FFR_chap06.txt

CHAPTER 6
CHEMICAL PROCESSING*

6-1. INTRODUCTION

One of the principal advantages of fluid fuel reactors is the possibility of
continually processing the fuel and blanket material for the removal of
fission products and other poisons and the recovery of fissionable material
produced. Such continuous processing accomplishes several desirable
objectives: (a) improvement of the neutron economy sufficiently that the
reactor breeds more fissionable material than it consumes, (b) minimiza-
tion of the hazards associated with the operation of the reactor by main-
taining a low concentration of radioactive material in the fuel, and (c) im-
provement of the life of equipment and stability of the fuel solution by
removing deleterious fission and corrosion products. The performance
and operability of a homogeneous reactor are considerably more dependent
on the processing cycle than are those of a solid fuel reactor, although the
objectives of processing are similar.

The neutron poisoning in a homogeneous reactor from which fission
product gases are removed continuously is largely due to rare earths [1],
as shown in Fig. 6-1. In Fig. 6-1 the rare earths contributing to reactor
poisoning are divided into two groups. The time-dependent rare earths
are those of high yield and intermediate cross section, such as Nd43
and Nd'#% Prl4l and Pm!'47) which over a period of several months could
accumulate in the reactor and result in a poisoning of about 209,. The
constant rare-earth poison fraction is due primarily to Sm'#® and Sm!5!
which have very large cross sections for neutron absorption but low yield,
and therefore reach their equilibrium level in only a few days’ operation.
Poisoning due to corrosion of the stainless-steel reactor system was cal-
culated for a typical reactor containing 15,500 {t* of steel corroding at a
rate of 1 mpy. It is assumed that only the nickel and manganese contribute
to the poisoning, since iron and chromium will hydrolyze and precipitate
and be removed from the reactor system; otherwise, corrosion product
poisoning would be four times greater than indicated in Fig. 6-1. The
control of rare earths and corrosion product elements is discussed in sub-
sequent sections of this chapter. Removal of solids from the fuel solution
also improves the performance of the reactor by diminishing the deposition
of scale on heat-transfer surfaces and reducing the possibility of erosion of
pump impellers, bearing surfaces, and valve seats.

*By R. A. McNees, with contributions from W. E. Browning, W. D. Burch,
R. E. Leuze, W. T. McDuffee, and 8. Peterson, Oak Ridge National Laboratory.

301
302 CHEMICAL PROCESSING [cHAP. 6

 

  
 
 
 
  
 
 
    
  
    
   

4.5 | !
A. Total Rare Ecrths
40 | B. Time Dependent Rare Earths
' C. Corrosion Products at 1 mpy
D. Constant (98.2% Rare Earths
3.5 + 1.8% Cd)
E. Sr8% 4 1c99
30} F. Alkali Metals ]
’ G. VIIB(13Y)
R H. Nobie Metals
e 2.5 (RR103)
2
z

 

 

 

 

 

0 30 60 90 120 150
Irradiation Time, days

Fig. 6-1. Poison effect as a function of chemical group in core of two-region
thermal breeder.

D,O
2 To Reactor

u, Th, Pa, Blanket

Fission Products

Slurry from DQ_O

Blanket | Recavery

 
  
 

Decay
! Storage
160 days)

Overflow Returned
to Reactor Core

    
 

U and
Fission
Products

F —
eed from—b\ Z——Hyd roclone
s /

       
   

Fission

Products
Reactor Core Solvent

v ‘ '{

 

Separation | Extraction
- - " Pa, T B
| Underflow Fission Products
Containing Decay | U
Insoluble | Storage
| Fission u (120 days) To Reactor
| Products To Reactor Core

Core

Fic. 6-2. Conceptual flow diagram for processing fuel and blanket material from
a two-region reactor,

The biological hazards associated with a homogencous reactor are due
chiefly to the radioactive rare earths, alkaline earths, and iodine [2].
The importance, as a biological hazard, of any one of these groups or nu-
clides within the group depends on assumptions made in describing ex-
posure conditions; however, I'3! contributes a major fraction of the radia-
tion hazards for any set of conditions. While the accumulation of hazardous
materials such as rare earths and alkaline earths will be controlled by the
processing methods to be described, less is known about the chemistry of
6-1] INTRODUCTION 303

Blanket
1.4 mUOQSO4
in D70 at
250°C

    
  

[—PDQO

1 D50 = Dissolution
Recovery [ | inHNO,

 

 

 

 

 

 

 

 

 

- Solvent

 
  
 

       
 

 

Extraction
Partition
Stripping

 

Solvent

=== Lrgnium Fission Plutonium Uranium
— Plutonium Product Waste

FiG. 6-3. Conceptual flow diagram for processing blanket material from a two-
region plutonium producer.

iodine in the fuel systems and methods for removing it. Ixisting informa-
tion on iodine processing is discussed in Section 6-3.

