U 3 4ysgg 0383042 4 - This report was prepared as an accouni of work sponsored by the United States Gowvernment. Neither the United States nor the United Siates Atomic Energy Commission, nor any of their employess, nor any of their contractors, subcontractors, or their employess, makes any warranty, exXpress or implied, or assurmes ' any legal Hability o:r vesponsibility for the accuracy, complietengss or usefuiness of amy information, apparatus, product or process disclosed, or represents that its use would; not infringe arivately owned rights, ORNL-TM-3141 Contract No. W-TLh05-eng-26 CHEMICAL TECHNOLOGY DIVISION ENGINEERING DEVELOPMENT STUDIES FOR MOLTEN~SALT BREEDER REACTOR PROCESSING NO. 6 L. ¥. McNeese DECEMBER 1971 OAK RIDGE NATIONAIL LABORATORY Oak Ridge, Tennessee 37830 operated by UNTON CARBIDE CORPORATION for the U.S. ATOMIC ENERGY COMMISSION ii Reports previously issued in this ORNL~436M ORNL~4365 ORNI~%366 ORNL~TM-3053 ORNI~TM-3137 ORNT~TM~3138 ORNIL~TM-3139 ORNL~TM~ 3140 Period Period Period Period Period Period Period Period series ending ending ending ending ending ending ending ending are as follows: Merch 1968 June 1968 September 1968 December 1968 March 1969 June 1969 September 1969 December 1969 iii CONTENTS Page SUPJMAR IES ° . ° . . . . ° . . s e . » . . . . . . . . - . - ° 2 . . V l . INTRODUC‘I‘ION . . . a » . . * . ¢ . ° @ . . . . ® - » . . . ’ . l 2. MSBR FUEL PROCESSING USING FLUORINATION~-REDUCTIVE EXTRACTION AND THE METAL TRANSFER PROCESS . . . ¢ v v v ¢ v v v v o o« . . 1 2.1 Eguilibrium Data and Concentrations . « « + ¢« v o + + « & D 2.2 Flowsheet Analysis .« .+ « ¢ v & ¢ v ¢ ¢ ¢ v e e 4 e e a 4 2.3 [Effect of Contamination of LiCl with Fluoride . . . . . . 13 3. AXIAL DISPERSION IN OPEN BUBBLE COLUMNS . . « « v o « v + + « « 13 3.1 Previous Studies on Axial Dispersion . . . . . . . . « . 16 Equipment and Experimental Technique . . . . . . . . . . 16 3.3 Effects of Gas Inlet Diameter and Column Diameter on Axial Dispersion . « « & o o v o o v v e e e v e e e e e AT 3.4 Cas Holdup in Bubble Columns . . . + +« & 4« o 4o o & & & « 19 3.5 Discussion of Results and Future Experiments . . . . . . 292 4. CONSIDERATIONS OF CONTINUOUS FLUORINATORS AND THEIR APPLICARBIL- ITY TO MSBR PROCESSING + « ¢« v v v v v v o v v o 0 0 0 o v o v 23 L.1 Types of Fluorinators . . « « v« v « v v v 4 v v v v v « o DY L.2 Experience Related to Fluorination of Molten Salt for Uranium Removal . o o o o v ¢« v v v 0 v v v v v v v v o o D2f 4.3 Mathematical Analysis of Open-Column Continuous Fluorina- OS¢ v v v 6 v vt e b e e s e e e e s e e e e . DB L.4 Evaluation of Fluorination Reaction Rate Constant . . . . 30 L.5 Prediected Performance of Open-Column Continuous Fluorina- tOI'S ¢ & 8 8 e B 8 2 & & & & & » & e & a4 & s = 3w s & » 3}_,_ 5. USE OF RADIO~-FREQUENCY INDUCTION HEATING FOR FROZEN-WALL FLUORI- NATOR DEVELOPMENT STUDIES . . o o o o v v v« v o v v o o v v v 30 5.1 Mathematical Analysis . . . « « . ¢ ¢ « ¢ v o v v v v .. N1 5.2 Calculated Results for a Molten-Salt Fluorinator . . . . L§ 5.3 Experimentally Measured Heat Generation Rates . . . « . « L& TR b B e iv CONTENTS (continued) DEVELOPMENT OF THE METAL TRANSFER PROCESS . . . . . . . . . . . 52 7.1 Equipment and Experimental Procedure . . . . o « « . . . 52 7.2 Development and Testing of a Pump for Circulating LiCL . s5i ELECTROLYTIC CELL DEVELOPMENT: STATIC CELL EXPERIMENTS . . . . 57 STUDY OF THE PURIFICATION OF SALT BY CONTINUQUS METHODS . . . . 99 9.1 Previous Work on Salt Purification . .+ . « . « + .+ + .+ . 50 9.2 Experimental Equipment . . . « . « o v v o 0 0 e v 0w 61 9.3 Gas Bupply and Purification Systems . « .« . « « . + + . 6l 9.4 Installation of Fquipment and Initial Checkout . . . . . §5 9.5 Anticipated Experiments and Operating Procedures . . . . 71 SEMICONTINUOUS REDUCTIVE EXTRACTION EXPERIMENTS TN A MILD-STEEL FACIDTTY v v v v v v v e e e e e e e e e e e e e e e e e e e T3 10.1 Equipment Modifications . . . .« . . « « ¢« « « o o o v . T3 10.2 Tregtment of Bismuth and Salt; Adjustment of Zirconium Distribution Hatio . . « « « + « + « o« v ¢ « v o « « « .« 75 10.3 Hydrodynamic Experiments HR-9, -10, -11, and -12 . . . . 75 10.4 Maintenance of Equipment . . . . . « ¢ v « v v+« . . . 7O REFERENCES v v v v v v v v e e e e e e e e e e v e e v e v e v 80 SUMMARTES MSBR FUEL PROCESSING USING FLUORINATION--REDUCTIVE EXTRACTION AND THE METAL TRANSFER PROCESS A combined flowsheet for processing MSBR fuel salt by fluorination-- reductive extraction and the metal transfer process has been devised. Calculations have been made. based on recently measured distribution co- efficients,tor a number of rare-earth and actinide elements. Reference conditions for the isolation of protactinium on a 10-day cycle are given, and the effects of several parameters associated with rare-earth removal are discussed. Conditions that result in rare~earth removal times of about 15 to 50 days are described. The effect of contamination of the LiClL with fluoride ions was examined. It was found that the fluoride concentration will have to be maintained below about 2 mole % in order to avoid a high thorium discard rate. AXTAL DISPERSICN IN OFPEN BUBBLE COLUMNS Measurements of axial dispersion during the countercurrent flow of air and water were made in 1.5-, 2-, and 3-in.-diam columns with a range of gas inlet diameters. In the "slugging' region, the dispersion coeffi~ cient was found to be independent of gas inlet diameter and dependent only on the volumetric gas flow rate for all column diameters. In the "pubbly'" region, the dispersion coefficient also appears to depend only on the volumetric gas flow rate when the column diameter is 2 in. or larger. Gas holdup in bubble columns was also determined for a range of operating conditions., CONSIDERATIONS OF CONTINUOUS FLUORINATORS AND THEIR APPLICABILITY TC MSBR PROCESSING A great deal of experience has been accumulated in removing uranium from molten salt by batch fluorination; however, information on continuous fluorinators is sparse, particularly on fluorinators capable of handling vi salt flow rates up to about 100 ftB/day. Experience with fluorinators is reviewed, and possible types of fluorinators are discussed. A math- ematical analysis of open-colunn continuous fluorinators is presented, and predictions are made concerning the performance of cpen-column continuous fluorinators for MSBR processing applications. USt OF RADIO-FREQUENCY INDUCTION HEATIUWG FOR FROZEN-WALL FLUORINATOR DEVELOPMENT STUDIES Radio~frequency induction heating is being considered as a method for generating heat in molten salt in studies of frozen-wall fluorinators with nonradiocactive salt. Two configurations for an inductively heated contimious fluorinator are discussed. Caleculations for the first con- figuration show that sufficient heat would be generated in a 1.9~in.~-diam molten zone by a coil current of 24.7 4 at 500 kHz tc maintain a 1.5-in.- thick frozen salt film with a2 100°C temperature difference across the film. ‘The calculated efficiency of heating the salt was about 34%; the heat generated in the metal walls was about 1.05 times the heat generated in the salt. In experiments with a 3-in.-diam clhirge T 30% H2SDM sur- rounded by a 6-in.-long section of 6~in. sched 40 pipe, the measured ratio of heat generated in the pipe to that in the acid was 1.3, whereas the calculated ratio for the system was 0.58. This discrepancy shows that the design of an experimental fluorinator using induction heating will depend heavily on empirical design relations. The second configuration could not be examined mathematically. Ex- QSOM showed that the ratio of heat generated in the pipe to that generated in the acid was perimental measurements using this configuration with 30% H 0.069. The coupling of the magnetic field with the acid was weaker with this configuration than with the first configuration. MSRE DISTILLATION EXPERIMENT Data obtained in the MSRE Distillation Experiment for the effective relative volatilities, with respect to LiF, of BeF;, ZrFu, and fluorides L _ 8 of QSZr, lthe, 1 TPm, lSSEu, le, 9OSr, gSr, and 13705 were examined in an sttempt to explain the anomalous relative volatilities of all 95 fission products except Zr. 'These data were scrutinized closely for possible evidences of enfrainment, concentration polarization, and sample contamination. Although all three effects were prcbably present, we believe that sample contamination was the major reason for the dis- crepancies between the values obtained in this experiment and those measured under eguilibrium conditions. The low relative volatility 137 obgserved for Cs is not explained by any of the three mechanisms examined. DEVELOPMENT OF THE METAL TRANSFER PROCESS Eguipment has been fabricated for studying and demonstrating the metal transfer process for removal of rare earths from MSBR fuel salt. Work that will demonstrate all phases of the process 1s under way. Lanthanum and luTNd will be extracted from fuel carrier salt by contact with bismuth containing thorium. The rare earths will then be selec- tively transferred fo LiCl. The final step of the experiment will con-~ sist of removing the rare earths from the LiCl by contact with bismuth containing 0.4 mole fraction lithium. Several pumps made of quartz have been designed and tested with molten LiCl at 650°C in an effort to develop a device that is capable of circulating the LiCl in the experiment. Although difficulty has been encountered with devitrification of the quertz, we believe that the LiCl can be sufficlently purified to permit a quartz pump to perform satisfactorily. One pump was found tc be cperable after tests with LiCl at 650°C over a 16-day period. ELECTROLYTIC CELL DEVELOPMENT: BSTATIC CELL EXPERIMENTS A static cell electrolysis experiment was made in an all-metal cell to determine whether the presence of quartz contributed to the formation viii of the black material found to be present in the salt Thase in other electrolysis experiments. No such material was observed in this experi- ment ; however, the lack of a bismuth cathode may have resulted in a sys- tem too different from the previous cells to allow us te draw firm con- clusions. STUDY OF T PURIFICATION OF SALT BY CONTINUQUS METHODS To date, the molten salt required for development worxk as ..l as for the MSRE has been purified from harmful contaminants (sulfur, oxygen, and iron fluoride) by a batch process. It is believed that the costs of the labor associated with salt purification can be reduced considerably by using a continuous process for the most time-consuming operation (i.e., the hydrogen reduction of iron fluoride). We have installed equipment in which molten salt and hyérogzzr carn be countercurrently contacted in a 1.25-in.~diam, 8l~in.-long wacxad column. The system is fabricated of nickel, and provision is made for feeding about 15 liters of molten salt through the column at flow rates of 50 to 250 ems/min. The equipment and gas supply systems are descrited, and the anticipated experimental program is outlined. SEMICONTINUOUS REDUCTIVE EXTRACTION EXPERIMENTS IN A MILD-STEEL FACILITY A new column, packed with 1/h-in. molybienum Raschig rings, was installed in the system. Minor changes were made in some of the piping. Three successful hydrodynamic experiments were made in which bismuth and molten salt were contacted countercurrently. The results are in excellent agreement with a flooding correlati-n developed from work with the mercury-water system. Results of a hydrodynamic experiment with salt flow only established that the pressure drop for the new column was in satisfactory agreement with that predicted from a liter- ature correlation. 1. INTRODUCTION A molten-salt breeder reactor (MSBR) will be fueled with a molten fluoride mixture that will circulate through the blanket and core regions of the reactor and through the primary heat exchangers. We are develop- ing processing methods for use in a close-coupled facility for removing Tfission products, corrosion products, and fissile materials from the molten flucride mixture. Several operations associated with MSBR processing are under study. The remaining parts of this report discuss: (1) a flowsheet for process— ing MSBR fuel salt by fluorination--reductive extraction and the metal transfer process, (2) measurements of axial dispersion coefficients in open bubble columns, (3) considerations of continuous fluorinators and their applicability to MSBR processing, (4) an evaluation of radio- frequency induction heating for frozen~wall fluorinator development studies, (5) an examination of several explanations for the anomalous relative volatility data obtained In the MSRE Distillation Experiment, (6) the design and testing of equipment for demonstration of the metal transfer process for removal of rare earths from MSBR fuel carrier salt, (7) the operation of a static electrolytic cell in an all-metal system, (8) a study of the purification of salt by countinuous methods, and (9) experiments conducted in a mild-steel reductive extraction facility to increase our understanding of the hydrodynamics of packed column operation during the countercurrent flow of molten salt and bismuth. This work was carried out in the Chemical Technology Division during the period January through March 1670, 2. MSBR FUEL PROCESSING USING FLUCRINATION--REDUCTIVE EXTRACTICN AND THE METAL TRANSFER PROCESS M. J. Bell I,. E. McNeese Recently, we reportedl the development of the metal transfer process for extraction of rare-earth fission products from MSBR fuel salt and presented removal times for several rare earths for a range of operating conditions. Noting that this process eliminated the need for large elec- trolytic cells, we introduced another process not reguiring an electrolytie cell, namely, the fluorination--reductive extraction process for isolation of protactinium from fuel salt. ©Since then, we have devised a combined MSBR processing flowsheet that uses fluorination--reductive extraction for the isolation cof protactinium and the metal transfer process for rare- earth removal. A range of operating conditionsg for the processing plant has been examined, and the MATADOR code2 has been used te calculate the breeding ratio corresponding to each set of conditions. Additional in- formation ou the distribution of rare earths between LiCl and Bi contain- ing reductant has become available; this information indicates that sat- isfactory removal times can be obtained for Ba, Nd, and Sm, as well as for Eu and La (as previously reported). 