ORNL-TM-3141.txt

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- 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
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SALT CONTAIKING
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SALT CONTAINING RARE EARTHS

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EXTRACTOR

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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
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beg
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EXTRACTOR 4:
1 STAGE |
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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
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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. Ruch, Cor~
rosion of the Volatility Pilot Plant INOR-8 Hydrofluorinator and
Nickel 201 Fluorinator During Forty Fuel Processing Runs with Zir-

~ conium-Uranium Alloy, ORNL-3623 (March 1965).

 

J. C. Mailen, "Volatilization of Uranium as the Hexafluoride from
Drops of Molten Fluoride Salt," paper presented at the American
Chemical Society National Meeting, Chicago, Sept. 2, 196k,
81

 

 

 

 

 

/

S
'f ] \‘\(
/ 15., G. I. Cathers, M. R. Bennett, and R. L. Jolley, The Fused Salt-
\ Fluoride Volatility Process for Recovering Uranium, ORNL-2661 (1959).
16. jw. E. Whatley et al., Unit Operations Section Monthly Progress
Report, September 1963, ORNL-TM-785 (196L).
17. (hem. Technol. Div. Ann. Progr. Rept. May 31, 1965, ORNL-3830, PO -
:\\i'{'\ ] P ".‘7 l— T 5 .
7 \,/

K‘18,/ R. P. Milford, S. Mann, J. B. Ruch, and W. H. Carr, Jr., "Recovering
>~ 7 Uranium Submarine Reactor Fuels," Ind. Eng. Chem. 53, 357 (1961).

| 19. MSR Program Semiann. Progr. Rept. Aug. 31, 1968, ORNL-L34h, pp. L-11.

20. Chem. Technol. Div. Ann. Progr. Rept. May 31, 1967, ORNL-41L45, pp.
95-97.

 

21. MSR Program Semiann. Progr. Rept. Aug. 31, 1968, ORNL-434h4, pp. 302-5.

22. N. R. Stansel, Induction Heating, McGraw-Hill, New York, 19490,

 

23. P. G. Simpson, Induction Heating, Coil, and System Design, McGraw-
Hill, New York, 1960,

 

24, N. R. Stansel, Induction Heating, McGraw-Hill, New York, 1940, p. 31.

 

25. P. G. Simpson, Induction Heating, Coil, and System Design, McGraw-
Hill, p. 1b1.

26. J. R. Hightower, Jr., and L. E. McNeese, "MSRE Distillation Experi-
ment," Engineering Development Studies for Molten-Salt Breeder
Reactor Processing No. 5, ORNL-TM-3140 (in press).

27« J. R. Hightower, Jr., and L. E. McNeese, Low-Pressure Distillation
of Molten Fluoride Mixtures: Nonradioactive Tests for the MSRE Dis-~
tillation Experiment, ORNL-434L (January 1971).

28. S. 8. Kirslis and F. F. Blankenship, MSR Program Semiann. Progr.
Rept. Feb. 29, 1968, ORNL-L254k, p. 100.

 

 

29, 5. S. Kirslis and F. F. Blankenship, MSR Program Semiann. Progr.
Rept. Feb. 28, 1969, ORNL-4396, pp. 145-53.

 

30. M. S. Lin and L. E. McNeese, Engineering Development Studies for _
Molten-Salt Breeder Reactor Processing No. 2, ORNL-TM-3137, pp. 82-89.

31. J. H. Shaffer, MSR Program Semiann. Progr.,Rept. July 31, 1964, ORNI-
3708, pp. 288~303.
32.

33.

3h.

35.

36.

37.

38.

G. E. Brown and N. A. Bhagat, A Prelininary Conceptual Design of a
Pilot Plant for the Production of Purified PFused Fluoride Mixtures,
ORNL-MIT-25 (May 30, 1967).

D. A. Jones and J. A, Alvarez, Reduction of Iron Fluoride with Hy-
drogen in Mixtures of Molten Salts in a Packed Column (Part I),
ORNT.-MIT-53 (May 6, 1968).

J. A, Alvarez and W. H. Pitcher, Jr., Reduction of Iron Fluoride
with Hydrogen in Mixtures of Molten Salts in a Packed Column (Part
IT), ORNL-MIT-56 (May 29, 1968).

W. S. Pappas, "Continuous Moisture Analyzer for Gases Containing
Hydrogen Fluorides,”" Anal. Chem. 38, 615 (1966).

0. H. Howard and C. W. 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).
»

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12.
13.
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16.
17.
18.
19.
20.
21.
22,
23.
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25.
26.
27,
28-38.

39.

40.

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85.

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