Schematic flowshects for proposed processing schemes for two types of
two-region aqueous homogencous reactors are shown in Figs. 6-2 and 6-3.
In both cases, solids are removed by hydroclones and concentrated into
a small volume of solution for further processing. The nature of such
processing will be determined by the exact design and purpose of the
reactor. Thus, for a two-region plutonium producer, the core and blanket,
materials would have to be processed separately to avoid isotopic dilution,
while for a thorium breeder, core and blanket material could be processed
together. However, if an attractive method should be developed for leach-
ing uranium and/or protactinium from a thorium-oxide slurry without
seriously altering the physical properties of the slurry, the two materials
could be processed separately. In a similar way, the relation between
iodine control and fission product gas disposal is such that neither problem
can be disassociated from the other. A specific, complete, and feasible
chemical processing scheme cannot be proposed for any reactor without
an intimate knowledge of all aspects of design and operation of the reactor.
However, some of the basic chemical knowledge needed to evaluate various
304 CHEMICAL PROCESSING [crAP. 6

possible processing methods has been developed and is presented in the
following sections.

6-2. CorE PRrROCESSING: SoLIDs REMOVAL

6-2.1 Introduction. Early in the study of the behavior of fisston and
corrosion products in uranyl sulfate solutions at temperatures in the range
250 to 325°C, it was found that many of these elements had only a limited
solubility under reactor conditions. Detailed studies of these elements
were conducted and devices for separating solids from liquid at high
temperature and pressure were constructed and evaluated. Based on this
work, a pilot plant to test a processing concept based on solids separation at
reactor temperature was installed as an adjunct to the HRE-2. These
processing developments are discussed in this section.

6-2.2 Chemistry of insoluble fission and corrosion products. Of the
nongaseous fission produets, the rare earths contribute the largest amount
of neutron poison to a homogencous reactor after a short period of operation
(Fig. 6-1). Therefore, a detailed study of the behavior of these elements

TaBLE 6-1

SOLUBILITY OF LANTHANUM SULFATE IN
0.02 m U0:804,—0.005m HaSO4 as A FuNcTION OF
SoLuTioON TEMPERATURE

 

 

 

 

mg La2(S04)s/kg H20
Temperature,
°C True Concentration required to
solubility initiate precipitation
190 250 760
210 130 360
230 54 167
250 25 77
270 12 36

 

 

 

 

 

has been made. All the rare earths and yttrium showed a negative tem-
perature coefficient of solubility in all the solutions studied and a strong
tendency to supersaturate the solutions, as shown in Table 6-1. With the
exception of prascodymium and neodymium, which are reversed, the solu-
hility at a given temperature and uranyl sulfate concentration increased
with increasing atomic number, with yttrium falling between neodymium
6-2] CORE PROCESSING: SOLIDS REMOVAL 305
and samarium, as shown in Table 6-2. Increasing the uranyl sulfate
concentration increased the solubility of a given rare-earth sulfate, as

shown in Table 6-3.
TABLE (-2

SOLUBILITY OF VARIOUS RARE-IARTH SULFATES IN
0.02m U0s804—0.005m HoSO4 aT 280°C

 

 

 

 

 

Solubility, Solubility,
Salt mg/kg H20 Salt mg/kg H20
Las(804)3 10 Nd2(S04)5 110
Ce2(804)3 50 Y2(S04)3 240
PI‘2(804)3 170 SHlQ(SO4)3 420
TABLE 6-3

Errect oF URANYL SULFATE CONCENTRATION ON THE SOLUBILITY OF
NEODYMIUM SULFATE AT VARIOUS TEMPERATURES

 

 

 

 

 

Nd2(S04)3 solubility, mg/kg H20
U, g/kg H20
250°C 280°C 300°C
5.7 270 115 73
10.8 400 200 120
16.6 770 300 180
22 4 > 1000 o00 500

 

 

 

 

 

In a mixture of rare-earth sulfates the solubility of an individual rare
earth 1s less than it would be if it were present alone. For example, the
solubility of praseodymium sulfate at 280°C is 170 mg/keg H»0 with no
other rare earths present, as compared with 12 mg,/ke H>0 in a solution
made up with a rarc-earth mixture containing 6 praseodymiuin sulfate.
Samples of the precipitating salts i1solated from solution at 280°C' have
usually been the sulfates and contained no uranium. However, under
special conditions a mixed sulfate salt of neodymium and uranium has been
observed [3].

The alkaline earths, barium and strontiun, also show a negative tem-
perature coeflicient, but not so strongly as do the rare earths; almost no
effect can be seen when the temperature of precipitating solutions is in-
306 CHEMICAL PROCESSING [cHap. 6

creased from 250 to 300°C. At 295°C in 0.02 m U02504—0.005 m H2S04
solution, the solubility of barium sulfate is 7 mg/kg HoO and that of
strontium sulfate is 21 mg/kg Ho0O. Both the alkaline and rare-earth
sulfates show a strong tendency to precipitate on and adhere to steel
surfaces hotter than the precipitating solutions, and this property can be
used to isolate these solids from liquids at high temperatures.

Other fission and corroston product elements hydrolyze extensively at
250 to 300°C and precipitate as oxides, leaving very low concentrations
in solution. Iron(III) at 285°C has a solubility of 0.5 to 2 mg Fe/kg Ho0
and chromium(III), 2 to 5 mg/kg HoO. At 285°C less than 5 mg of zir-
conium or niobium per kilogram of H20 remains in solution.

