x 10° Mw Z 100 f / z 2.7 %102 Mw w = 0 [ 0 40 80 100 5.55 x {0° Mw TIME (msec) {0.5 msec 2.7 x10% Mw %\ 1.0 msec g INITIAL / \ POWER 5 Mw é | S LEVEL TRIP AT \ 7.5 Mw \ SAFETY GRID PROFILE (TIME AFTER TRIP) \\ ] \ 0 20 40 60 80 100 120 140 160 180 TIME (msec) 200 Various Amounts of Positive Excess Reactivity with Power Overshoot Being Limited Only by Safety Grid Insertion After . TSR-Il Reaches 7.5 Mev. (b) Rate of Reoctivity Insertion by Safety Grids as a Function of Time After Scram. 207 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2-01-060-52 1000 POWER EXCURSION 900 : b —m 3 PERIOD 0.9 x 10 16.5 msec / 800 — e e [ w 3 Q_ TOO — - [ o —_— . - — [ I E— .| = 0.7x10° = Z o 600 [t — - }—. .| l 2 3 Y 0.5x10 500 _ [ w = (@] a & 400 - — = (&) g w [0} © 300 — L« & 3 0.25 x 10 22.3 msec 200 100 INITIAL POWER -0 \ o &_ 0 , —— LEVEL TRIP AT 1.2 x I— INITIAL POWER \ X - \ — 3 SAFETY GRID PROFILE \ a (TIME AFTER TRIP) : \ > }—. a9 L @x -3 \_\ -4 0 20 40 60 80 100 120 140 160 180 200 TIME (msec) Fig. 4.5.25. (a} Ratio of Power Level to Initial Power as a Function of Time After Insertion of Various Amounts of Positive Excess Reactivity with Power Overshoot Being Limited Only by Safety Grid Insertion After TSR-H Reaches 6 Mw. (b) Rate of Reactivity Insertion by Safety Grids as a Function of Time After Scram. 208 PERIOD ENDING SEPTEMRER 30, 1958 UNCLASSIFIED 2-01-060-53 , 50 500 T ‘ 1 ‘ 1 i ot ’/ | /V ’/{// 400 ‘ . ’ ‘ ‘ / - | Z | EXIT MEAN ‘ T #’ FUEL TEMPERATURE -— // L 300 — ;, , : ~—REACTOR POWER . [} | 1 /»' y x = ‘ o P | v BOILING POINT—="1" E L -~ J o r% : / H 200 : - v s EXIT MEAN ‘ / / R FUEL TEMPERATURE — o F | A /< _74.7 ! i —— - EXIT SECTION MEAN /\\:\__ —] & / I P 5 BOILING POINT —__ y . « S\ v ! — N - RS e e (e o . S 7 E / //- ‘ ,//» // 5 / o e - — 200 T // —— / ,// I - 4 - P ‘-——-——-_-_- FUEL EXIT — ey . WATER TEMPERATURE S~ EXIT SECTION MEAN s / I WATER TEMPERATURE Sy 100 » -] ced - . S / ! A L INLET WATER | , e TEMPERATURE ~— | — e : : o , 0 0.5 L0 5 2.0 2.5 3.0 35 TIME {sec) 40 30 20 REACTOR POWER {Mw) Fig. 4.5.27. Reactor Power and Fuel Temperature as a Function of Time After the Initial Reduction of the Inlet Water Temperature from 145.8 to 32°F. Reactor power initially at 5 Mw. 209 ANP PROJECT PROGRESS REPORT UNCLASSIFIED 2-01-060-55 300 T —— | EXIT MEAN FUEL SRR IR _— ; TEMPERATURE\\ R e s 250 : Lol s 4 BOILING POINT ~—_ L ——/7"""_# 4 e ! o S R ; | g ) /'/’ A = ~—— EXIT SECTION s s ~ e 200 MEAN WATER /'/ S 7 3z S S e - o / l TEMPERATURE SN S s S x 2 o 3 < & & . - x s S BRI TR N S {1, 4 \ — ~ = Wi /, / e y ’, ) m \/REACTOR POWER o _ I ] | \\ \ s // S 100 — Y ./// 50 RA PR I 0 0.5 1.0 15 2.0 35 TIME (sec) Fig. 4.5.28. Reactor Power ond Exit Section Mean Fuel and Mean Water Temperoture os a Function of Time After Loss of Coolant Woter Flow. 3000-cps differential voltage. This voltage was applied to a cathode-ray oscilloscope with a 200-msec sweep period initiated by the same change in sigma bus voltage as that which initiated the scram. The plot of position vs time that resulted was calibrated in tenths of an inch between the limits of travel of the control plate. The plate was considered to have started and stopped moving when the slope of the curve left and retumed to zero. The delay time was measured by using the change of sigma bus potential that initiated the scram to simultaneously trigger the sweep of the oscilloscope. The length of sweep from the