2.1 Equilibrium Data and Concentrations Ferris and co—worker83 have measured the distribution coefficients of several fission products and actinide elements between a number of acceptor salts and molten bismuth containing lithium. t a given temperature, the distribution coefficients for an element M can be expressed as * = + log DM n log XLi log KM, where XLi is the mole fraction of lithium in the bismuth phase, n is * the valence of M in the salt phase, and log KM is a constant. The distribution coefficient i1s defined as . mole fraction of M in bismuth phase M mole fraction of M in salt phase Their results, summarized in Table 1., indicate that either LiCl or LiBr would constitute a suitable acceptor salt, and that satisfactory removal Table 1. Distrivution Coefficient Data Values of log K* Derived from * log D = n log XLi + log K Temffé;“”‘e Salt Element log K* 630 LiCl Eu®* 2.301 640 LiCl Ba®t 1.702 La™ 7.973 Ng* 8.633 Sm** 2.886 Th¥* 15.358 pa 17.838 u* 11.278 640 LiCFLiF (98.1-1.9 mole %) Th* 13.974 640 LiCHLIF (964 mole %) Th* 12.90 pa* 14.7 u* 10.80 640 LiCI-LiF (90-10 mole %) La® 7.288 Th* 11.309 600 LiCL-LiF (80-20 mole %) La* 7.235 Nd:: 7.644 Th 10.964 640 LiCI-LiF (80-20 mole %) La: 7.124 Th 10.629 700 LiCFLiF (80-20 mole %) Nd: 6.732 Th 9,602 575 LiBr Ra 1.497 600 LiBt Ba®* 1.443 1a* 9.079 Ndz: 8.919 Th 16.16 640 LiBr La* 8.266 Ng* 8.834 650 LiBr Ba®™ 1.358 700 LiBr Baz; 1.316 Nd 8.430 600 LiBr-LiF (90-10 mole %) La™ 8.158 Th* 12,380 600 LiBr-LiF (80-20 mole %) La™ 7.840 W™ 11373 times can be obtained for Ba, La, Nd, Sm, and Eu. These data were used to evaluate the performance of the metal transfer process for removing strontium, barium, and the rare earths from MSBR fuel salt. Strontium was assumed to distribute in a manner similar to barium. and the trivalent rare earths for which distribution data were not available were assumed to have distribution characteristics like those of neo- dymium . These assumptions are believed to be conservative. 2.2 Flowsheet Analysis A combined flowsheet for processing MSBR fuel salt using fluorina- tion-~reductive extraction and the metal transfer process is shown in Fig. 1. The effects of various operating parameters for the Pa isola- tion system on the Pa removal time apd the uranium inventory in the Pa decay tank have been reported p‘r'eviously.:L A 10~-day protactinium removal time is obtained with a fuel salt flow rate of 0.88 gpm (10~ day processing cyele), a bismuth flow rate of 0.23 gpm, two stages in the lower contactor, six to eight stages in the upper contactor, and column diameters of less than & in. A decay tank volume of 200 to 300 ft3 is required. Reductant nust e supplied at the rate of 340 to 420 equivalents per day, which costs 0.012 to 0.015 mill/kWhr. This system also results in a 10-day removal time for materials that are more noble than thorium and do not form veolatile fluorides during fluorination; these include‘Zr,esiPa, Pu, Bh, Pd. Ag, Cd, In, Ni, and other corrosion products. The conceptual flowsheet (Fig, ?2) for the metal transfer process includes four salt-metal contactors that operate at 640°C. Fuel salt from the Pa isolation system, which is free of U and Pa but which con- tains the rare earths at the reactor concentration, is countercurrently contacted with Bi containing approximately 0.002 mole fraction Li and 0.0025 mole fraction Th (90% of the solubility of thorium at 6L0°C) in contactor 1. Significant fractions of the rare earths transfer to the downflowing metal stream and are carried into contactor 2. Here, the PROCESSED SALT ORNL DWG 70-28i11 tion—Reductive Extraction and the SALT U PURIFICATION REDUCTION o ! A ! 2 | { FQ } } i 1 { i UF g I REACTOR COLLECTION i i { i lur, i S i } i i | 1 i | ! FLUORINATOR Fa SALT CONTAIKING PaF, AND UF, - Fig. 3 SALT CONTAINING RARE EARTHS H - Pg - DECAY - EXTRACTGR UF, -{ oxumzeaj::jnuommmfl_—. SALT 10 SALY ? Hy” HF fa - 1 EXTRACTOR i EXTRACTOR 1 - Bi-L]| Lict l % 0.3 MOLE FRAC. EXTRACTOR 1 | | | | i 1 i ~d |__________.__.___f I I | | | I 1 I | | 1 | ! —l - -— Bi-ij (0.0% WOLE FRAC. il EXTRACTOR Bi-Li — — 4 + TRIVALENT RARE EARTHS | b — e e e © L_T______._T_.l I 1 I i t I i 1 1 t ! ! i ! § i 1 | ] 1 1 i Bi- L4 ! Lo +DEVALENT RARE : ! i | 1 I I | 4 ] { i 1 1 i b et e ke A A e RECUCTANT OXIDIJER F—t e — —— — ADDITION H,~HF Lf,Th Fiowsheet for Processing & Single-¥Fluid MSBR by Fluorina- Metal Transfer Process. ORNL DWG 70-2812 PROCESSED SALT TO < == REACTOR 5\‘_th | | EXTRACTOR 1| 3-6 STAGES | FUEL SALT } ; (No U cor Pa) | EXTRACTOR 2| 3-6 STAGES | | - ) === 8i-Lj LiCl H (0.5 MOLE FRAC. Li) EXTRACTOR 3 2-3 STAGES frJ Bi-Li b — — o + DIVALENT : - RARE EARTHS { - Bi-Li (0.05 MOLE FRAC. Li) I l . | | | beg | EXTRACTOR 4: 1 STAGE | I i—-. l Bi-Li o i 4 TRIVALENT J ] RARE EARTHS Fig. 2. Metal Transfer Process for Removal of Rare Earths from a single-Fluid MSBR, bismuth stream is contacted countercurrently with LiCl, and significant fractions of the rare earths and a trace of the thorium transfer to the LiCl. The resulting LiCl stream is then routed to contactor 4, where it is contacted with a bismuth soluticn containing 0.05 mele fraction lithium for removal of the trivalent rare earths. About 2% of the LiCl is routed to contactor 3, where it is contacted with a bismuth solution containing 0.5 mole fraction lithium for removal of the divalent rare carths (Sm and Eu) and the alksline earths. The LiCl from contactors 3 and 4 (still containing some rare earths) is then returned to con- tactor 2. The trivalent and divalent rare earths are removed from the LiCl in separate contactors in order to minimize the amount of lithium re- quired. Removal of these elements in separate contactors appears ad- visable for several reasons. A high lithium concentration in the biz- muth is required for obtaining adequately high distribution coefficients for the divalent rare earths. However, the solubilities of the triva- lent rare earths in bismuth are much lower than those of the divalent elements. Also, the production rate for the trivalent rare earths is several times that of the divalent elements. Calculations were made to identify the important system parameters for the metal transfer process. Figure 3 illustrates the effect of the bismuth flow rate through contactors 1 and 2 on the removal time for necdymium, a typical trivalent rare earth, and samarium, a typical divalent rare earth, for a fixed LiCl flow rate. The divalent materials distribute less readily to the metal phase, and high bismuth flow rates are regquired to achleve significant removal of these materials. On the other hand, Fig. 4 illustrates that, for a fixed bismuth flow rate, the divalent rare earths transfer quite readily to the LiCl but that high LiCl flow rates are required to achieve removal of the trivalent rare earths,. The overall effect of the bismuth and LiCl flow rates on the removal of rare-earth fission products 1s illustrated in Fig. 5. It is seen that the reactor performance is relatively insensitive to in- ORNL -DWG-70-2816 30 u T ¥ T . ] i 80 80 ~ 50} - REMOVAL TIME, days Li CI FLOW RATE = 33 gpm aol- - 30 - 2a}- - 10 J 1 1 1 1 o 5 10 15 20 25 BISMUTH FLOW RATE, gpm Fig. 3. Effect of BRismuth Flow Rate Through Contactors 1 and 2 on the Removal Times of Neodymium and Samarium, Using the Metal Transfer Process. ORNL DWG TO-2BI7 9 0 T I | I T 70} ~ BISMUTH FLOW RATE=12.3 gpm REMOVAL TIME, days a0}~ 30 20— =~ | l | { 1 i0 5 20 25 30 as 40 LiC! FLOW RATE,gpm Fig. b. Effect of LiCl Flow Rate on the Removal Times of Neodymium and Samarium, Using the Metal Transfer Process, 10 ORNL DWG 70-10,994 0.064 T I T I 1 | ] | T Bi FLOW RATE - qpm) e e""T20.7 0.063 | o 16.6 12.4 0,062 T ey £ o - 3 & 2 0061} e o o 8.3 i) L) [ o - 0.060 {— e STAGES IN UPPER CONTACTORS=3 ™ STAGES IN TRIVALENT STRIPPER=1 STAGES IN DIVALENT STRIPPER=2 0.05% — — ey by 20 25 30 35 40 LiCl FLOW RATE {gpm) Fig. 5. Overall Effect of LiCl and Bismuth Flow Rates in the Metal Transfer System on MSBR Performance. 11 creages in the LiCl flow rate above 33 gpm. A substantial increase in the breeding gain (breeding ratio minus 1) is obtained by increasing the bismuth flow rate from 8.3 gpm to 12.4 gpm. Further increases in the bismuth flow rate do not produce corresponding gains in reactor performance. Bismuth and LiCl flow rates of 12.4 gpm and 33 gpm, re- spectively, have been selected for the reference processing conditions. Figure 6 shows the effect of the number of stages in the fuel salt-- bismuth and the LiCl-bismuth contactors. Little benefit is obtained from using more than three stages; therefore, threse stages is considered optimunm. Only one stage is required to extract the trivalent rare earths from the LiC1 in contactor 4. The flow rate of the Li-Bi solution through this contactor is 8.1 gpm, and 5.7 gal of the metal stream must be removed daily to prevent the solubilities of the trivalent rare esarths in the bis- muth from being exceeded. The bismuth can be recovered by hydrofluorinst- ing or hydrochlorinating the metal stream in the presence of salt, which can be processed further (if economical) or discarded. The divalent rare earths, plus strontium and barium, can be stripped from the LiCl by passing 2% of the LiCl (0.66 gpm) through a small two-stage extractor where it is 3 o bismuth--50 at. % lithium per minute. The bis- contacted with 1.5 cm muth in the metal stream can again be recovered by hydrofluorination or hydrochlorination in the presence of a waste salt. The rare-earth removal times that can be obtained using the reference processing ccnditions range from about 15 to 50 days (see Table 2). Table 2. Fission Product Removal Times for Metal Transfer Process Under Reference Condltions Removal Time Flement (days) Ba2+ 16.8 La3+ 22.0 wast 29.9 + Sm2 27.0 muot 51.0 ORNL DWG 70-10,993 - Licl FLOW RATE 0.063 — (gprm) — 33 0.062 — — < < o — o 1 Z 0.061}— s L) 1) ar = X 0.060 +— - BISMUTH FLOW RATE =12.3 gpm . STAGES IN TRIVALENT STRIPPER={ | STAGES IN DIVALENT STRIPPER=2 0.053 ] s __ 4o e 2 3 4 5 & NUMBER OF STAGES IN UPPER CONTACTORS Fig. 6. Effect of the Number of Stages in the Fuel Salt~~Bismuth and LiCl-~Bismuth Contactors on MSBR Performance. 13 2.3 Effect of Contamination of LiCl with Fluoride The presence of fluoride in the LiCl acceptor salt causes a sig- nificant decrease in the thorium distribution coefficient, as shown in Fig. 7. This results in an increase in the extent to which the thorium transfers to the LiCl and is undesirable since the thorium is subsequently extracted, along with rare earths, from the LiCl into the Li-Bi solutions and is discarded. As shown in Fig. 8, the thorium loss rate increases from 0.41 mole/day with no fluoride in the LiCl to 280 moles/day when the LiCl contains 5 mole % LiF. It is likely that the fluoride concentration in the LiCl will have to be kept below about 2 mole %, which corresponds to a thorium transfer rate of 7.7 moles/day. Discard of thorium at this rate would add 0.0013 mill/kWhr to the fuel cycle cost. The effect of the presence of fluoride in the LiCl on the removal of rare earths is negligibley in fact, the rare-earth removal efficiency increases slightly s the fluoride concentration in the LiCl increases. 3. AXTAL DISPERSTION IN OPEN BUBBLE COLUMNS J. 8. Watson L. E. McNeese Axial dispersion 1s important in the design and performance of con- tinuous fluorinators. Since molten salt saturated with fluorine is cor- rosive, fluorinators will be simple, open vessels that have a protective layer of frozen salt on all exposed metal surfaces. In such systems, the rising gas bubbles may cause appreciable axial dispersion in the salt. We have been involved, for some time, in a program for measuring axial dispersion during the countercurrent flow of air and water in open bubble columns. The objectives of this program are to evaluate the effect of axial dispersion on fluorinator performance and to account for this effect in the design of fluorinators. The data reported in this gection result largely from a study by ) A. A. Jeje and C. R. Bozzuto,Jr of the MIT Practice School, who investi- 1k ORNL-DWG 70— 4503 1?1T1II1iii]i|Tl|IIl‘I|1 TEMP ELEMENT SALT (°c) . La3* LiBr—LiF 600 La®* LiCl-LiIF 640 — Th** LiBr-LiF 800 Th** LiCI-LiF 640 A 1 16 R A O A @ Log K* ? L 1 b L I 0 0.05 0. 0145 0.2 MOLE FRACTION LiF IN SALT Fig. 7. Values of log K* for Lanthanum and Thorium, Using LiBr-LiF and LiCl-LiF as the Salt Phase. 15 ORNL DWG 70-10,995 ' [ * I T I T l : [ 200 — 100 {— — 50 (— - ,;; - —{ -0 I . P 4 o 2 g 201 ] Lt - ] - < x 10— — @ L i O — i - » . E e a s el 2 5 o = " O X | ] !._ 2 — - . TEMPERATURE=640°C 1 Th CONC IN Bi=30% OF — Th SOLUBILITY AT 640°C STAGES IN UPPER COLUMNS=3 i 0.5 STAGES IN DIVALENT STRIPPER:=2 ] L STAGES IN TRIVALENT STRIPPER=1 03 | | | s | | I . | 0.00 0.0t 0.02 0.03 0.04 0.05 MOLE FRACTION LiF Fig. 8. Effect of LiF Contaminant in LiCl on Thorium Loss Rate in Metal Transfer Process. gated the effects of gas inlet diameter and column diameter on axial dispersion in open bubble columns during the countercurrent flow of alr and water. 3.1 Previous SBtudies on Axial Dispersion Initial studies on axial dispersion in open columns were carried out by Bautista and Mc:Neese,5 who studied axial dispersion during the countercurrent flow of air and water in a 2-in.~ID, T2~in.~long column. Two regions of operation were observed., The first of these consisted of a "bubbly"” region, at low gas flow rates, in which the air moved up the column as individual bubbles and coalescence was minimal. The second region consisted of a "slugging" region, at higher gas flow rates, in which the air coalesced rapidly into bubbles having diameters equal to the column diameter. A plot of the logarithm of the dispersion coeffi- cient vs the logarithm of the gas flow rate was linear in both regions. However, the slope of the line representing data in the slugging region was higher than that for data in the bubbly region. The transition be- tween the two regions was well defined. The same column and equipment were used by A. M. Sheikh and J. D. Dearth,6 of the MIT Practice School, for investigating the effects of the viscosity and the surface tension of the liguid. The dispersion coefficient was found to decrease in the bubbly region as the viscosity of the ligquid was increased from 1 to 15 ¢cP by the addition of glycerol to the water; little effect was noted in the slugging region. An in- crease in the dispersion coefficient was observed as the surface tension of the liquid was decreased by the addition of n~butanol to the water. 3.2 Equipment and Experimental Technique The egquipment and the experimental technique used in the study described here are the same as those used in the previous studies; a T detailed description was given previously. The technique involved 7 continuously injecting a tracer solution (cupric nitrate) into the bottom of the column and determining the resulting steady-state tracer concentra- tion profile at points upstream along the column axis. The tracer concen- tration was measured at each of 20 sampling points that were equally spaced (3.5 in. apart) along the 72-in.-long column. At each gampling point, a miniature centrifugal pump was used for circulating solution be- tween the column and a photocell for determination of the tracer concen- tration. The solution was withdrawn from, and returned to, opposite sides of the column at the same elevation. Thrae column diameters (1.5, 2, and 3 in. ID) and four gas inlet diameters (0.019, 0.04, 0.06, and 0.085 in.) were used by the MIT Practice School group. Later, measurements were also made with a 0,17- in.