For other elements of variable valence, such as technetium, the amount
of the clement in solution is determined by the stable valence state under
reactor conditions. In general, the higher valence states better resist hy-
drolysis and remain in solution. Thus at 275°C in 0.02 m UO2S04 Te(VII)
is reduced to Te(IV) if hydrogen is present, and only 12 mg/kg H20 re-
mains in solution. However, a slurry of TcO» in the same solution but with
oxygen present dissolves to give a solution at 275°C with a technetium
concentration of more than 9 g/kg H»0. The same qualitative behavior is
observed with ruthentum. Selenium and tellurium in the hexapositive state
are much more soluble than when in the tetrapositive state [4].

A few eclements, e.g., cesium, rubidium, nickel, and manganese, intro-
duced into the fuel solution by fission or by corrosion of the system, are
very soluble under reactor conditions. Their removal and control are dis-
cussed in Section 6-1.

6-2.3 Experimental study of hydroclone performance. It is evident
from the preceding section that the amount of uranium withdrawn from
the reactor diminishes if the collection, eoncentration, and 1solation of the
insolubles can be effected at high temperature. One device capable of
collecting and concentrating solids at high temperature is a solid-liquid
cyclone separator called a “hydroclone,” or “clone.” A diagram of a hydro-
clone is shown in Fig. 6-4. In operation, a solids-bearing stream of liquid
is injected tangentially into the wide portion of a conical vessel. Solids
concentrate in a downward-moving layer of liquid and are discharged from
the bottom of the clone into the underflow receiver. Partially clarified
liquid leaves from the top of the clone through a vortex finder. Use of the
underflow recciver eliminates mechanical control of the discharge flow
rate and, by proper choice of hydroclone dimensions, any desired ratio of
overflow rate to underflow rate can be achieved. The driving foree for the
system is provided by a mechanical pump.

The factors influencing the design of an effective hydroclone for homo-
geneous reactor processing use have been studied, and hydroclone designs
6-2] CORE PROCESSING: SOLIDS REMOVAL 307
t/Overflow

] ——-DO

 

1%

Feed —-i

Hydroclone | //

 

 

 

 

 

 

 

 

 

 

 

 

 

Underfiow
Port

V5

3 £

e— Underflow
Receiver

 

 

 

 

 

 

Fig. 6-4. Schematic diagram of a hydroclone with associated underflow receiver.

based on these studies have been tested in the laboratory and on various
circulating loops [5]. All tests have shown conclusively that such hydro-
clones can separate insoluble sulfates or hydrolyzed materials from liquid
streams at 250 to 300°C. In the HRE-2 mockup loop a mixture of the sul-
fates of iron, zirconium, and various rare earths, dissolved in uranyl-
sulfate solution at room temperature, precipitated when injected into the
loop solution at 250 to 300°C. The solids concentrated into the underflow
receiver of a hydroclone contained 75% of the precipitated rare-earth
sulfates.  When the lanthanum-sulfate solubility in the loop solution was
exceeded by 109, the concentration of rare earths in the underflow receiver
was four to six times greater than in the rest of the loop system; some
accumulation of rare earths was observed in the loop heater. A large
fraction of the hydrolyzed iron and zirconium was collected in the gas
separator portion of the loop. In the separator the centrifugal motion
given to the liquid forced solids to the periphery of the pipe and allowed
them to accumulate. Only about 109 of the solids formed in the loop was
recovered by the hydroclone, and examination of the loop system dis-
closed large quantities of solids settled in every horizontal run of pipe.
308

 

Reactor
Cell

Pump

 

CHEMICAL PROCLESSING

  
     
  
 

 

Underflow
Pot

Hydroclone

Recombine:
-Condenser

  
    

10 gal

Separator

Lsad
e

DZO Receiver

g Decay Tank {2)

Recombine:
-Condenser

  
  
 

HZO to Waste

Separator

100 gal
\

[cHAP. 6

 

 

Dissolver

 

 

 

 

 

   
    

 

  
 
 

7 gal @7\
Dissolver
Solution
DO Addition 67
»+ Fuel Addition Sampler o

]

Carrier

 

 

 

Transfer Tank

 

 

Fie. 6-5. Schematic flow diagram for the HRE-2 chemical processing plant.

TaBLE 64

Dmaensions oF HRE-2 HypROCLONES

 

 

 

 

 

 

 

 

Dimension, in.
Symbol Location . . _
0.25-in. 0.40-1n, 0.56-1n.
hydroclone hydroclone hydroclone
D1 Maximum inside 0.25 0.40 0.56
diameter
L Inside length 1.50 2.40 3.20
Dy Underflow port
diameter 0.070 0.100 0.148
Do Overflow port
diameter 0.053 0.100 0.140
Dy Feed port effec-
tive diameter 0.051 0.118 0.159

 

 

 
2] CORE PROCESSING: SOLIDS REMOVAL 309

Samples taken from the loop after addition of preformed solids and without
the hydroclone operating showed an exponential decrease in solids concen-
tration with a half-time of 2.5 hr; with the hydroclone operating, the half-
time was 1.2 hr. In the HRIE-2 chemical plant [5], operated with an aux-
iliary loop to provide a slurry of preformed solids in uranyl sulfate solution
as a feed for the plant, the half-times for solids disappearance and removal
were 11 hr without the hydroclone and 1.5 hr with it. The efficiency of the
hydroclone for separating the particular solids used in these experiments
was about 109, With gross amounts of solids in the system, concentration
factors have been as large as 1700.