point it began to the first discontinuity was taken as an indication of the *‘dead’’ time between initiation of a scram signal and the first perceptible motion of the control plate. SHIELD DESIGNS The design criteria for the beam shield which will be used initially with the TSR-ll were that (1) on a 210 Reactor initially at 5 Mw. rem basis and with an rbe of 10 for fast neutrons, there was to be an equal number of gamma rays and fast neutrons at the reactor shield surface, and (2) the background dose at the crew compartment from the beam shield 64 ft away was to be a factor of 50 to 100 less than that produced from a beam at right angles to the reactor-crew compartment axis. With these criteria and a gamma-ray shield of 50-50 volume per cent lead and water, the following shield dimensions were obtained.'® The lead-and-water gamma-ray shield should be 43 ¢m thick with an inner radius of 48 cm. The water neutron shield should be 81 cm thick, which gives an outer shield radius of 172 cm, Thus, the shield will contain 33,500 Ib of lead and 42,900 |b of water. The possibility of using 5/us-in. raschig rings with a ’/a-in.-dicu central opening in the lead-and-water lsThis work was performed by Lloyd Byrnes of GE- ANP, Cincinnati, Ohio. gamma-ray shield is being investigated. The beam holes will be stepped halfway through the shield, with the outer cylinder 15 in. in diameter and the inner cylinder 10 in. in diameter., The Laboratory is proceeding with the design of this shield with the additional criterion that it must be self-supporting when it is placed on the ground. A second high-performance shield for the TSR-I is being designed by the Y-12 Engineering Depart- ment according to specifications set by Pratt & Whitney Aircraft. The shield will utilize a de- pleted uranium gamma-ray shadow shield and a lithium hydride neutron shield. [t will be fabricated in the Y-12 shops and is scheduled to be completed by May 1, 1959. The estimated cost of the shield is $330,000. INVESTIGATION OF STRESSES IN THE TOWER STRUCTURE FROM WATER AND ELECTRICAL LINES An investigation has been made !¢ to determine whether water hoses and electrical cables can be suspended from the reactor support cables and tower legs without exceeding the allowable stresses in the tower structure. The investigation was carried out for the following conditions: 1. weight of reactor = 55 tons, 2. weight of crew compartment = 30 tons, 3. 80-mph wind (20 psf), 4, two water hoses, each 170 ft long and weighing 25 Ib/ft, 5. four electrical cables, each 170 ft long and weighing 2.75 |b/ft. The reactor and crew compartment were considered to be in the plane of normal operation, that is, in the 16, A, McCarthy, Report of Investigation for the TSF-1I, McPherson Co., July 1958, SO PERIOD ENDING SEPTEMBER 30, 1958 vertical plane bisecting the longitudinal axis of the Tower Shielding Facility (east and west) and ex- tending 200 ft from the ground level. The water hoses and electrical cables were as- sumed to be symmetrically positioned from tower legs | and Il. One end of each hose and each cable was considered to be attached at a distance of 123 ft 6 in. above the base of the leg and the other end was considered to be attached to the reactor support cables at a point 33 ft from the reactor. The calculations were compared to a similar cal- culation made in connection with the original de- sign of Knappen-Tippetts-Abbott-McCarthy (KTAM). 