~ID gas inlet. The fraction of the column volume occupied by the gas phase during the countercurrent flow of air and water was measured by two different techniques. The first consisted of measuring (1) the height of the air-water mixture above the air inlet while air was flowing through the column, and (2) the height of the water above the gas inlet after the alr flow was turned off. The gas holdup wasg then determined from the difference in these values. The second technique consisted of attaching a water~filled manometer to points along the column axis. The mancometer reading indicated the settled height of liquid above the point of attach- ment. The second method was found to be more rapid, and also allowed holdup measurements to be made for upper porticns of the column. 3.3 Effects of Gag Inlet Diameter and Column Diameter cn Axial Dispersion As shown in Fig. 9, the dispersion coefficient values measured for a 2-in.~diam column are not dependent on gas inlet diameter in either the slugging region or the bubbly region. A few tests were alsc made using the 1.5~ and 3~in.-diam columns with a O0.1lT-in.-diam gas inlet; no effect of gas inlet diameter was noted. DISPERSION COEFFICIENT {(cm?/sec) 18 ORNL DWG 70-11,000 100 80 . 60 |- ] 50 — 40— ] 30— s & 83 MIL ORIFICE 20+ B 60 MIL ORIFICE — 4 40 ML ORIFICE 10 | 1 ] b b L L1 | 10 20 30 40 eC 80 100 200 GAS FLOW RATE (cc/sec) Fig. 9. Effects of Gas Flow Rate and Orifice Diameter on Axial Dispersion in a 2-in.-diam Open Column. 19 It is not surprising that changes in the gas inlet dlameter have no effect in the slugging region since a great deal of coalescence occurs in this region and the bubble size distribution gquickly becomes independ- ent of the initial size distribution. However, in the bubbly region (low gas Tlow rates), where coalescence is minimal, it was thought that gas inlet diameter might be important. It has been postulatedh that "chain bubbling" occurs at low gas flow rates (i.e., that gas bubbles are not formed consecutively). In this case, the bubble diameter would not be highly dependent on gas inlet diameter. Even with different bubble diam- eters, one would not expect large differences in bubble rise velocities or, possibly, large changes in dispersion coefficient. Davies and Taylor8 report that the bubble rise velocity is proportional to the sixth root of the bubble volume. Thus, even for conditions resulting in the release of single bubbles (where bubble volume is proportional to the cube root of the inlet diameter), one would find little variation in rise velocities and, possibly, little variation in dispersion coefficient with changes in gas inlet diameter. The effects of gas flow rate and column diameter are shown in Fig. 10 for column diameters of 1.5, 2, and 3 in. In the slugging region, there is little difference in dispersion coefficient for the three col- umn diameters at a given volumetric gas flow rate. However, the dats do not extend to high gas flow rates for all column diameters, and extrapolation to higher gas flow rates is guestionable since uncer- tainty in the data is greatest at high gas flow rates. The dispersion coefficients for the 2-in.~ and 3-in.~diam columns do not differ appreciably in the bubbly region. On the other hand, data from the 1l.5-in.-dianm column show significantly lower valuesg for the dis- persion coefficlent and a greater dependence of dispersion coefficient on gas flow rate in this region. 3.4 Gas Holdup in Bubble Columns Fxperimentally determined gas holdup values are summarized in Fig. 11, which shows the effects of superficial gas velocitly and column diameter DISPERSION COEFFICIENT (cm?%sec) 100 80 60 50 490 30 20 ORNL DWG 70-11,00t1 i 1 1 1 117 z T T T T T T m N A ] A _ % — A W ] /A' ‘ ~PREVIOUS 2-in. COLUMN DATA ® g{ ® 7 : -~ . e - | \ B m " Bal \ _________ f - e _ e ® o ® B 3-in COLUMN B A 2-in. COLUMN o ® ® 15-in COLUMN ® ® ® Lt | Lt | 2 5 10 20 40 50 60 B8O 400 200 GAS FLOW RATE {cc/sec) Fig. 10. Effect of Column Diameter on Dispersion Coefficient in an Cpen Coiumn. 0c GAS HOLDUP ORNL DWG 70-14710 1 B3 in. COLUMN A 2 in. COLUMN ® |5 in. COLUMN i O 1 1 L L 1 L 1 L 3 | 4 O { 2 3 4 5 6 7 8 9 10 H i2 13 4 SUPERFICIAL GAS VELOCITY (cm/sec) Fig. 11. Variation of Gas Holdup with Superficial Gas Velocity in Open Columns Having Diameters of 1.5, 2, and 3 in. 1 22 on holdup. At low gas flow rates, holdup is linearly dependent on super— ficial gas velocity and is independent of column diameter. However, a transition from this bhehavior is observed at a superficial gas velocity of 2 to 3 cm/sec. At superficial velocities above the transition, the holdup data for the various column diameters diverge; the holdup is greatest for the smallest column diameter. Data for the 3~in.-diam col- umn do not extend beyond the transition region. The transition region corresponds roughly to the transition between bubbly and slug flow, as determined from measurements of the dispersion coefficient. Values for holdup in the transition region are thought to be inaccurate and will be checked during future studies. There was no detectable variation of heldup with axial position along the column for any operating conditions tested. 3.5 Discussion of Results and Future Experiments Although the present data on axial dispersion are incomplete, one can make the following tentative conclusions: (1) In the slugging region, the dispersion coefficient appears to be proportional 1o the square root of the volumetric gas flow rate and independent of column dlameter. (2) TIn the bubbly region, the dispersion coefficient is only a function of the volumetric gas flow rate for columns that are 2 in. or larger in diameter. The dispersion coefficient date for a 1.5~in.~diam column deviate from this condition. Additional information 1s needed in order to confirm these conu- clusions. The reported data become less accurate as the column diameter is increased since air, which collects in the photocell sample circuits, prevents flow through the photocells. We will consider methods for preventing the buildup of air in the sample circuits as well as alter~ native methods for measuring dispersion coefficients. 23 Future studies will be concerned primarily with obtaining data in the region of slugging flow, where continuous fluorinators of interest will operate. Gas inlets other than the small-diameter tubes used thus far will be studied. Most of the effort will be concentrated on side inlets of large diameter since this type of inlet appears to be the most emenable to protection against corrosion by use of a frozen wall. 4}, CONSIDERATICNS OF CONTINUOUS FLUORINATORS AND THEIR APPLICABILITY TO MSBR PROCESSING L. . McNeese J. 8. Watson Most of the flowsheetsgnl2 considered to date for processing MSBR fuel salt reguire fluorination of molten salt for removal of uranium at one or more points. These applications include: (1) removal of trace guantities of uranium from relatively small salt streams prior to dis- card, (2) removal of uranium from a captive salt volume in which 233Pa 2 is accumulated and held for decay to 23”U, (3) removal of most of the uranium from relatively large fuel salt streams prior to isolation of protactinium and removal of rare earths, and (k) nearly quantitative 233 removal of uranium from a salt stream contalning Pa in order to produce isotopically pure 233Ua Not all of these applications reguire continuous fluorinators; in fact, the use of batch fluorinators results in definite advantages in certain cases. However, as the guantities of salt and uranium to be handled increase, the use of continuous flu- orinators becomes mandatory in order to avoid undesirably large inven- tory charges on uranium and molten salt as well as the detrimental increase in reactor doubling time that is assoclated with an increased fissile inventory. Although the literature contains many references to the removal of uranium from molten salt by batch fluorination, information on continuocus fluocrinators, particularly on fluorinators capable of handling salt flow rates on the order of 100 ftg/day, is rather meager. The remalinder 24 of this section is devoted to a review of the experience with flu- orinators, a discussion of the possible types of fluorinators, a mathematical analysis of open-column continuous fluorinators, and predictions of the performance of open-column continuous fluorinators for MSBR processing applications. 4.1 Types of Fluorinators No known materials of construction are resistant to attack by the extremely corrosive enviromment resulting from the combined action of fluocrine and molten salt313 for this reason, any discussion of flu- crinator types must also give consideration to the relative ease of protecting the fluorinator from corrosion. Several types of fluori- nators, all of which can be clagsified as either batch or centinuous, have been considered in the past. In a batch fluorinator, the molten salt is contacted with fluorine for a sufficient time to reduce the uranium concentration to the desired level. A bateh fluorinator is especially useful in cases where it is desired to remove small guan- tities of uranium from molten salt prior to its discard. In batech operations, the salt can be analyzed repeatedly and the probability of inadvertent discard of fissile material can be reduced to an ac- ceptably low level. However, this type of fluorinator is not well suited to the removal of large quantities of uranium from salt streams having flow rates of 50 ft3/day or greater unless one can tolerate the large salt and uranium inventories that result. Batch fluorinators are amenable to protection against corresion by use of a frozen wall, provided the heat generation rate in the salt is adequate for supporting the thermal gradient necessary for maintaining a layer of frozen salt adjacent to molten salt. In this method of cor- rosion protection, a layer of frozen salt is maintained on all metal surfaces that are expected to be contacted by molten salt. This allows the buildup of a protective NiF, layer, which would otherwise be dis— 2 solved by the molten salt. 25 At least three types of continuous fluorinators have been con- sidered inthe past; these are: (1) open columns in which an appreciable axial uranium concentration gradient exists, (2) a series of well-mixed vessels, and (3) falling-drop fluorinators in which molten salt droplets are allowed to fall through a gas phase containing f‘luorine,lh Open- column continuous fluorinators have the advantage of a relatively small salt holdup and hence a low uranium inventory. This type of fluorinator, which consists of a simple open cylinder, can be designed for frozen-wall protection against corrosion for a fairly wide range of gpecific heat generation rates in the salt. Also, 1t appears to ve capable of a relatively high salt throughput with a low salt inventory. However, this type of contactor depends on the establishment of a significant axlal uranium concentration gradient; for this reascon, axial dispersion in the salt phase 1s important. Many of the same arguments apply to a continuous fluorinator consisting of a series of well-mixed vessels. The galt inventory in such a system is likely to be larger than in the case of an open=~column fluorinator designed for a given salt throughput. No attempt would be made to minimize dispersion in the salt phase in a given vessel. Protection of the intervessel salt transfer lines against corrosion for this type of fluorinator might be difficult. The falling-drop fluorinator appears to have several advantages over other types of continuous fluorinators. TFor example, the salt inventory is quite low, and a lafge quantity of gas can be contacted with a small amount of salt. On the other hand, this system has two disadvantages that have not presently been circumvented: (1) thermsl convection currents in the gas tend to sweep the molten salt droplets into the fluorinator wall, resulting in a highly corrosive condition, and (2) the device used for dispersing the salt into small droplets is subject to corrosion. It appears that the open column is the best type of continuous fluorinator for MSBBR processing applications requiring removal of 50 to 09% of the uranium from salt streams that have flow rates on the order of 100 ftg/day; thus this flucorinator has been selected for further development. 26 4.2 Experience Related to Fluorination of Molten Salt for Uranium Removal The removal of uranium from molten fluoride salts by bateh fluori- nation has been studied extensively in the laboratory,ls in engineering experiments,l6 and in the ORNL Fused 5alt Fluoride volatility Process Pilot Plant. Several speat reactor fuels cooled for periods as short 17,18 as 25 days were used in these studies. The most recent demonstra- tions of batch fluorination consist of: (1) the recovery of about 6.5 kg of uranium from Th £t> of MSRE flush salt (66-34 mole % LiFmBng) during a 6.€-hr operation, which resulted in a final uranium concentra- 19 tion of 7 ppm in the salt,”” and (2) the recovery of about 21L4 kg of uranium from Th £t° of MSRE fuel salt (65-30-5 mole % LiF~BeF2erFh) during a 6-day operation (fluorination time, 47 hr), which resulted in 19 a final uranium concentration of 26 ppm in the salt. Data obtalined from the above systems demonstrate that the concentra- tion of uranium in molten salt can be decreased by batch fluorination to very low levels. Although the equilibria involved have not been measured, the formation of UF6 is strongly favored. It is believed that the uranium concentration in molten salt in equilibrium with FQMUF6 mixtures contain- ing low concentrations of fluorine is very low; for the present data, it is i_distinguishable from zero. Experience with continuous fluorination is limited to a single study by McNeese.gO The fluorinator used in this study had a salt depth of 48 in. and was constructed from l-in.-diam (nominal) nickel pipe. No attempt was made to protect the fluorinator walls from corrosion, and a relatively high corrosiocn rate occurred. The uranium removal efficiency was deter- mined by analyzing the inlet and outlet salt streams. The fluorine ntilization could not be determined because of the high corrosion rate. Two temperatures (525 and 600°C) and two inlet uranium concentrations in the salt (0.12 and 0.35 mole %) were used. The results of this study are summarized in Fig. 12. 27 - ORNL--DWG 67—8405 10 . . ................................................... Lo T _____ e fj_.;fi.}ffififijj;ffifij N { T N | 0.35 mole % UE, | i . 52500 1 / | S ' T a !_ : Q : /O 12 mole % UF4 | LLi 5 i < Lofttv 600°C ‘ < i : i g ; | Ll i J x e [ i ] (o _ - | O : = E = / | 2 G.35 mole % UF4 Z A 0 4 X 600°C [l o1 (A S e b { ----- ] L [ [T ; O L ' . - | UF, CONC IN g FEED SALT TEMPERATURE ™~ Q 4 (mole %) (°C) i | = 0.35 525 T A 0.2 600 - i ® 0.35 800 oL = | | ‘ ! | i | | I | . | L L 10 20 50 100 200 500 1000 SALT THROUGHPUT (g--moles /hpr) Fig. 12. Variation of the Residual Uranium with Salt Flow Rate, Temperature, and Inlet Uranium Concentration During the Continuous Flucrina- tion of Molten Salt. 28 Three runs for each temperalure and inlet uranium concentration (a total of nine runs) were made using a range of salt flow rates. The fraction of the uranium remcved was high in all cases, ranging from 97.5 to almost 99.9%. The fluorination rate (and hence the fraction of the uranium removed) increased as the temperature was increased; the fraction of the uranium removed decreased as the inlet uranium concentration was decreased. These data serve primarily to demonstrate that high uranium refioval efficiencies can be obtained with an open~column conbtinuous fluorinator. However, they are sufficiently reliable for use with data on axial dispersion in open columns to allow estimating the performance of larger fluorinators, which are required for M3BR processing. A study has also been made of the hydrodynsmics of heat transfer in a simulated open-column continuous flucorinator having a frozen wall as a means of protection against corrosion.gl The esquipment consisted of a 5-in.