Correlation of these data with anticipated reactor chemical plant oper-
ating conditions indicates that the HRE-2 chemical plant will hold the
amount of solids in the fuel solution to between 10 and 100 ppm. If neces-
zaryv. performance can be improved by increasing the flow through the
chemical plant and by eliminating, wherever possible, long runs of hori-
zontal pipe with low liquid veloeity and other stagnant areas which serve
to accumulate solids.

6-2.4 HRE-2 chemical processing plant.* An experimental facility to
test the solids-removal processing concept has been constructed in a cell
adjacent to the HRI-2. A schematic flowsheet for this facility is shown
in Fig. 6-5.

A 0.75-gpm bypass stream from the reactor fuel system at 280°C and
1700 psi ig circulated through the high-pressure system, consisting of a
heater to make up heat losses, a screen to protect the hydroclone from
plugging, the hydroclone with underflow receiver, and a canned-rotor
circulating pump to make up pressure losses across the system. The
hydroclone is operated with an underflow receiver which is drained after
each week of operation, at which time the processing plant is isolated
from the reactor system.

At the conclusion of each operating period 10 liters of the slurry in the
underflow pot is removed and sampled. The heavy water is evaporated
and recovered, and the solids are dissolved in sulfurie acid and sampled
again. The solution is then transferred under pressure to one of two 100-gal
decay storage tanks, Tollowing a three-month decay period, the solution
1= transferred to a shielded carrier outside the cell and then to an existing
sulvent extraction plant at Oak Ridge National Laboratory for uranium
decontamination and recovery.  The sulfuric acid solution step is incor-
porated n the HRIS-2 chemical plant to ensure obhtaining a satisfactory
<umple.  This step would presumably not be necessary in a large-scale
plant.

*Contribution from W. D. Burch.
310 CHEMICAL PROCESSING [cHAP. 8

 

Fic. 6-6. HRE-2 chemical plant cell with equipment.
6-2] CORE PROCESSING: SOLIDS REMOVAL 311

All equipment is located in a 12- by 24- by 21-ft underground cell located
adjacent to the reactor cell and separated from it by 4 ft of high-density
concrete. Other construction features are similar to those of the reactor
cell, with provisions for flooding the cell during maintenance periods in
order to use water as shielding. Figure 6-6, a photograph of the cell prior
to installation of the roof plugs, shows the maze of piping necessitated by
the experimental nature of this plant.

Dimensions of the three sizes of hydroclones designed for testing in this
plant are shown in Table 6—4. These three hydroclones, which have been

 

Blind Flange ;‘
Closure ‘{? !

1

i Y:‘X‘E:\'Y- AN !

b m&\ A0 Sy \\\\T{\*\
G

e < .:\‘.\\ .

i
i
o

N

 

  
 
 
  

Hexagonal Head —
ol oo
it :

Cap Plug
L R
|
|

      
 

   
   
 

 

o f For
*~"  ]lLeak Detector

Spider

Overflow

Pressure Screw

Hydroclone
Retaining Plug

 

Gold Wire Gasket

F1c. 6-7. HRE-2 chemical plant hydroclone container.

selected to handle the range of possible particle sizes, are interchangeable
at any time during radioactive operation through a unique, specially ma-
chined flange, shown in Fig. 6-7. Removal of the blind closure flange ex-
poses a cap plug and retainer plug. Removal of these with long-handled
socket wrenches permits access to the hydroclone itself. This operation has
heen performed routinely during testing with nonradioactive solutions.

In processing homogeneous reactor fuel, a transition from a heavy- to a
natural-water system is desirable if final processing is to be performed in
conventional solvent extraction equipment. Such a transition must be ac-
complished with a minimum loss of D20 and a minimum contamination of
312 CHEMICAL PROCESSING [cHAP. 6

the fuel solution by H20 in recycled fuel. Initial tests of this step in the
fuel processing cycle have been carried out [6]. In these experiments a
mixture of 59, D0, 959%, H:20 was used to simulate reactor fuel liquid.
The dissolver system was cycled three times between this liquid and or-
dinary water, with samples being taken during each portion of each cycle.
Isotopic analysis of these samples showed no dilution of the D20 in the
enriched solution and no loss of D20 to the ordinary water system.

At expected corrosion rates, approximately 400 g of corrosion products
will be formed in the reactor system per week, and the underflow receiver
was therefore designed to handle this quantity of solids. The adequacy of
the design was shown when more than three times this quantity of solids
was charged to the underflow receiver and drained in the normal way with-
out difficulty.