7 The calculation by KTAM did not take into account items 4 and 5 above and showed the allowable stresses in the tower legs to be: 1. as given by KTAM, 19.90 ksi, and 2. allowed by AISC code, 19.55 ksi. The stresses in the tower legs determined in this calculation are: 1. before addition of water hoses and electrical cables, 18,072 ksi, and 2. after addition of water hoses and electrical cables, 19.545 ksi. In the calculation it was assumed that the electrical cables and water lines were rigidly fastened to the tower legs. In the present design all four electrical cables will be suspended from Leg | and one water hose will be suspended from Leg | and the other from Leg Il. However, the suspension arrangement, which is shown in Fig. 4.5.29, eliminates the greater portion of the horizontal load in the tower legs by shifting it to the deadmen which now anchor the tower legs. ”Loadz'ng Criteria and Analysis for Tower Shielding Facility, Knappen-Tippetts-AbbotteMcCarthy Co,, New York, 1953, UNCLASSIFIED 2-01-060-56 -s— TOWER LEG MESSENGER CABLE FOR WATER HOSE GUY ANCHOR ] ' 1 HOSE TO REACTO)/7/ HOSE FROM STANDPIPE) Fig. 4.5.29. Tower Suspension Arrangement of Reactor Cooling Water Lines. 212 VONO U AN~ A A @EMMO = CMEEPIMZAOMPETPREI>COCCOMECODROPMAM . Allen . Billington . Blankenship . Blizard . Boudreau . Boyd . Breeding . Briggs . Bruce . Callihan . Center (K-25) . Charpie . Clifford Coobs . Cottrell . Culler . Cuneo DeVan . Doney . Douglas . Emlet (K-25) . Fraas Frye . Furgerson Gray . Greenstreet . Grimes . Grindell Guth . S. Harrill . Hikido R. Hill E. Hoffman W. Hoffman . Hollaender . H. Jordan . W. Keilholtz . L. Keller OAr—“—AToorzTaroImMmrprmoxTtmemmMUmMw =X . J. Keyes . S. Livingston . N. Lyon INTERNAL DISTRIBUTION 42. 43. 44, 45. 46. 47. 48. 49, 50. 51. 52. 53. 54. 55. 56. 57. 58. 59. 60. 61. 62. 63. 64. 65. 66. 67. 68. 69. 70. /1. 72, 73. 74, /5. 76. 77. /78-80. 81-87. 88. 89-91. ORNL -2599 C-85 —~ Reactors-Aircraft Nuclear Propulsion Systems M-3679 (22nd ed.) . G. MacPherson . C. Maienschein . D. Manly . R. Mann . J. Miiler . Z. Morgan . J. Murphy P. Murray (Y-12) L. Nelson . J. Nessle . G. Overholser . Patriarca . K. Penny . M. Perry . M. Reyling . W. Savage . W. Savolainen . D. Schultheiss L. Scott . D. Shipley . Simon . J. Skinner . H. Snell . Storto . Susano . Swartout . Trauger . Trubey . Watson . Weinberg White . Wigner {consultant) . Williams . Wilson . Winters W. Zobel ORNL - Y-12 Technical Library, Document Reference Section Laboratory Records Department Laboratory Records, ORNL R.C, Central Research Library TrOIXT-MA>METNT OCOME> 000N M>TPM-TP> T VP> mONoNDzzxmPO 213 EXTERNAL DISTRIBUTION 92-95. Air Force Ballistic Missile Division 96. AFPR, Boeing, Seattle 97. AFPR, Boeing, Wichita 98. AFPR, Douglas, Long Beach 99-101. AFPR, Douglas, Santa Monica 102-103. AFPR, Lockheed, Marietta 104. AFPR, North American, Canoga Park 105. AFPR, North American, Downey 106. AFPR, North American, Los Angeles 107-108. Air Force Special Weapons Center 109-110. Air Research and Development Command (RDZN) 111, Air Technical Intelligence Center 112, Air University Library 113-115. ANP Project Office, Convair, Fort Worth 116. Albuquerque Operations Office 117. Argonne National Laboratory 118. Armed Forces Special Weapons Project, Sandia 119. Armed Forces Special Weapons Project, Washington 120-121. Army Ballistic Missile Agency 122, Army Rocket and Guided