-diam, 8-ft-long pipe in which molten salt (66-34 mole % LiFerFM) and argon were countercurrently contacted. A volumetric heat source in the molten salt was simulated ty Calrod heaters located along the center line of the pipe, and provision was made for cooling the pipe wall. The system was operated over a range of conditions that included heat genera- tion rates in the molten salt as high as 55.7 kW/ftB, which is greater than the heat generation rate in salt Just removed from the reference 1000-MW(e) MSBR. The thickness of the frozen salt film on the fluorinsator wall ranged from 0.3 to 0.8 in., depending on the operating conditions. The film was symmetricel and adhered to the metzl wall. 4.3 Mathematical Analysis of Open-Column Continuous Fluorinators Consider a differential height of a continucus fluorinator in which fluorine and molten salt containing uranium are in countercurrent flow. If the rate of removal of uranium from the salt is assumed to be first order with respect to the uranium concentration in the salt, a material balance on the differential volume yields the relation 2 pEt v I xe =, (1) ng dXx 29 where D = axial dispersion coefficient, cmg/sec, C = concentration of uranium in salt, moles/cmB, X = position in column measured from top of column, cm, V = superficial salt velocity, cm/sec, k = reaction rate constant, secfil. The terms in BEgq. (1) represent the transfer of uranium in the salt by axial diffusion, the transfer of uranium in the salt by convection, and the removal of uranium from the salt by reaction with fluorine, respec- tively. The assumption of a first-order reaction does not imply a particular rate-limiting reaction mechanism; however, it is consistent with the assumption that the rate-limiting step is diffusion of uranium in the salt to the gas~liquld interface. In this case, the first-~order expression would imply that the concentration of uranium in the salt at the interface is negligible in compariscn with the uranium concentra- tion in the salt at points a short distance from the interface. The boundary conditions chosen for use with Eg. (1) assume that the diffusive flux across the fluorinator boundaries is negligible: at X = 0 {top of fluworinator), ac V [ '] =-= [ - C 3 (2) Ao D |"feed O+ and at X = & (bottom of fluorinator), dc - =0, (3) dK‘X=2 where Cfeed = concentration of uranium in salt fed to the fluorinator, :CO+ = concentration of uranium in salt at the top of the flu- orinator. 30 e C 1. al tc Note that o+ 18 not equal to Cfeed uranium concentration in the salt at the top of the column where the since there is a discontinuity in salt enters. The solution to Eg. (1) with the stated boundary conditions is, then: Pal ea'X {;a sinh B (X ~ %) + B cosh 8 (X - Qfl c(x) = feed s () 2 (o™ + 32) sinh BL + 20R cosh RL where o is defined as V/2D, B is defined as J&g + LkD/2D, and the ratio of the uranium concentration in salt leaving the column (at X = ) to the concentration in the feed salt is: c(o) 1 C e T - > feed 1/2:mn . 1/21 o5 LJl/h + N - 1/2“] _(1/z*tn. 1/2] &6 E/P +¥1/k +nj (o + bn A J o+ kn () where n o= 52 ve _ L ™D L.h Fvaluation of Fluorination Reaction Rate Constant Application of Eq. (5) to the design and evaluation of continucus fluorinators requires information on the rate constant k and the axial dispersion coefficient D. Values for the dispersion coefficient can be obtained from studies in which air and water were contacted countercur- rently in open bubble columns.5 Results from these studies are shown in Fig. 13 for 1.5~, 2-, and 3-in.-ID columns that were 72 in. long. The dispersion coefficient is independent of the gas inlet diameter over the range tested (from 0.020 to 0.170 in.) and appears to be es— DISPERSION COEFFICIENT (cm?/sec) 100 ORNL DWG 70-11,001 R - A 80 A = 60 Ao B 50 ~PREVIOUS 2-in. COLUMN DATA ° ‘{ * ! -~ 40 \ 2 » . =Rt 30 N e m— T % % s ¢ =0 B 3-in. COLUMN ® ® tn. A 2-in. COLUMN ° ¢ ® {.5-in. COLUMN . . ol—® 1 1114l | L1t ! 2 5 10 20 40 50 60 B8O 100 200 GAS FLOW RATE (cc/sec) Pig. 13. Axial Dispersion Coefficient Data Used in Evaluation of Continuous Fluorinator Performance. e 32 sentially independent of column diameter for the larger column sizes. Only with the smallest column diameter and low gas rates did the data deviate from a single curve. Under all other conditions, the dispersion coefficlent appears to be a function of the volumetric gas flow rate rather than the superficial gas velocity as one might expect. The dats fall into two regions, corresponding to low gas flow rates and to high gas flow rates, respectively. The experimental data from which the rate constant k was evaluated were used with data from the region of low gas flow rates. The flucorinators proposed for MSBR processing are expected to operate in the region of high gas Tlow rates, well beyond the range of the existing data. In analyzing the experimental data from the l-in.-diam open-column continuous fluorinator, it was assumed the axial dispersion coefficients were the same as those measured in a 1-1/2-in.~diam column (the closest column size tested). No correction was made for differences in physical properties between molten salt and water; for this reason, considerable uncertainty may have been introduced into the results. Efforts to improve this situation are under way. A group of MIT Practice School students has studied the effects of viscosity and attempted to determine the effect of interfacial tension on the dispersion coefficient.¥ In~ creases in viscosity result in a decrease in the dispersion coefficient; however, the effects of high interfacial tension and density of the liquid are not known. The reaction rate constant k was estimated from the series of contin- uwous fluorination experiments made by McNeese in a l-in.-diam, 48-in.-long fluorinator by using the mathematical model developed above and an estimate for the axial dispersion coefficient. Two inlet uranium compositions (0.35 and 0.12 mole %) were studied at 600°C, and one composition (0.35 mole %) was studied at 525°C. Three data points corresponding to different salt flow rates were obtained for each set of temperatures and inlet composi- tions. The fluorine flow rate was different for each data polint; however, according to the present model, this flow rate only affects the results by changing the dispersion coefficient. To evaluate the rate constant, 33 k, we chose to use the data obtained at 525°C since the temperature of the molten salt in the proposed fluorinator will be 10 to 15°C above the salt liquidus temperature of 505°C. Application of the model to the three data points obtained afi 525°C gave the results shown in Table 3. Note that the resulting valueg for k are reasonably constant and show no trend with salt or fluorine flow rate. Although these data do not confirm the validity of the present model, they do not contradict the model. Table 3. Summary of Fluorination Results Obtained at 525°C Salt Flow Rate g(z) Fo Flow Rate D k {em/sec) Tfeed (em3/sec) (cm©/sec) (sec™1) 0.0625 '0,0257 6.8 | 18 0.0081k4 0.0445 0.0096 5.0 1k 0.01010 0.0225 0.00L57 3.82 12 0.00943 avg 0.00922 In the mathematical model developed earlier, the effect of axial dispersion was consgidered; however, it is interesting to note how the results are affected when axial dispersion is neglected. The solution to Bg. (1) when D is zero is: C{a) _ ~KL/V Cfeed The three data points considered above produce K values in this case of 0.00188, 0.00170, and 0.000995 secul. The effective rate constants are, as expected, considerably lower than those obtained when axial dispersion is taken into account. An indication that axial dispersion is significant in the present case lies in the fact that the lowest K value was obtained when the salt flow rate was lowest; the effects of axial dispersion would be the greatest for this condition. 3L 4.5 Predicted Performance of Open-Column Continuous Fluorinators The performance of large open-column continuous fluorinators was estimated from Eg. (5) using the previously discussed estimates of the reaction rate constant k (shown in Table 3) and the dispersion coeffi- cient D {shown in Fig. 13). The required fluorinator height is shown in Figs. 1Lk-17 for a range of salt flow rates for four fractional uranium removal values (0.9, 0.95, 0.99, and 0.999). The uranium con- ceatration in the inlet salt was assumed to be 0.003 mole fraction in all cases, and the fluorine flow rate was assumed to be equal to 1.5 times the stoichiometric requirement. The effect of fluorine flow rate 1s important since the axial dispersion coefficient is dependent on fluorine flow rate. ZExtrapolation of the dispersion coefficient data to high gas flow rates results in very high dispersion coefficients for some of the conditions considered. (This extrapolation is believed to be a conservative one.) The results shown in Figs. 15 and 16 are encouraging since they suggest that single fluoriunation vessels of moderate size will suffice for removing uranium from MSBR fuel salt pricr to the isolation of protactinium by reductive extraction. The reference flowsheet for isolating protactinium by fluorination--reductive extraction reguires fluorination of fuel salt at the rate of 170 ft3/day, which represents a 10-day processing cycle. A 6-in.-diam fluorinator having a height of 10 ft would be required for a uranium removal efficiency of 95%, and an 8-in.-diam fluorinator having a height of 1L ft would be re- quired for a uranium removal efficiency of 99%. Fluorinators having a high uranium removal efficiency are re- quired in the production of high-purity 233U since incomplete removal 233 of uranium from a salt stream containing “Pa would result in contam~ 233 ination of the U with other uranium isctopes. Fluorination of salt streams having flow rates of 550 to 1700 ft3/day, with uranium removal efficiencies as high as 99.9%, may be required. As shown in Fig. 17, if a single open-column continuous fluorinator were used, a column 35 ORNL-DWG- 70-1471i-R2 1000 1 YT T T T T T [~ URANIUM REMOVAL EFFICIENCY 30 % ] L. INLET URANIUM CONCENTRATION 0.003 mole fraction - . FLUORINE FEED RATE 150 % of Stoichiometric — 100 }_ - " x o ) x e 5 s & 'O O n = - 2 ™ - 4 in. in, i i FLUORINATOR DIAMETER | il L] i ool i b 413 41 10 100 000 10,000 SALT FLOW RATE, ft/day Fig. 14. Variation of Calculated Fluorinator Height with Salt Flow Rate and Fluorinator Diameter for a Uranium Removal Efficiency of 90%. 36 CRNL- DWG-T70-8998R2 1000~ T T T T 7T T T T T T T T T T B URANIUM REMOVAL EFFICIENCY 95 % N B INLET URANIUM CONCENTRATION 0.003 mole fraction | - FLUORINE FEED RATE 150 % of Stoichiometric o IOOW— B =L FLUORINATOR DIAMETER ’,.m T O . L1 T 14 S 210 —] g ——n Z 7 o ] | ] . w ] [ L i 1 b | ] | ] | ll i ! L1111 10 100 1000 10,000 SALT RATE, ft3 /day Fig. 15. Variation of Calculated Fluorinator Height with Salt Flow Rate and Fluorinator Diameter for a Uranium Removal Efficiency of 95%. 37 ORNL DWG 70-14712R2 1000 [ 1 1T T V11T ' t I 0T T 1 | LR - URANIUM REMOVAL EFFICIENCY 99% — INLET URANIUM CONCENTRATION 0.003 mole fraction - FLUQORINE FEED RATE 150% of Stoichiometric 100 |- L bbbl | FLUORINATOR HEIGHT, ft i3l } FLUORINATOR DIAMETER i | | bl 1 1 1l | L d Lol L) ! Lot 11 1id 10 100 1000 10,000 SALT FLOW RATE, ft¥day Fig. 16. Variation of Calculated Fluorinator Height with Salt Flow Rate and Fluorinator Diameter for a Uranium Removal Efficlency of 99% . 38 ORNL-DWG-70-147I3R2 1000 f T T T T T 7 T T77TT) |~ URANIUM REMOVAL EFFICIENCY 99.9% ’ L INLET URANIUM CONCENTRATION 0.003 mole fraction - - FLUORINE FEED RATE 150 %% of stoichiomstric — i 00— - [ I: e o i | T ?::) = % < S / 8 10 FLUCRINATOR DIAMETER ] i ] | 1 ool 1 ot ] Lol bt 10 100 1000 10,000 SALT FLOW RATE, f13/day Fig. 17. Variation of Calculated Fluorinator Height with Salt Flow Rate and Fluorinator Diameter for a Uranium Removal Efficiency of 99.9%. 39 dismeter of 10 in. and heights of 36 to 60 ft would be required. In this case, the fluorinator would be divided into several open—column fluorinators operating in series. If two columns were used, the re~ gquired column heights would be less than half the height required for a single column since there would be no axial dispersion scross the fluorinator inlets and outlets. As shown in Fig. 18, the fequired uranium removal efficiency for each column would be 96.8% and column neights of 15 to 28 ft would be required for a 10-in.-diam fluorinator. The use of three columng, each of which must have = Q0% uranium removal efficiency, would reduce the total column height even further; heights of 8 to 17 £t would be reguired for a 10-in.