Full-scale dissolution procedures have also been tested [6]. To minimize
the possibilities of contaminating the reactor fuel solution by foreign ions,
a dissolution procedure was developed using only sulfuric acid. This con-
sists of a 4-hr reflux with 10.8 M HoS04 in a tantalum-lined dissolver fol-
lowed by a 4-hr reflux with 4 3 H2S04, and repeated as required until
dissolution is complete. Decay storage tanks and other equipment required
to handle the boiling 4 3 HoSO4 are fabricated of Carpenter—20 stainless
steel. Tests have repeatedly demonstrated more than 99.59, dissolution
of simulated corrosion and fission products in two such eycles.

The HRE-2 hydroclone system has been operated as an integral part
of the reactor system for approximately 600 hr and for an additional
1200 hr with a temporary pump loop during initial solids-removal tests.
During this operating period, in which simulated nonradioactive fuel
solutions were used, the performance of the plant was satisfactory in all
respects.

6—3. Fission Propuct Gas DisposaL*

6-3.1 Introduction. To prevent the pollution of the atmosphere by
radioactive krypton and xenon isotopes released from the fuel solution, a
system of containment must be provided until radioactive decay has re-
duced their activity level. This is accomplished by a method based on the
process of physical adsorption on solid adsorber materials. If the adsorber
system 1s adequately designed, the issuing gas stream will be composed of
long-lived Kr®, oxygen, inert krypton isotopes, inert xenon isotopes, and
insignificant amounts of other radioactive krypton and xenon isotopes. In
case the activity of the Kr®5 is too high for dilution with air and discharge
to the atmosphere, the mixture may be stored after removal of the oxygen

*Contribution from W. E. Browning.
3] FISSION PRODUCT GAS DISPOSAL 313

 

 

 

40
r— 171717 1 ©° © B 1 oo
20—
10 —
- oh
E 5 SieYes —
'g Mo\ecu\c“
o -1
3 Vot Sieqes
E N\O\ecu 5 1 0",\ —
o cutes Sieve
3 o
E 2 12 ]
- gitico &2
0
<
3 5
» a0
g 1.0 3\2’5 S‘“COG —
e Tr e ]
e
.g_ | oW Hi-Sil %-303
> | —
x — /4 Dr\OCe\ —
05 —
gihca Gel-70
0.2 \at Gieves- AR —]
Zeo-Dur MoleculS

Micro Cel-A

(1 1 1 [ [ [ |
0.1l
0 10 20 30 40 50 60 70 80 90 100 110 120 130

Krypton Pressure, mm

 

Fic. 6-8. Adsorption of krypton on various adsorbents at 28°C.

or further separated by conventional methods into an mert xenon fraction
and a fraction containing Kr®5 and inert krypton.

6-3.2 Experimental study of adsorption of fission product gases. Evalu-
atiou of vartous adsorber materials based on experimental measurements of
the equilibrium adsorption of krypton or xenon under static conditions is
in progress [7]. Results in the form of adsorption isotherms of various
solid adsorber materials are presented in Fig, 6-8,

A radioactive-tracer technique was developed to study the adsorption
efficiency (holdup time) of small, dynamic, laboratory-scale adsorber
systems [8]. This consists of sweeping a brief pulse of Kr8 through an ex-
perimental adsorber system with a diluent gas such as oxygen or nitrogen
314 CHEMICAL PROCESSING [cHAP. 6
and monitoring the efluent gases for Kr®5 beta activity. The activity in
the gas stream versus time after injection of the pulse of Kr®5 is recorded.
A plot of the data gives an experimental elution curve, such as shown in
Fig. 6-9, from which various properties of an adsorber material and ad-
sorber system may be evaluated.

Among the factors which influence the adsorption of fission product
gases from a dynamic system are (1) adsorptive capacity of adsorber ma-
terial, (2) temperature of adsorber material, (3) volume flow rate of gas
stream, (4) adsorbed moisture content of adsorber material, (5) composi-
tion and moisture content of gas stream, (6) geometry of adsorber system,
and (7) particle size of adsorber material. The average time required for
the fission product gas to pass through an adsorber system, fmax, 1s influ-
enced by the first five of the above factors. The shape of the experimental
elution curve is affected by the last two.

The temperature of the adsorber material is of prime importance. The
lower the temperature the greater will be the adsorption of the fission gases,
and therefore longer holdup times per unit mass of adsorber material will
result. The dependence of adsorptive capacity, k, on temperature as de-
termined by holdup tests with some solid adsorber materials is shown in
Table 6-5.

TABLE 6-H

ADsSoRPTIVE CaprAcITY OF VARIOUS MATERIALS AS A FUNCTION
oF TEMPERATURE

 

 

 

 

 

 

 

 

 

 

ce gas/g adsorbent*®
Gas Diluent Adsorber
273°K 323°K 373°K
Xe Oy Charcoal 4.7 x 103 4.0 x 102 80.0
Kr He Charcoal 1.8 x 10° 34 9.6
Kr Oz or N» Charcoal, 68 24 11.0
Kr O; Linde Molecular 23 9 4.5
Sieve SA
Kr O Linde Molecular 11 5.7 3.5
Steve 10X

 

*(Gas volume measured at temperatures indicated.