Missile Agency 123. Assistant Secretary of Defense, R&D (WSEG) 124. Assistant Secretary of the Air Force, R&D 125-130. Atomic Energy Commission, Washington 131-133. Bettis Plant (WAPD) 134. Brookhaven National Laboratory 135. Bureau of Aeronautics 136. Bureau of Aeronautics General Representative 137. BAR, Aerojet-General, Azusa 138. BAR, Chance Vought, Dallas 139. BAR, Convair, San Diego 140. BAR, Goodyear Aircraft, Akron 141. BAR, Grumman Aircraft, Bethpage 142. BAR, Martin, Baltimore 143. Bureau of Ships 144. Bureau of Yards and Docks 145-146. Chicago Operations Office 147. Chicago Patent Group 148. Curtiss-Wright, Quehanna 149, Director of Naval Intelligence 150. duPont Company, Aiken 151-158. General Electric Company (ANPD) 159-160. General Electric Company, Richland 161. Hartford Aircraft Reactors Area Office 162. {daho Test Division (LARQO) 163. Jet Propulsion Laboratory 164. Knolls Atomic Power Laboratory 165. Lockland Aircraft Reactors Operations Office 166-167. Los Alamos Scientific Laboratory 168. Marquardt Aircraft Company 169. National Advisory Committee for Aeronautics, Cleveland 170. 171. 172. 173. 174. 175. 176. 177. 178. 179. 180-181. 182-185. 186. 187. 188-189. 190. 191-192. 193. 194-195. 196-208. 209-233. 234, ‘ National Advisory Committee for Aeronautics, Washington Naval Air Development Center Naval Air Material Center Naval Air Turbine Test Station Naval Research Laboratory New York Operations Office Oak Ridge Operations Office Office of Naval Research Office of the Chief of Naval Operations Patent Branch, Washington Phillips Petroleum Company (NRTS) Pratt and Whitney Aircraft Division Sandia Corporation Sandia Corporation, Livermore School of Aviation Medicine USAF Headquarters USAF Project RAND U. S. Naval Radiological Defense Laboratory University of California Radiation Laboratory, Livermore Wright Air Development Center Technical Information Service Extension Division of Research and Development, Atomic Energy Commission, QOak Ridge Operations 216 Reports previously issued in this series are as follows: ORNL-528 ORNL-629 ORNL-768 ORNL-858 ORNL-919 ANP-60 ANP-65 ORNL-1154 ORNL-1170 ORNL-1227 ORNL-1294 ORNL-1375 ORNL-1439 ORNL-1515 ORNL-1556 ORNL-1609 ORNL-1649 ORNL-1692 ORNL-1729 ORNL-1771 ORNL-1816 ORNL-1864 ORNL-1896 ORNL-1947 ORNL-2012 ORNL-2061 ORNL-2106 ORNL-2157 ORNL-2221 ORNL-2274 ORNL-2340 ORNL-2387 ORNL-2440 ORNL-2517 Period Ending November 30, 1949 Period Ending February 28, 1950 Period Ending May 31, 1950 Period Ending August 31, 1950 Period Ending December 10, 1950 Period Ending March 10, 1951 Period Ending June 10, 1951 Period Ending September 10, 1951 Period Ending December 10, 1951 Period Ending March 10, 1952 Period Ending June 10, 1952 Period Ending September 10, 1952 Period Ending December 10, 1952 Period Ending March 10, 1953 Period Ending June 10, 1953 Period Ending September 10, 1953 Period Ending December 10, 1953 Period Ending March 10, 1954 Period Ending June 10, 1954 Period Ending September 10, 1954 Period Ending December 10, 1954 Period Ending March 10, 1955 Period Ending June 10, 1955 Period Ending September 10, 1955 Period Ending December 10, 1955 Period Ending March 10, 1956 Period Ending June 10, 1956 Period Ending September 10, 1956 Period Ending December 31, 1956 Period Ending March 31, 1957 Period Ending June 30, 1957 Period Ending September 30, 1957 Period Ending December 31, 1957 Period Ending March 31, 1958