-diam fluorinator in this 5. USE CF RADIO-FREQUENCY INDUCTICN EHEATING FOR FROZEN~WALL FLUCRINATOR DEVELOPMENT STUDIES J. R. Hightower, Jr. C. P. Tung L. E. McHeese Fluorinaticn of molten salt for removal of uranium is required a several points in processes being considered for the isclation of prot- actinium and for the removal of rare earths. The fluorinators will be protected from corrosion by a layer of salt frozen on metal surfaces that will potentially contact both fluorine and molten salt. Although the separate aspects of such operations (continuous or batch fluorina- tion, and frozen film formation) have been shown experimentally to be feasible, the testing of a fluorinator protected from corrosion by a frozen wall has been hampered by the lack of a corrosion-resistant means for generating heat in the molten salt. (The decay of fission products in the galt will constitute the heat source in a reactor process- ing plant.) Radio~-frequency induction heating appears to be a suitable method for providing heat in experimental work on fluorinator development, and its use is being studied. The heat would be generated in the molten salt ORNL DWG 70-147i4R2 1000 T T T TTT1 T T TTTTTT ] T T T TTT] Fyot URANIUM REMOVAL EFFICIENCY 96.838 % INLET URANIUM CONCENTRATION 0O.003 mole fraction FLUORINE FEED RATE 150 % of Stoichiometric —d 1 100 FLUORINATOR HEIGHT, ft FLUORINATOR DIAMETER Ll 1Ll ] Lol L L) l oL L1 0 100 1000 10,000 SALT FLOW RATE, ft3%/day Fig. 18. Variation of Calculated Fluorinator Height with Salt Flow Rate and Fluorinator Diameter for a Uranium Removal Efficiency of 96.838%, i (a conductor) by eddy currents induced by an alternating magnetic field. The magnetic field would be generated by a coil not in contact with the molten salt. This method of generating heat in the salt has the disad- vantage that heat would also be produced in the metal walls of the fluori- nator, although this can be minimized by choosing a favorable gecmetry. Two promising coil configurations (see Fig. 19) will be considered in the remainder of this section. In the first configuration (config- uration I), the induction coil is located Just inside the metal wall of the eylindrical Tluecrinator ve?sel and 1s embedded in a frozen salt film on the vessel wall. Heat is:generated inside the coil by the magnetic field, and neither the coll nor the vessel wall would be in contact with molten salt and fluoride. In the second configuration (configuration II), the induction coil would be much smaller in diameter than the fluorinator vessel and would be located at the center of the fluorinator. A coolant would be passed through the induction coil in order to cover it with a layer of frozen salt. Heat would be generated in the molten salt (and in the vessel wall) by the magnetic field outside the coil. The second configuration requires a greater total heat genera- tion rate than the first configuration since a larger area would be covered with frozen salt. 5.1 Mathematical Analysis Initial work on the problem was directed toward a mathematical analysis of several coll configurations in order to assess the feasi- bility of rf heating and to idenfiify important system paremeters. Sev- eral configurations, including the two shown in Fig. 19, were examined. These include a configuration in which the induction coill was located outside the fluorinator vessel wall, which would be relatively thin to permit'power to be transmitted through it. Since it appeared that most of the heat generation for systems having dimensions of interest would occur in the metal wall, this configuration wag not considered further. Configuration T (see Fig. 19) is more amenable to mathematical analysis than configuration II, and expressions are available from standard texts 22,273 on induction heating” for predicting coil performance. ORNL-DWG 70~ 2827 . r/iiiF—lG’U—RjEi_\ c_gggga:e NVT\_‘/CONF iGU%AT:‘ON o \7 /- N ct% giff@? e Vi %& V/ N7 D H \/}j/ Z E \/ Z z |/ b A Y A z D z 2 0 v 2y W s P GEg ; gfiD 17 4y Up N\ ' ‘ N\ Gf A g,/fi:\? s N\ AP Fig. 19. Induction Coill Configurations for Tests of Fluorinators Protected from Corrosion by a Frozen Wall. oft L3 Calculations were made for predicting the performance of two systems: (1) a salt system having the approximate dimensions of an ex- perimental frozen-wall fluorinator that will be built later, and Famn N o S a system using sulfuric acid (which has properties similar to molten salt). Rate of Heat Generation in Molten Salt. — The rate of heat genera- tion in the molten~salt region of a frozen-wall fluorinator was approxi- mated by an equation for the rate.of heat generation in an infinitely long cylindrical charge inside an infinitely long coil. The heat genera- tion rate per unit length of charge is:2 P =1 fle‘*’l2'/109 SN ngaigF, (6) where P = power generated in charge, W/cm, p = permeability of charge, f = frequency, Hz, n = resistivity of the charge, {i-cn, n = coil spacing, turns/cm, a = radius of charge, cm, i = rms coil current, A. The factor F in Eq. (6) is defined by the following equation: p = ber(d) ver'(A) + bei(a) bei'(a) | (7) ) (ber(A))2 + (beil(A )2 where A is defined as aJuf/éSSB J;r and ber, ber', bei, and bei' are Bessel functions. Ly Heat Generatlion Rate in Fluorinator Wall. — No egquatioms for cal- culating the heat generstion rate in a cylindrical shell surrounding a cylindrical coil were found in the literature. Therefore, we estimated the heat generation rate in the fluorinstor wall by assuming that the heat generation rate in the pipe Just outside the coll was the same as that in a pipe Just inside = coil; the intensity of the magnetic field was assumed to be the same at the surface adjacent to the coil in each case. For the high frequencies we expect to use (f >300 kHz), the penetration depths are very small and a metal pipe would be heated at essentially the same rate as a s0lid cylinder. For this case, the term F in Eq. (7) approaches the value of Y2/2, and the equation that approximates the rate of heal gen- eration in the pipe is as follows: P = (n?/N10%) Ju_en 0% 5%, (8) - P = heat generated in pipe, W/ cm, u_ = permeability of pipe, n_ = resistivity of pipe, {i-cm, R = inside radius of pipe, cm. Equation (8) will be used only to obtain a rough estimate of the heat generation rate in the pipe; the actual heat generation rate will be measured experimentally. Resistance lLosses in the Induction Coil. — The resistance of the . ‘o . . . 2 induction coil is given by the following equation: 2 _ ~ 1/2 2 -8 R, = 63'2Kr (fnc) d,n” x 10 7, (9) where o = resistance of the coil, § per em of axial length, ;= correlation factor (assumed to be 1.5), f = frequency, Hz, n, = resistivity of the coil, ufi-cm, dc = diameter of the coil, in., n = coil spacing, turns/in. The resistive losses in the coll are given, then, by: P, = iERC (in W/em). (10) Conduction of Heat Through a Film of Frozen Salt. — It was assumed that all heat generated in the molten salt in the center of the fluori-~ nator is transferred by conduction through the film of frozen salt on the fluorinator wall, which is maintained at a temperature below the ligquidus temperature of the salt. The equation relating the hesat transferred through the frozen film to the temperature difference, the dimensions of the film, and the properties of the frozen film is: 2flks (Ti - TC) P = TRE (11) P = rate at which heat is transferred through the frozen film, k¥ = thermal conductivity of the frozen film, a = radius of the molten salt--frozen salt interface, R=a+ t, where t is the thickness of the frozen salt (R is assumed to be the inside radius of the coil), T, = the temperature of the solid-liguid interface, i.e.,, liquidus temperature of the salt, T = temperature in the frozen film at the outside of the coil. L6 5.2 Calculated Results for a Molten-Salt Fluorinator The fluorinator system that was examined had the following features: The salt was LiF-BeF -ThF) (68-20-12 mole %); the inside diameter of the 1/4-in.~thick vessel was 4.9 in.; the mean diameter of the induction coil was 3.9 in.; the frozen salt extended in from the wall, covering the coil and leaving a 1.9~in.-dlam molten core; and the temperature difference across the frozen salt layer was 100°C. The metal was assumed to have the same electrical properties as Monel., The heal generation rate in the mol- ten salt necessary to maintalin this frozen salt layer was 63 W per centi- meter of fluorinator length. The caleulated induction current in the coil (at an assumed frequency of 500 kHz ) necessary to procduce this heat gen- eration rate in the salt was 24.7 A. Using the assumption given previously, the calculated generation rate in the fluorinator vessel wall was 65 W/em (about 1.05 times the heat generated in the salt). Removal of this amount of heat generated in the vessel wall, as well as that generaled in the salt, would be practical. For a 5~-ft-long fluorinator vessel, these cal- culations indicate that a 28~kW rf generator would be required. Similar calculations could not be made conveniently for configuration IT. 5.3 Experimentally Measured Heatl Generation Rates In order to verify the values predicted for configuration I and to evaluate the performance of configuration II, we carried ocut experiments in which a 29 wt % HESOM solution was substituted for molten salt. Heat generation rates were measured in the acid and in a Monel pipe surround- ing the acid. The acid was contained in a 2-liter graduated cylinder having an inside diameter of 3-1/4 in. The pipe was a 6-in.-long section of 6-in. sched 40 Monel. The coil for the test with coufiguration I was L in. in inside diameter by 6 in. lcng and consisted of 20 turns of 1/h- in.~diam copper tubing; it was placed arcund the acid container, and the 6~in. pipe was placed around the coil. The coil for configuration II was 1-1/% in. in ocutside diameter by 6 in. long and consisted of 20 turns of 1/4~in.~diam copper tubing. It was placed im a 1-3/8-in.-0D glass tube, LT which was immersed in the center of the acid. The healt generation rates were obtained by measuring the rates of temperature increase in the acid and in the pipe. In three runs with configuration I, the ratio of heat generated in the pipe to that generated in the acid averaged 1.3. The frequency in the test was 350 kHz, and the conductivity of the acid was about 0.75 mho/cm. The heat generation rate in the acid was about 19 W per centi- meter of coil length; this was kept low to prevent the acid from boiling. The predicted value for the ratio of heat generated in the pipe to that generated in the acid was 0.58 (this value was calculated by using the properties, conditions, and dimensions of the experimental system, and by making the assumptions that were outlined in the previous section), which is approximately one~half the measured value. Deviations in heat generation rate of this magnitude between the predicted and measured values ére not surprising since the method used for calculating heat generation in the pipe was an approximate one. In five runs with configuration IT, the ratic of heat generated in the pipe to that generated in the acid had an average value of 0.069. The frequency in this test was LL4O kHz. The heat generation rate in the acid, using configuration II, was about 16 W per centimeter of coil length, even though a higher plate voltage was used than in the experiments with con- figuration I. This indicates that the coupling of the magnetic fleld with the acid was poorer with configuration IT than with configuration I; such a result is to be expected since the magnetic field strength outside a coll is smaller than the strength inside a coil. However, it means that, even though only a small amount of heat was generafied'in the pipe wall with configuration 1I, the efficiency of heating the salt could still be much less than if configfiration I were used, because the heat generation in the coil itself might be large compared with the heat generation in the salt. This aspect will be explored in future experi- ments when means to measure the coil current have been obtained. L8 In general, results of these prelimiunary experiments and calcula- tions encourage us teo believe that inductive heating is & reasonable method for supplying heat to an experiment designed to study a fluori- nator containing a frozen wall but no internal heat generation result- ing from fission product decay. Disagreement between the measured and The calculated values of relative heat generation rates indicates that the design of an experimental fluorinatcr system must rely heavily on empirically cbtained relationships. An experimental program to obtain this information Is under way. 6. MSRE DISTILLATION EXPERIMENT J. R. Hightower, Jr. L. E. MclNeese Effective relative volatilities, with respect to LiF, of BReF = 2’ and of the fluorides of 952r, lthe, lh?Pm, $55Eu, le, 9081‘3 89Sr, and 137 ZI"Fh, Cs have been calculated from condensate analyses made during the MSRE Distillation Experiment. These results, which have been discussed pre-~ viously,26 show that all of the components excep: BeF,. and Zth had ef- fective relative volatilities that deviated (drasticaily in some cases) from values predicted from tests with equilibriur systems. Possible causes for the discrepancies include: (1) entrainment of droplets of still-pot liquid in the vapor, (2) concentration zradients in the still pot, and (3) contamination of samples during their preparaticn for radio- chemical analysis. Entrainment was suspected for a number of reasons. For example, entrainment of only 0.023 mole of liquid per mole of vapor would account for the high relative volatilities calculated for the slightly volatile fission products lthe, lher, le, and 90 Sr. Entrainment rates of this order would not be reflected in the effective relative volatilities of more volatile materials (a > 1). The high correlation of the scatter of the calculated effective relative volatilities of different slightly volatile fission products is consistent with the hypothegis that en- trainment occurred. 49 Since entrainment was not apparent in the nonradiocactive operation of the still,zT reagons for entrainment in the radiocactive operation were sought to support the hypothesis. Evidence of a salt mist above the salt in the pump bowl at the MSRE and above salt samples removed from the MSRE 28,29 Further, studies have indicated that these mists has been reported. are present over radicactive salt mixtures but not over nonradiocactive mixtures; however, examination of data from these studies showed that entrainment rates large enough to explain the results of the MSRE Dig- tillation Experimentrcould be obtained only by assuming that the con- centration of salt in the gas space above the salt during this experi~- ment was equal to that observed above salt in the pump bowl at the M3IRE. If the rate at which the mist is formed decreases as the power density in the liquid decreases, the concentration of salt in the mist should - also decrease as the power density in the liquid decreases. B8Since the salt used in the MSRE Distillation Experiment had a much lower power density (L00-day decay period for distillation feed as compared with less than 30-day decay period for salt samples tested for mist formation) than salt samples from the MSRE, it seems unlikely that the concentra- tion of salt in the gas above the salt would have been high enough to explain the high relative volatilities for the slightly volatile fission products. In addition to the argument against the entrainment hypothesis given above, not all discrepancies would be explained by it. For example, 898r/9o Sr activity ratio 1374 S it would not account for the variations in the and for the low value for the effective volatility of Concentration polarization would cause thé effective relative vol- atilities of the slightly volatile materials to be greater than the true relative volatilities. As the mofeuvolatile materials vaporized from the surface of the liquid, the slightly volatile materials would be left behind on the surface at a higher concentration than in the liquid just below the surface. The concentration of these slightly volatile materials in the vapor would then increase since further vaporization would occur from a ligquid with a higher surface concentration of slightly volatile materials. Since effective relative volatilities were based on average 50 concentrations in the still pot, the concentration in the vapor would be higher than that corresponding to the average concentration in the liquid; also, the calculated effective volatility would be higher than the true relative volatility. Concentration polarization would cause the effective relative volatility to be lower than the actual relative volatility in the case of a component whose relative volatility is greater than 1. The extent to which concentration polarization affects the effective relative volatility of a particular component depends on the dimensionless group D/vL, which qualitatively represents the ratio of the rate of dif- fusion of a particular component from the vapor-liquid interface into the bulk of the still-pot liquid to the rate at which this material is trans- ferred by convection to the interface by liquid moving toward the vapor- ization surface. 