 
6-3] FISSION PRODUCT GAS DISPOSAL ald

 

1 | i

Average Holdup Time (rmax) _l

 
   

Kr85 Activity in Effluent Gas Stream —»

Breakthrough Time (1}

    

 

 

l I J

Time After Injection of Kr85 —

 

F1ac. 6-9. Experimental Kr83 elution curve.

At a given temperature, the average holdup time, fyax, 18 inversely pro-
portional to the volume flow rate of the gas stream. If the volume flow
rate ix doubled, the holdup time will be decreased by a factor of two.

All the solid adsorber materials adsorb moisture to some degree. Any
adsorbed moisture reduces the active surface area available to the fission
gases and thus reduces the average holdup time,

The geometry of the adsorber system influences the relation between
breakthrough time, f,, and average holdup time, ¢max, as shown in I'ig. 6-9.
Ideally, for fission product gas disposal, a particular atom of fisslon gas
should not emerge from the adsorber system prior to the time #,... Since
this condition cannot be realized in practice, the difference between break-
through and average holdup times should be made as small as possible.
For a given mass of adsorber material a system composed of long, small-
diameter pipes will have a small difference between ¢, and ., whereas a
system composed of short, large-diameter pipes will not.

The particle size of the adsorber material 1s important for ensuring n-
timate contact between the active surface of the adsorber material and
the fission gases. A system filled with large particles will allow some mole-
316 CHEMICAL PROCESSING [cHAP. 6

cules of fission gases to penctrate deeper into the system before contact is
made with an active surface, while the pressure drop across a long trap
filled with small particles may be excessive. Material between 8 and 14
mesh in size 1s satisfactory from both viewpoints.

6-3.3 Design of a fission product gas adsorber system. The design of
an adsorber system will be determined partly by the final disposition of
the effluent gas mixture. If ultimate disposal is to be to the atmosphere,
the adsorber system should be designed to discharge only Kr®* plus inert
krypton and xenon isotopes. If the effluent gases are to be contained and
stored, the adsorber system may be designed to allow discharge of other
radioactive krypton and xenon isotopes. In the following discussion it is
assumed that final disposal of the effluent gas mixture will be to the at-
mosphere. The following simple relation has been developed which is use-
ful in finding the mass of adsorber material in such an adsorber system:

fl’[ = ;,Fj’tmax,

14

where M = mass of adsorber material (grams), F = gas volume flow rate
through adsorber system (cc/min), k= adsorptive capacity under dy-
namic conditions (cc/g), and tnax = average holdup time (min).

The operating characteristics of the reactor will dictate the composition
and volume flow rate of the gas stream; fmax will be determined by the al-
lowable concentration of radioactivity in the effluent gas; &k values for
krypton and xenon must be determined experimentally under conditions
simulating these in the full-scale adsorber system. It should be noted
(Fig. 6-9) that a portion of the fission gas will emerge from the adsorber
system at a time #, prior to the average holdup time, fmax. The design
should ensure that radioactive gas emerging at time ¢ has decayed suffi-
ciently that only insignificant amounts of activity other than Kr® will be
discharged from the bed.

The adsorber system should be operated at the lowest convenient
temperature because of the dependence of adsorptive capacity on tempera-
ture. Beta decay of the fission product gases while passing through the
adsorber system will increase the temperature of the adsorber material
and reduce the adsorptive capacity. Temperature control is especially
eritical if the adsorber system uses a combustible adsorber material, such
as activated charcoal, with oxygen as the diluent or sweep gas.

6-3.4 HRE-2 fission product gas adsorber system. The HRE-2 uses a
fission product gas adsorber system containing Columbia G activated
charcoal. Oxygen, contaminated with the fission produet gases, is removed
6-4] CORE PROCESSING: SOLUBLES 317

from the reactor, dried, and passed into this system, and the eflluent gases
are dispersed into the atmosphere through a stack.

The adsorber system contains two activated charcoal-filled units con-
nected in parallel to the gas line from the reactor. Iiach unit consists of
40 ft of 4-in. pipe, 40 ft of 1-in. pipe, 40 ft of 2-in. pipe, and 60 ft of G-in.
pipe connected in series. The entire system is contained in a water-filled
pit, which is buried underground for gamma shielding purposes.  lach
unit is filled with approximately 520 Ib of Columbia G activated charcoal,
8 to 14 mesh.

The heat due to beta decay of the short-lived krypton and xenon 1so-
topes is diminished by an empty holdup volume composed of 160 {t of
3-1n. pipe between the reactor and the charcoal adsorber system. This pre-
vents the temperature of the charcoal in the inlet sections of the adsorber
system from exceeding 100°C. Iixecessive oxidation of the charcoal by the
oxygen in the gas is further prevented by water-cooling the beds.

Before the adsorber system was placed 1n service, its efficiency was
tested under simulated operating conditions [Y]. A pulse of Kr®% (25
millicuries) was injected into each unit of the adsorber system and swept
through with a measured flow of oxygen. In this way the krypton holdup
time was determined to be 30 days at an oxygen flow rate of 250 e¢/min/
unit. Based on laboratory data from small adsorber systems, the holdup
time for xenon is larger than that for krypton by a factor that varies from
30 to 7 over the temperature range of 20 to 100°C. Irom these data, it 1s
estimated that the maximum temperature of the HRE-2 adsorber system
will vary between 20 and 98°C after the reactor has been operating at
10 Mw power level long enough for the gas composition and charcoal
temperature to have reached equilibrium through the entire length of the
adsorber unit. The holdup performance of the adsorber system was cal-
culated with corrections for the increased temperature expected from the
fission gases. The calculated holdup time was found to be 23 days for
krypton and 700 days for xenon; this would permit essentially no Xe!33
to escape from the trap.

6—4. CortE PROCESSING: SOLUBLES

6—4.1 Introduction. While the solids-removal scheme discussed i Sec-
tion 6—1 will limit the amount of solids circulating through the reactor sys-
tem, soluble elements will build up in the fuel solution. Nickel and man-
ganese from the corrosion of stainless steel and fission-produced cesium
will not precipitate from fuel solution under reactor conditions until con-
centrations have been reached which would result in fuel instability and
loss of uranium by coprecipitation. Loss of neutrons to these poisons
would seriously decrease the probability of the reactor producing more
§

318 CHEMICAL TPROCESSING [cHAP. 6

fuel than it consumes. Thercfore, a volume of fuel solution sufficient to
process the core solution of the reactor at a desired rate for removal of
soluble materials is discharged along with the inscluble materials concen-
trated into the hydroelone underflow pot. This rate of removal of soluble
materials depends on the naturc of other chemical processing being done
and on the extentrof corrosion. For example, operation of an iodine re-
moval plant (Section 6-5) reduces the buildup of cesium in the fuel to an
insignificant value by removing .cesium precursors.

6—4.2 Solvent extraction. Processing of the core solution of a homo-
geneous reactor by solvent extraction is the only method presently avail-
able which has been thoroughly proved in practice. However, the amount
of uranium to be processed daily is so small that operation of a solvent ex-
traction plant just for core solution processing would be unduly expensive.
Therefore, the core solution would normally he combined with blanket ma-
terial from a thermal breeder reactor and be processed through a Thorex
plant, but with a plutonium-producing reactor separate processing of core
and blanket materials will be needed. These process schemes are discussed
in detail in Sections 6-6 and 6-7.

The uranium produect from either process would certainly be satisfactory
for return to the reactor. Since solid fuel clement refabrication is not a
problem with homogeneous reactors, decontamination factors of 10 to 100
from various nuclides are adequate and some simplification of present
solvent extraction schemes may be possible.

6-4.3 Uranyl peroxide precipitation. A process for decontaminating the
uranium for quick return to a reactor has been proposed as a means of
reducing core processing costs. A conceptual flowsheet of this process,
which depends on the insolubility of UO4 under controlled conditions for
the desired separation from fission and corrosion products, is shown in
Fig. 6-10. A prerequisite for use of this scheme is that losses due to the
mmsoluble uranium contained in the solids concentrated in the hydroclone
plant be small. Iowever, laboratory data obtained with synthetic solids
simulating those expected from reactor operation indicate that the uranium
content of the solids will be less than 19, by weight. Verification of the
results will be sought during operation of the HRE-2.

In the proposed method, the hydroclone system is periodically isolated
from the reactor and allowed to cool to 100°C. The hydrolyzed solids re-
main as such, but the rare-earth sulfate solids concentrated in the under-
flow pot redissolve upon cooling. The contents of the underflow pot are
discharged to a centrifuge where solids are separated from the uranium-
containing solution and washed with D»0), the suspension being sent to a
waste evaporator for recovery of Dz0.
6-5] CORE PROCESSING: IODINE 319

D0 D207 D20 D250,
SU ‘\ fi* LW t
SFC L SU | e uo, ‘
RE —— 100°C —— SFC—tm 0 C SFc—--l 3o°c ,......uo4—->1 100°C
sy RE _ RE \ G
IsFC ’

Llc—a
-_—

c

O

N

w

O

>

DQO RE

D50 Recovery r——J
SU = Soluble Uranium -

SFC = Soluble Fission ~—H,O
and Corrosion Products

 

RE = Rare Earth Sulfates s
IsU = Insoluble Uranium -5 FC o T Waste Storage
IsFC = Insoluble Fission S FC

and Corrosion Products RE

Fig. 6-10. Schematic flow diagram for decontaminating uranium by uranyl
peroxide precipitation.

Uranium in the clarified solution is precipitated by the addition of either
D202 or NaoOq2. By controlling pID and precipitation conditions, a fast
settling precipitate can be obtained with less than 0.19 of the uranium
remaining in solution. The U0y precipitate is centrifuged or filtered and
washed with D20 and dissolved in 509 excess of 112804 at 80°C before
being returned to the reactor.