1In this ratio, D is the effective diffusivity of the component of interest, v is the velocity of liquid moving toward the interface, and L is the distance between the interface and the point where the feed i3 introduced. The occurrence of concentration pelarization is suggested by the sharp increase, at the beginning of the run in the effective relative volatilities of lthe, luYPm, 155Eu, and, possibly. of 91Y and 9OSI‘. This increase would correspond to the formation of a concentration gradient in the still-pot liquid. The effective diffusivities of NdF3 in the still pot, calculated from results of the nonradiocactive experi- ments, ranged from 1.4 x 10“”1!r to 16 x lO”u cmg/sec and form the basis for estimating the magnitude of the concentration polarization effect in the radiocactive coperation. During the semicontinucus operation at the MSRE, the liquid velocity resulting from vaporization averaged 2.2 X lO“JJr cm/sec; the depth of liquid above the inlet was about 9.4 om. 1f -one assumes that the effective diffusivities of the fission products in the still pot during the MSRE Distillation Experiment were in the same range as they were during the nonradiocactive tests, the observed relative volatilities of the slightly volatile materials would be only 2.0 to 18 times the true relative volatility and the observed relative volatility of 13705 would be 0.011 to 0.021 times its true value (as- o1 suming, in each case, that the true relative volatilities are those given in refs. 27 and 28). Although concentration polarization may have been significant in the work with radiocactive salt, the effect was not great enough to account for the discrepancies bhetween the observed relative vol- atilities and what we consider to be the true values. Also concentration polarization would not explain the variation in the 8QSI'/Q(‘JSI' activity ratic between samples of condensate. The possibility that the condensate samples became contaminated while they were being prepared for radioactive analysis 1s suggested by the wide variation in the value of the 898r/908r activity ratio. Al- though routine precautions against contamination were taken in the hot cells, where the capsules were cut open, no special procedures were used and the manipulators used to handle MBRE salt samples were also used to open the condensate samples, If it 1s assumed that the source of the contamination was the last salt sample taken from the MSRE before the distillation samples were submitted, only 10"6 to 107 g of salt per gram of sample would be required to yield the observed values of the 89 9 Sr and OSr activities. Contamination from such small quantities of material would be extremely difficult to prevent. Other observatlons explained by assuming that the samples were con- taminated are the high relative velatilities of the slightly volatile fission products and the close correlation between the variations of calculated relative volatilities of different fission preducts. However, 137 the low relative volatility for Cs is not explained by this hypothesis. We conclude that, although several factors may be involved, the dis- crepancy between the effective relative volatilities of the slightly vol- atile materials measured in this experiment and the values measured pre- viously is primarily the result of contamination of the condensate samples by minute quantities of other MSRE salt samples in the hot cells. 52 7. DEVELOPMENT OF THE METAIL TRANSFER PROCESS E. L. Youngblood W. F. Schaffer, Jr. L. E. McNeese E. L. Nicholson J. R. Hightower, Jr. Rare earths have been found to distribute selectively into molten LiCl from bismuth solutions containing rare earths snd thorium, and an improved rare-ecarth removal process based on this observation has been devised. Work that will demonstrate agll phases of the improved process, known as the metal transfer process, is under way. f.1 Equipment and Experimental Proceduare Equipment has been fabricated for study and demonstration of the metal transfer process for selectively removing rare earths from single-~ fluid MSBR fuel salt. The first series of experiments will be carried out in a 6-in.~-diam compartmented vessel (Fig. 20) fabricated from carbon steel. The vessel is 24 in. high; and the internal partition, which divides the vessel into two equal-volume compartments, terminates 1/2 in. above the bottom of the vessel. During each experiment, thze vessel will contain a 2~in. depth of bis- muth (i.e., v 0.8 liter) that is saturated with thorium at the operating temperature of 640°C and two sa’*t phases having depths of 3 to U in. (0.7 to 1.0 liter each). One salt phase will be MSBR fuel carrier salt (72-16~ 12 mole % LiFmBngmThFh) initially containing 0.3 mole % LaF 147 and about 5 3 mCi of Nng; the other salt phase will be LiCl. The LiCl compartment also contains an electrically insulated cup containing a Li-Bi solution and a pump for circulating LiCl through the cup. The cup will contain 3 about 200 em” of Bi~Li solution having a lithium concentraticn of 0.L0O mole fraction. Provision is made for agitating the liquid phases and for sampling all phases. During operation, La and Nd will be transferred from the fuel salt to the LiCl by circulation of the Th-Bi solution that will alsc contain the rare earths. Fractions of the La, Nd, and Th will then be extracted from the LiCl by contacting the LiCl with the Li-Ri solution. ORNL- DWG-70-4504RI S 4—LEVEL CONTROLLER LEVEL ELECTRODES ARGON ,]———. VENT VENT — 4 Hg sUBBLER 1 —QUARTZ CARBON-STEEL—] &1 PUMP PARTITION | ™ 6-in. CARBON- ~" STEEL PIPE 24 in. 8L —LicCl 72-16-12 MOLE Y% —~_ FUEL CARRIER SALT D ~ gl _—Li-Bi Th"Bl"\ ::':-: ¥ Fig. 20. Carbon~Steel Vessel for Use in the Metal Transfer Experiment. To begin an experiment, the three phases in contact (fuel salt, bismuth, and LiCl) will be allowed to approach equilibrium and samples of the phases will be taken. The LiCl will then be pumped through the reservoir containing the Li~Bi sclutlion at a flow rate of about 25 cm3/ min in order to remove the rare ecarths and thorium from the LiCl. The LiCl will overflow the Li-Bi reservoir and return tc the initial LiCl volume. After a period of about 3 hr, circulation of the LiCl will be stopped and the system will be allowed to approach equilibrium. Samples of the phases will then be taken. It is estimated that 5 to 10% of the lanthanum initially present in the fuel salt will have been transferred to the Li-Bi solution at this point. The above seqguence of operations will then be repeated until the desired fraction of the lanthanum has been transferred to the Li-Bi solution (50 to 90%). The lhTNd tracer will be used to give a rapid indication of the rate at which the rare earths are transferred during the experiment. T.2 Development and Testing of a Pump for Circulating LiCl Several pumps for circulating the LiCl have been fabricated and tested for use in the experiment. The first pump, made from 1-1/2-in.- diam quartz tubing, was tested in a b-in.-diam quartz vessel, as shown in Fig. 21. The pump was driven by varying the difference in argon pressure between the inside and the outside of the pump chamber. The direction of flow was determined by weighted quartz check valves on the pump inlet and outlet. In operation, the flow of argon was controlled by solenoid valves that were actuated by signals from two level probes in the punp tube. During the first part of a pump cycle, the vessel was pressurized with argon, which forced liguid into the pump chamber through the pump inlet until the liquid contacted the high-level probe (Fig. 21). When contact with the high~-level probe weas made, the solenoid valves actuated to vent the test vessel and pressurize the pump chamber; this forced liquid out through the pump outlet. When the level of the liguid in the pump chamber | dropped below the low-level probe, the solenoild valves were again actuated and the cycle was repeated. 55 ORNL PHOTO 98323 CONDENSER FOR LiCl VAPOR 4-IN.-DIAM QUARTZ TEST VESSEL MAGNESIA PROBE SUPPORT SR GROUND HIGH-LEVEL PROBE LOW-LEVEL PROB 1-1/2-IN.-OD R QUARTZ PUMP TUBE PUMP OUTLET . [ B | L QUARTZ = CHECK VALVES 2 ey _ PUMP INLET R k* Fig. 21. Pump for Circulating LiCl, Installed in 4-in.-diam Quartz Test Vessel. 56 Figure 2] shows the pump in operation with a colored agueous solution of LiCl. During tests with this mixture, we observed that the quartz check valves had a tendency to stick. Before testing the pump with molten LiCl, we added supports to the weighted quartz check valves (which were teardrop- shaped) to prevent them from tipping over. This modification was not entirely successful; the valves continued to stick, although not as frequently. After modification, the quartz pump was tested with molten LiCl. The h-in.-diam quartz vessel containing the pump was loaded with 800 g of LiCl that had not been previously dried, and the system was heated to 650°C in order to melt the LiCl and remove water from the salt. The pump and the gas space of the vessel were purged with argon during this period. During the heatup period, the presence of a liquid having a pH of about 2 was observed on the vessel walls. When the pump was lowered into operating position, we found that the level probes had shorted; this shorting was probably caused by an accumulation of liquid on the magnesia probe support. The pump was then withdrawn from the molten LiCl, and the system was cooled to room temperature in order to recover the pump. Examination revealed that the quartz vessel and pump had sustained severe damage in regions contacted by the vapor; however, the quartz surfaces in contact with molten LiCl showed no sign of attack. Two quartz pumps, similar to the first but having sapphire check valves, were then fabricated and tested with molten LiCl at 650°C. 1In the test involving the second pump, the LiCl was air-dried at 225°C. The pump cperated successfully for a few hours, but difficulties with short circuits in the level probes prevented continued operation. The quartz components of the system showed evidences of gradual deterioration. After 7 days of exposure, the quartz was badly etched and scme areas appeared to be ready to disintegrate. However, the sapphire balls used as the check valves were in good condition. The electrical short circuits were apparently caused by a film of LiCl covering the insulators used to separate the electrical probes. The third guartz pump was tested in LiCl that had previously been purified by contact with a Th-Bi solution at 650°C to remove oxides. A molybdenum cup containing the Th-Bi solution was placed in the bottom of o7 the pump test vessel to further purify the LiCl. This pump operated satisfactorily at rates of 30 to 125 ml/min for 16 days, after which it was disassembled for inspection. At the time of shutdown, the pump was still operating satisfactorily; however, inspection of the pump and the quartz vessel showed evidence of considerable attack of the quartz in the hot vapor region. The quartz components that were submerged in the liquid, the molybdenum components, the carbon-steel thermowell, and the sapphiré balls used as check valves all appeared to be in reasonably good condition. Discussions with quartz manufacturers and with those who have had experience with quartz equipment revealed that devitrification (i.e.., a change in the crystal structure of the quartz) is a difficulty commonly encountered with this material at high temperature. The rate of devitrification is accelerated by the presence of contaminants (such as LiCl) on the surface of the quartz. The full effect of devitrification is not apparent until the quartz has cooled to room temperature. This appears to be the type of attack in our experiments. When purified LiCl is used, the rate of devitrifilcation appears to be sufficiently low to permit a quartz pump to be used for the metal transfer experiment. A pump using captive bismuth pools as check valves has been fabricated of low-carbon steel. This pump will be tested with molten LiCl in the near future. 8. ELECTROLYTIC CELL DEVELOPMENT: STATIC CELL EXPERIMENTS J. R. Hightower, Jr. M. 5, Lin L. E. McNeese We repeated an earlier experiment,3o which was carried out in an a2ll- métal cell to determine whether the presence of quartz in static-cell tests was involved in the formation of black material in the salt. The cell vegsel was fabricated from an 18-in. section of mild-steel, sched L0 pipe. One electrode was a 15~kg pool of‘bismuth; the other was a 1/h-in.-diam mild-gteel rod located at the center of the vessel and placed 1/2 to 1 in. above the bismuth pool. Observations in the salt phase were made using the bismuth surface as a mirror to reflect light to a sight glass in the top flange of the vessel. The electrolyte, a mixture of LiF-BeF, (66-34 mole %), 58 filled the vessel to a level about 3 in. above the bismuth surface. The bismuth had been sparged with Hg at 600°C to remove oxides. The cell vessel was electrically isolated from the hood and from the gas and water supply lines in order that the bismuth pool (and hence the cell vessel) could e operated in an anodic manner; the resistance between the cell and ground was 2 X 106 Q. A de voltage of 2.3 V impressed across the electrodes for about 1 min, with the iron rod located 1/2 in. above the bismuth (the iron rod was cathodic), produced a current of about 16.8 A; when the electrode was raised to 1 in. above the bismuth surface for about 2 min, the current decreased to 15.6 A. An increase in voltage to 2.5 V produced a current that increased from 19.0 A to 21.8 A over a period of 7.5 min; the cathode current density during this time was about 1.8 to 2.0 A/cmg. The total charge transferred during the run was 12,400 C. The cell temperature was 550°C. During the passage of electric current, there was no sign of the dark material that had previously been seen; during the cell operation, density gradient patterns were visible, indicating that convective mixing was taking place in the salt phase. After the operation had been completed, the salt appeared to have a more distinct green color than before. Although the salt was slightly turbid, it was still quite transparent. During the operation of the cell, material was deposited on the iron cathode. BSome of the material was metallic and was probably not as dense as the salt since it formed near the salt-gas interface and tended to gpread out over the surface of the salt. The remainder of the material was black and nonmetallic, and formed more uniformly over the submerged part of the steel electrode. The deposited material (16.5 g) had the composition 9.0 wt % Be, 17.0 wt % Ti, and 71.8 wt % F, and contained traces of Fe and Bi. The formation of such a depcsit is explained as follows. As Bel, was reduced at the cathode during the cell operation, 2 the concentration of BeF2 in the salt in the vicinity of the cathode decreased. At the operating temperature of 550°C, LiF began to crystallize when the BeF2 concentration reached 30 mole %. Further reduction of BeF2 was accompanied by precipitation of sclid LiF in the vicinity of the cathode. When the cathode was removed from the cell, some of the LiFmBeFE mixture was carried with the reduced beryllium. If one assumes that the Ll was present as LiF and that the remaining T was associated with Be as BeF,_, the cathodic depcsit contained about » 0.5 g of Be metal (33% of the Bi_present in the deposit) or 0,11 gram equivalent of Be. Since 0.129 electricsal squivalent was passed, this represents a current efficiency of about 85% for beryllium reduction. The current efficiency was probably actuslly closer to 100%; however, the additional beryllium was not recovered with the cathodic deposit but floated away from the cathode on the surface of the salt. BSome particulate material was noted on the salt surface when the cell was dismantled. The fact that the salt in this test remained transparent indicates that quartz may be important in the formation of black material in the salt. However, the lack of a bismuth cathode into which lithium could have been reduced may have resulted in a system too different from previous cells to allow us to draw firm conclusions. 9. STUDY OF THE PURIFICATION OF SALT BY CONTINUOUS METHODS R. B. Lindauer .. BE. MclNeese To date, the molten salt required for development work as well as for the MSRE has been purified from harmful contaminants (mainly sulfur, oxygen, and iron flucride) by & batch process. The cost of salt from this process (containing natural lithium) has averaged about $l600/ft3; less than 407 of this cost is due to the cost of materials. TIn April 1968, =a commercial vendor submitted a bid of $2660/ft3 on 8 280-ft° quantity of salt. Tt is believed that the labor cogts associated with salt purificstion can be reduced considerably by use of a continuous process for the most time~consuming operation, which is the hydrogen reduction of iron fluoride. Although the removal of sulfur and oxygen by the batch process is fairly rapid, there would probably be advantages to performing this operation in continuous equipment also. A packed column in which molten salt and gas can be contacted counter- currently is being installed (ecell LB, second floor of Bldg. 4505) to obtain data that will provide the basis for the design of a full-scale continuons salt purification facility. Flooding studies will be made using argon and hydrogen with two different sgalt mixtures: LiFwBer (66-34 mole %) 60 and LiFmBngmThFu (72-16-12 mole %). Studies of iron reduction and oxide removal will also be made with both salts. ©Since beryllium fluoride with a low sulfur content is now available, no sulfur removal work is planned. 9.1 Previous Work on Salt Parification All of the molten salt used in the molten-salt reactor projects at ORNL up to the present time has been prepared in small (2~ft3) batch equipment.31 After the raw materials have been blended and melted, the sulfides and oxides are removed simultaneously by a H,~HF sparge at A00°C. The temperature of the salt is then increased to sbout T700°C, and the fluorides of Fe and Ni are reduced by contacting the salt with hydrogen. Next, the salt is passed through a sintered nickel filter. Iron fluoride reduction, the most time-consuming of the various steps, reguires about four days of contacting with hydrogen for a 2-~ft3 batch, Completion of reduction of the metal fluorides is determined by titration of a sample of the effluent gas for HF content. Consideration of continuous methods for use in the purification of molten salt mixtures began in 1967 with a preliminary conceptual design of a semicontinuous pilot plant by the MIT Practice School.32 In a ten- tative design that was proposed for a 0.25~ft3/hr pilot plant, packed columns were suggested for the hydrofluorination and reduction steps. A detailed design of the plant was not possible because of the lack of rate data for the steps involved. During the following year, two series of 33,34 experiments were made by the MIT Practice School for studying the reduction of ironm fluoride in molten-salt mixtures by countercurrent con- tact with hydrogen in a packed column. Although only a few runs were made and equipment performance was not completely satisfactory, significant reduction of the iron fluoride was shown. The results suggested that the reduction rate is Tirst order with respect to iren fluoride concentraticn and that the rate is controlled by the amount of interfacial area that is present between the salt and the hydrogen. It was recommended that additional studies be made to establish the rate-controlling mechanism more firmly. It was also suggested that (1) more effective purification of the hydrogen is necesssry, (2) future runs should be made with packing 61 materials other than York Demister mesh, and (3) the column should be operated close to the leading or flooding polnt. 9.2 Experimental Hguipment A simplified flowsheet for the present experimental eguipment is shown in Fig. 22. Molten salt is fed to the column by pressurizing the feed tank with argon at a controlled rate. The salt flow rate is set by the srgon flow rate and is determined by the rate of depletion of zalt from the feed tank. The hydrogen is preheated before entering the column by a 1.5-in.~diam, 24-in.-long heater filled with 1/k-in.~diam nickel spheres. The salt from the column flows through a gas-seal loop and =a salt filter consisting of a removable 2.75 by 1l2-in. nickel Feltmetal fiber metal cartridge with a mean pore gize of 50 u. The salt passegs through a flowing stream sampler before entering the recelver tank. The gas stream leaving the column can be throttled 1f necessary to depress the salt-gas interface to a point below the bottom of the column. A sample of the gas stream can be withdrawn for continuous determination of the H,0 and HF contents. The stream tThen passes through a NaF trap and an agsolute filter. Packed Column. — The column was fabricated from 1.25-in.-diam low- carbon nickel pipe and can be operated at temperatures up to 750°C. The packed section is 81 in. long and contains 1/4 x 1/4 x 1/16-in. nickel Raschig ring packing. There are deentrainment sections of 3~in.-diam pipe at each end, and a 2-ft-long section of 1/2~in.-diam pipe below the column for deentrainment of hydrogen in case the interface is depressed by a high pressure drop through the column. A differential-pressure transmitter having a range of 0 to 50 in. HBO is connected to the gas inlet and outlet lines to provide data on liquid holdup and flooding. In the absence of gas flow through the column, the salt-gas interface is located 3.4 in. above the hydrogen inlet; the interface can be depressed below the inlet by increeasing the pressure at the top of the column or by an increase in the pressure drop across the column. The column is heated by resistance heaters having a total heating capacity of 3 kW. Salt Feed and Recelver Tanks. — The feed and receiver tanks have, in each case, an inside diameter of 10.25 in., a height of 15 in., and a volume of 21 liters. The 1/bL-in.-thick walls and the l-in.-thick flat ends are 62 ORNL DWG 70-11048 R BACK PRESSURE CONTROL VALVE 3 4 WATER FLUORIDE ANALYZER ANALYZER I ] ARGON S COLUMN : i 1 i | ! FILLAME | ARRESTER ' e VENT SALT e ED . ABSOLUTE TANK | FILTER HYDROGEN (& HF) FILTER g% SAMPLER I RECEIVER Fig. 22. Studies. Simplified Process Flowsheet for Molten-Salt Purification 63 made of low-carbon nickel. These process vessels require approval for use as pressure vessels since their diameters are greater than 5 in. and the operating pressure is greater than > psig. (The vessels have been approved by the ORNL Pressure Vessel Committee for use at 25 psig and 650°C.) Heaters on the column feed line are used to heat the salt to a temperature of 700 to T50°C bvefore it enters the column. Pressure relief valves on the argon and hydrogen supply systems are used to limit the maximum pressure to 25 psig. The tanks are heated by electrical resistance heaters eqguipped with temperature confirollers to permit unattended heatup. Each tank has a bubble~type liquid-level instrument and both well~ and surface-mounted thermocouples. Salt Bamplers. = Samplers are provided on the salt feed tank and in the line exiting from the salt filter. These samplers consist of a ball valve through which a sample capsule can be lowered into the liquid. The capsule is 3/16 in. ID and 1.5 in. long, and has a 25 u metal fiber filter attached to the bottom. Vacuum 1s applied through a capillary tube attached to the top in order to obtain a filltered salt sample. Off-Gas Analyzers. — A sample of the column off-gas stream can be passed through a water and/or a hydrogen fluoride analyzer at the rate of about 5 cm3/min, Flow through these instruments is obtained by a small "Dyna~Vac" gas pump. The water analyzer is the same instrument that was used in the MSRE Fuel Processing Facility.35 Hydrogen flucride is removed from the diluted sample by a sodium fluoride trap before pas- sing through an electrolytic moisture cell. A reference leg with a similar trap and cell provides a compensated system with & single readout. The in- strument is quite sensitive, and the injection of gas containing 250 ppm of water produces a rapld response. The hydrogen fluoride monitor36 has a separate sample diluter with a heated capillary tube that withdraws a 5mcm3 sample each minute; the sample is diluted with argon to produce a flow rate of 1000 cmg/min. The diluted sample is scrubbed with an acetic acld solution in the moni- tor, and the solution is analyzed by use of an aluminum-platimum electrol- ysis cell. 6l 9.3 Gas Bupply and Purification Systems Argon. — Argon is supplied to a purification system from a six-cylinder manifold. Oxygen is removed by a 6-in.-diam, 24-in.~high trap charged with Dow-Q-1 (copper-coated aluminum), which has & total oxygen capacity of T ml (STP) per gram of packing. The oxygen capacity at 100 times the maximum argon flow rate expected {maximum expected rate, 6 liters/wmin) is 1 ml of cxygen per gram of absorbent, with a removal efficiency of 98%. The expected oxygen concentration in the unpurified argon is about 10 ppm; nence, the rurification system has an expected life of at least U months and should reduce the oxygen concentration in the purified gas to about 0.02 ppm. The oxygen content of the argon is measured by a Teledyne Model 306 W Oxygen Trace Analyzer, which uses a wet galvanic cell with silver- lead electrodes. The analyzer can be used for determining the oxygen concentration in either the purified or unpurified argon. Water is removed from the argon by one of two Molecular Sieve <{raps that are arranged in pavallel to allow operation during regeneration of one of the traps. The traps consist of a L2-in.-long packed section con- tained in the annulus between a 5-in.-diam pipe and a 2~in.-diam pipe that surrounds the heaters used for regeneration of the traps. Each Lrap con- tains about 0.k ft3 of type A Molecular Sieve and can reduce the water concentration in a 6-liter/min argon stream from 100 ppm to 1 ppm for a period of 300 days. The water content of the argon is measured before and after purification by a Panametrics model 1000 hygrometer, which uses an aluminum oxide sensor placed directly in the flowing gas stream. Hydrogen. — Hydrogen is supplied to one of two purification systems from a four-cylinder manifold. One purification system, consisting of a Serfass Hydrogen Purifier (palladium membrane), will provide hydrogen at flow rates up to 18 liters/min. In the other purification system, which is used for higher flow rates, the hydrogen is passed through a Deoxo unit, vhere oxygen is converted to water, and a Molecular Sieve trap with a capacity similar to that of the trap on the argon supply. The Deoxo unit has a diameter of 2.5 in., a length of 12.5 in., and a rated maximum capacity of 50 liters of hydrogen (maximum oxygen content, 3%) per minute. Hydrogen Fluoride. - Hydrogen fluoride is vaporized from a tank having a 35~-1liter working volume. The tank is heated in a water bath, and 65 the pressure in the tank is set by controlling the bath temperature. The vaporized HF flows through a cubicle, maintained at 100°C, that containsg a pressure transmitier, an HF flow:control valvé, four capillary tubes and a différential~pressure transmitter for determining HF flow rate, and o Hastings mass flowmeter. The Hastings flowmeter has a maximum tlow rate of 1000 cmB/min, while the capillaries have maximum flow rates of 250 to 2500 cmg/min. The flow of HF can be terminated from three different locations in the operating area in case of an:emergency‘ sulfur is removed from the vaporized HF by passing it through a 2«in.- diam pipe packed with nickel wool. An 18-in.-long section of the pipe contains tightly compressed wool and is heated to 650°C. A 2bh-in.-long section of nickel wool located downstream of the heated section is used to remove particulates from the gaseous HF. 9.4 Installation of Hguipment and Initial Checkoud The equipment (shown in Figs. 23-25 before addition of thermal insulation) was mounted in a 30 in. x 30 in. x 13-ft-high frame for installation in cell 4B on the second floor of Bldg. 4505. Instruments for measuring the liquid levels in the salt feed and receiver tanks, the pressure at the top of the column and above the salt filter, and the differential pressure across the column are also located in the cell. In addition, the cell contains off-gas filters, in-line instruments for measuring the HF and the H?O contents of the off-gas stream from the column, and a sodium‘fluoridé trapafor disposal of excess HF. The equipment is operated from the area just outside the cell on the second floor, where the rotameters are located for controlling the hydrogen flow rate and the argon flow rates fcy:purges and pressurizationZOf the salt feed tank. The panelboard is chown in Fig. 26. A separate panelboard, shown in Fig. 27, contains heaber controllers for the equipment , as well as the recorders and indicators for témperature, liquid level, énd pressure., After completion of the piping but prior to application of insulation to the éystem, the eéuipment was leak tested at 15 psig and room temperature, using helium and a thermal conductivity leak detector capable of detecting a leak rate of less than 0.001 em3/sec. The leak rate for the final system was aboul 0.07 cmB/sec after all detectable leaks had been repaired. 66 ORNL PHOTO 98475A sy PE-ENTRAINMENT SECTION , ABSOLUTE 7 FILTERS - FEED TANK eSfERRERte N - 11401 kll'l“fillfl.i Hl Fig. 23. Top View of Salt Purification Equipment Before Addition of Thermal Insulation. 67 ORNL PHOTO 98473A N FLOWING STREAM . SAMPLER s T . AND o ' FLUORIDE ) ANALYZER SALT FILTER SALT RECEIVER Fig. 24. ILower View of Salt Purification Equipment Before Addition of Thermal Insulation. 68 ORNL PHOTO 98476A SOLIDS ¥ ADDITION FUNNEL — PACKED COLUMN P SAMPLER il - & FEED TANK N") Fig. 25. ©Salt Purification Equipment Before Addition of Thermal Insulation. 69 PHOTO 98474 e D Fig. 26. Panelboard for Salt Purification Equipment. Temperature, level, and pressure recorders and heater controls are included. Fig. 27. tion Equipment. Gas Flow Panelboard and Cell Entrance for PHOTO 98472 Salt Purifica- 0L 71 9.5 Anticipated Experiments and Operating Procedures Experimental Program. — The system will be charged initially with about 15 liters of LiF—BeF2 (66-3L4 mole %), and will be operated with argon and hydrogen to determine the column flooding rate at several salt flow rates. The salt will then be treated with a H2~HF mixture in the feed tank to remove oxides. The flooding tests will be repeated in order to determine the effect of oxide in the salt on flooding. Subsequently, iron fluoride will be added to the salt, and the salt will be counter- currently contacted with hydrogen at several gas and salt flow rates to obtain mass transfer data for reduction of the iron. After these data have been collected, a portion of the initially charged salt will be withdrawn from the system and sufficient LiF and LiF—ThFh eutectic {T73-27 mole %) will be added to yield salt having the composition of 72-16-12 mole % LiF—BeFQ—ThFh. Flooding and iron fluoride reduction tests will be repeated with the new salt mixture. Removal of oxide from the salt by countercurrent contact with a H2—HF mixture in the column will also be investigated. Experimental Method. — The flcoding data will be obtained by main- taining a constant salt flow rate through the column while the gas flow rate is increased in several steps. The pressure drop across the column at each gas flow rate will be recorded; increases in the gas flow rate will be continued until a sharp increase in column pressure drop is observed. The column temperature will be maintained at TOO0°C during the experiment. If subsequent iron fluoride reduction tests show that a higher temperature is necessary, flooding tests will also be made at the higher temperature. Reduction runs will be made in a similar manner after iron fluoride has been added to the salt. Filtered salt samples will be taken from the salt in the feed tank before and after a run. Several flowing stream samples will be withdrawn during each run. Data obtained by analyzing the samples for iron will be used for calculating values of the mass transfer coefficients for the system. The fluoride monitor will provide a check on the extent of iron reduction achieved. In the oxide removal tests, the water analyzer in the column of f-gas stream will be used for determining the amount of oxide removed from the salt. 72 Operating Procedures. — A typical run for testing iron fluoride reduction will consist of the following steps: (1) (2) (3) (%) (5) (6) (7) (8) (9) (11) (12) (13) (14) (15) Heat the system to operating temperature. Add a weighed amount of FeF, (if necessary) to the salt in the 2 feed tank, and sparge for several hours. Sample the salt in the feed tank. Activate the fluoride monitor; start the gas sample pump, heat the sample flow capillary, set the sample and diluent gas flow rates, and set the scrubber solution flow rate. Check the hydrogen supply system; heat the Serfass membrane if it is to be used. Check the O2 and H20 contents of the argon and hydrogen supply. Check the instrument purge rates. Start the hydrogen flow at the specified rate. Pressurize the feed tank and adjust the pressurizing argon flow rate to provide the specified salt flow rate. Record the data on column temperature, flow rates, column pressure drop, salt head above the filter, pressure at the top of the column, and fluoride concentration in the exit gas stream. Withdraw flowing stream salt samples periodically. When the salt supply in the feed tank is exhausted, vent the feed tank and terminate the flow of hydrogen. Purge the system of hydrogen. Transfer the salt in the receiver vessel back to the feed tank. Sparge and sample the feed tank. T3 10. SEMICONTINUOUS REDUCTIVE EXTRACTION EXPERIMENTS IN A MILD-STEEL FACILITY B, A. Hannaford C. W. Kee L. E. McNeese A new column, packed with 1/4-in. molybdenum Raschig rings, was installed in the system, and minor changes were made in some of the piping. Three successful hydrodynamic experiments were performed in which bismuth and molten salt were contacted countercurrently. The results are in excellent agreement with a flooding correlation developed from work with the mercury-water system. Data from a hydrodynamic experiment in which salt flow only was used established that the pressure drop across the new column was approximately equal to that predicted from a literature correlation. 10.1 Eguipment Modifications Operating experiences with the original column and examination of the column following its removal suggested the need for minor changes in the column design. Of prineipal importance was the substitution of 1/b-in. Raschig rings for solid 1/4-in. right cireular cylinders. Installation of a column packed with 1/L-in. Raschig rings was advantageous at this time in that it provided a column whose characteristics should be altered only slightly by deposition of small amounts of iron in the column. The new column had an inside diameter of 0.82 in. and a packed length o of 24 in., excluding end sections. Iach of the end sections contained a 1.5-in.-long packed section consisting of a transition from the 0.82-in. column diameter to the 1.6-in. end section diameter. The void fraction of the column was0.84, as determined by direct meagurement., An X~ray radiograph of the new column (Fig. 28) confirmed that the molybdenum Raschig rings were uniformly distributed. The 1/L-in. and 3/8-in.~long mild~steel rings, which were tack-welded to the slotted support plate in order to prevent the bottom layer of molybdenum rings from sealing the slots, are not clearly shown. Measurements of the pressure drop across the column were made with argon at room temperature in order to establish a reference condition for future comparison. In order to minimize entraimment of bismuth into the salt receiver, two changes were made at the time the column was replaced: (1) the height SISMUTH INLET v Fig. 28. stallation. ORNL DWG 70-454BR1 27 in X-Ray Radiograph of Packed Extraction Column Before In- The column, which has an inside diameter of 0.82 in., is packed with 1/b4-in. Raschig rings. Measured void fraction, 0.8L. 1L 5 of the disengaging section of the column was increased to 3.5 in., and (2) the entraimnment detector was altered slightly to improve the sep- aration of bismuth from the entering salt. 10.2 Treatment of Bismuth and Salt; Adjustment of Zirconium Distribution Ratio Prior to the first hydrodynamic experiment (HR-9) in the new column, the combined salt and bismuth phases were sparged with about 300 g-mecles of 30% HF in hydrogen during a 20-hr operation for the removal of possible oxide contaminants. This was followed by a 6-hr hydrogen sparge (for re- moval of HF) and the addition of metallic thorium to reduce FeF2 and Zth into the bismuth phase. Addition of zirconium to the bismuth is reported to inhibit the mass transport of iron — a source of iron deposits observed in earlier Operations.37 Samples of bismuth and salt taken 24 hr after addition of the initial 154 g of thorium showed that most of the iron but almost none of the zirconium had been reduced. An additional 157 g of thorium was added, and samples (filtered and unfiltered) were taken of each phase 90 hr later. Analysis of the bismuth revealed the presence of 240 ppm of Th, 30 ppm of Li, and 91 ppm of Zr. The concentration of lithium in the bismuth was in good agreement with the calculated equi- librium value based on the thorium concentration, and the zirconium concentration in the bismuth accounted for more than 70% of the zirconium inventory in the system. The salt and bismuth were judged to be in sat- isfactory condition to permit hydrodynamic experiments to be carried out in the new column; the bismuth contained about 90 ppm of zirconium, and the salt had a low iron content (60 ppm). 10.3 Hydrodynamic Experiments HR-9, -10, -11, and -12 Three successful hydrodynamic experiments (HR-9, -10, and -11) were made with the new column, yielding flooding data in good agreement with values predicted from the flooding correlation that was developed from studies with a mercury-water system. Salt and bismuth were transferred from the treatment vessel to their respective feed tanks just prior to each run, and were then countercurrently contacted in the column beginning at a flow rate of about 60 ml of each phase per minute. The flow rate of one of the phases or of both phases, was then increased T6 incrementally until the column flooded. Following each run, the salt and the bismuth were transferred to the treatment vessel. They were returned from this vessel to the feed tank when a subsequent run was made. Flooding rates were defined as those that resulted in a continually increasing pressure drop across the column or a pulsating flow of salt and bismuth through the column. The flow rate data obtained during the three hydrodynamic experiments are summarized in Table 4. The recorded time intervals of less than 8 min resulted from a variety of reasons. At flooded or near-flooded conditions, it was usually impossible to maintain a constant flow rate for each phase. Also, near the end of an experiment and at high f{ ow rates, the supply of bismuth or salt limited the time available. The data from Table LI are plotted as the square root of the superficial velocity of each phase in Fig. 29. The predicted flooding curve for the bismuth--molten salt system, developed on the basis of an assumed constant slip velocity between the phases,38 is shown for comparison. The experimental points for nonflooded operation lie below the predicted flooding curve, and the points for flooded conditions lie above the curve. Thus, the data provide an excellent verification of the flooding correlation developed earlier. The data obtained with 1/4-in. solid cylindrical packing, reported previously, also agree with the predicted flooding curve. However, the range of flow rates investigated with the solid packing did not allow such conclusive confirmation of the predicted curve as is shown in Fig. 29. Reference pressure drop measurements for salt flow only were made as the principal objective of run HR-12. Measurement of the small pressure drop expected (<10 in. H20) required that the bismuth-salt interface at the bottom of the column be depressed below the salt inlet. This was accomplished by routing the argon off-gas flow through a mercury seal in order to pressurize the top of the column, the salt overflow sampler, and the salt receiver. Observed pressure drops through the column were 2, 2.5, and 5 in. H.O at salt flow rates of 68, 127, and 2Lk 2 ml/min, respectively, as compared with values of 0.5, 1, and 2 in. H20 ~ predicted by the Ergun equation. This was regarded as a satisfactory check of the predicted values since the measured values were obtained as the difference between two large numbers. Table 4. Summary of Hydrodynamic Data for Runs HR-9, -10, and -11 Time Volumetric Flow Interval Rate {(ml/min) Run HR- (min) Bisrmuth Salt Comments 9 18 80 85 9 15 115 121 9 175 7T 9 221 133 Apparent bismuth holdup, ~15 vol % 9 I 221 107 Apparent bismuth holdup, V16 to 26 vol % and increasing g 15 150 150 Apparent bismuth holdup, 40 vol % 9 3 130 300 Flooding 10 12 L5 68 10 6 175 68 10 7.5 27h 72 10 3 Lho 20. Incipient flooding; apparent bismuth holdup, 30 vol % and increasing 1i 8 210 51 11 ' 330 50 11 6 406 u7 Apparent bismuth hoidup, 16 to 26 vol % and increasing; incipient flooding 11 3.5 228 100 Apparent bismuth holdup, V15 vol %; not flooded L) 78 ORNL DWG 70-i4715 SALT FLOW, ML/MIN 50 100 200 300 400 500 600 700 20 T : : | . — 700 PREDICTED FLOODING CURVE - 600 V!’ + VpV2:=19.7, (FT/HR)V/2 -1 500 A A FLOODED A O NONFLOODED 85t - 400 T E O -4 300 2 L O 5 g J o o < * > -1 200 % = 10+ O Q 2 ; . O o T A 2 - O 5 3 ~ 100 @ v @ B O 5__ O h 50 O 1 1 1 0 5 10 15 20 (SALT FLOW, V)2 (FT/Hm)/2 Fig. 29. Flooding Data for the Bismuth~Salt System Compared with the Flooding Curve Predicted from Data from a Mercury-Water System. 19 10.4 Maintenance of Equipment The amount of maintenance work required during this period was small; there were no instances of iron deposition and hence no formation of plugs. Twe transfer lines failed, releasing a very small amount of salt. A salt transfer line from the treatment vessel developed a leak, apparently due to air oxidation of the steel tubing on the outside of a bend. A hole developed in a weld between the salt sampler and the specific gravity pot; this failure was probably due to contamination of the weld with salt or bismuth at the time the specific gravity pot was installed. 1. 10. 11. l2l 80 11. REFERENCES L. E. McNeese, "Rare Earth Removal Using the Metal Transfer Process," Engineering Development Studies for Molten-Salt Breeder Reactor Processing No. 5, ORNL-TM-3140 (in press M. J. Bell and L. E. McNeese, Engineering Development Studies for Molten-Salt Breeder Reactor Processing No. 1, ORNL-TM-3053, pp. 38-48, L. M. Ferris, MSR Program Semiann. Progr. Rept. Feb. 28, 1970, ORNL~L548, po. 289-92, Jeje and C. R. Bozzuto, Axigl Mixing in Open Bubble Columns , MIT-CEPS-X-102 (1970). M. S. Bautista and L. E. McNeese, Engineering Development Studies for Molten-Salt Breeder Reactor Processing No. 4, ORNL-TM-3139, pp. 38-83. A. M. Sheikh and J. D. Dearth, Axial Mixing in Open Bubble Columns, MIT-CEPS-X-91 (1969). J. S. Watson and L. E. McNeese, "Axial Mixing in Open Bubble Columns," ¥ngineering Development Studies for Molten-Salt Breeder Reactor Process- ing No. 5, ORNL-TM~-3140 (in publication). R. M. Davies and G. I. Taylor, Proc. Roy. Soc. (London), Ser. A 200, 375 (1950). e L. E. McNeese and M. E. Whatley, Engineering Development Studies for Molten-Salt Breeder Reactor Processing No. 2, ORNL-TM-3137, pp. 22-43. MSR Program Semiann. Progr. Rept. Feb. 29, 1968, ORNL-4254, pp. 260-63. H. F. Bauman, personal communication, August 1970. L. E. McNeese, "Protactinium Isolation Using Fluorination~~Reductive ! ? Extraction," Engineering Development Studies for Molten-Salt Breeder Reactor Processing No. 5, ORNL-TM-3140 (in press). E. L. Youngblood, R. P. Milfaord, R. G. Nicol, and J. B. 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Weber, "An Improved Continuous Internal- Electrolysis Analyzer for Gaseous Fluorides in Industrial Environ- ments," Am. Ind. Hyg. Assoc. J. 23, 48-57 (1962}, B. A. Hannaford, H. D. Cochran, L. E. McNeese, and C. W. Kee, Engineering Development Studies for Molten-Salt Breeder Reactor Processing No., 3, ORNL-TM~-3138, pp. 30-39. J. 5. Watson and L. E. McNeese, "Hydrodynamics of Packed Column Op- eration with High Density Fluids," FEngineering Development Studies for Molten-Salt Breeder Resctor Processing No. 5, ORNL-TM~31L0 (in press). » \O o= O\ oo 10. 12. 13. 1k, 15. 16. 17. 18. 19. 20. 21. 22, 23. ol 25. 26. 27, 28-38. 39. 40. ha, 85. INTERNAL DISTRIBUTION F. Baes . Bauman . Beall Bell Bennett . Blanco Blankenship . Boyad . Briggs . Brooksbank . Brown Carter Cochran, Jr. Culler Distefano Eatherly Ferguson Ferris Frye Grimes Grindell Haag Hannaford Hightower, Jr. Kee Lindauver McCoy McNeese Moulton Nichols . 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