In laboratory studies uranium losses have been consistently less than
0.19, for this method and decontamination factors from rare earths greater
than 10. Deecontamination factors from nickel and cesium have been 600
and 40, respectively. It is estimated that the product returned to the re-
actor would contain about 20 ppm of sodium as the only contaminant in-
troduced during processing. Although either the addition of D202 or use
of D202 generated by radiation from the solution itself appears attractive,
acid liberated by the precipitation of UO4 must be neutralized if uranium
losses are to be minimized. Since the entire operation is done in a D20
system, no special precautions to avoid contaminating the reactor with
ordinary water are needed.

6-5. Core Processing: lopine*

6-5.1 Introduction. The removal of iodine from the fuel solution of a
homogeneous reactor is desirable from the standpoint of minimizing the
biological hazard and neutron poisoning due to lodine and reducing the
production of gaseous xenon and its associated problems. lodine will also

*Contribution from S. Peterson,
320 CHEMICATL PROCESSING [crar. 6

poison platinum catalysts [10] used for radiolytic gas recombination in
the reactor low-pressure system and may catalyze the corrosion of metals
by the fuel solution. For this reason a considerable effort has been carried
out at ORNI: and by Vitro [11] to mmvestigate the behavior of iodine in
solution and to develop methods for its removal. In this regard, the iodine
isotopes of primary interest are 8-day I'*! and 6.7-hr 1'35,

6-5.2 The chemistry of iodine in aqueous solutions. Iodine in aqueous
solution at 25°C can exist in several oxidation states. The stable species
are iodide ion, I7; elemental iodine, 12; iodate, 103 ; and periodate, 104~
or H5IO¢s. The last of these exists only under very strongly oxidizing con-
ditions, and 1s immediately reduced under the conditions expected for a
homogeneous reactor fuel. lodide ion can be formed from reduction of
other states by metals, such as stainless steel, but in the presence of the
oxygen necessary in a reactor system it is readily converted to elemental
iodine; this conversion is especially rapid above 200°C. Thus the only
states of concern in reactor fuel solutions are elemental iodine and iodate.
Under the conditions found in a high-pressure fuel system the iodine is
largely, if not all, in the elemental form.

Volatility of todine. Since the volatile elemental state of iodine is pre-
dominant under reactor conditions, the volatility of iodine from fuel solu-
tion is the basis for proposed iodine-removal processes. The vapor-liquid
distribution coeflicient [11] (ratio of mole fraction of iodine in vapor to
that in liquid) for simulated fuel solution and for water at the temperatures
expected for both the high-pressure and low-pressure systems of homo-
geneous reactors 1s given in Table 66,

TABLE 6-6

Varor-Liquip DisTriBUuTION OF lODINE

 

Distribution coefficient,

 

 

 

vapor/liquid
Solution
High pressure Low pressure
(260-330°C) (100°C)
Clean fuel solution 7.4 0.34

(0.02 m U02804—0.005 m HaS04—
0.005m CLISO4, 1-100 pp Is
Fuel solution with mixed figsion and
corrosion products 2.4
Water (pH 4 to 8, 1-13 ppm I2) 0.29 0.009

 

 

 

 

 
6-3] CORE PROCESSING: IODINE 321

TTTTTTTTI Tt T T T T

 

Adsorber Bed
Silvered Alundum

|
—— e — ——
S I S G e

 

 

 

 

\ P SRl SN SN SEnR SN

Ejector

 

 

Gas
Separator

 

—— (g

Heater
e | iquid]

B

- Sampler

L
lodine Spike —'6

Fic. 6-11. Vitro iodine test loop.

 

Canned Rotor
Pump

 

A number of conclusions are evident from these data. Iodine 1s much
more volatile from fuel solution than from water at either temperature,
Fission and corrosion products appear to increase the volatility of 1odine
from fuel solution at 100°C. Increasing the temperature from 100 to 200°C
increases the volatility of 1odine relative to that of water. No systematic
variation of 1odine volatility has been found with iodine concentration in
the range 1 to 100 ppm or temperature in the range 260 to 330°C.

The volatility of 1odine from simulated fuel solution has been verified
by experiments in a high-pressure loop, shown schematically in Fig. 6-11
[11]. The circulating solution was contacted with oxygen in the ejector;
the separated gas was stripped of iodine by passing through a bed of sil-
vered alundum which was superheated to prevent steam condensation.
Potassium iodide solution (containing a radioactive tracer, I'3!) was
rapidly injected into the loop to give an 1odine concentration of 10 ppm.
The lodine concentration decreased exponentially with time in the circu-
lating solution. Table 6-7 gives the half-times for 1odine removal and the
volatility distribution coeflicient, calculated from the removal rate and the
flow rates, based on three experiments with clean fuel solution and two
with added iron. Within the accuracy of flow rate measurement, the coef-
 

 

 

 

 

322 CHEMICAL PROCESSING [cHAP. 6
TABLE 6-7
lopiNE REmovaL FroM A HicH-PrESSURE Loor

Todine Todine
Tempera~ L

\ i removal | distribution

Solution ture, . ..

% half-time, | coeflicient,

min vapor/hquid
('lean fuel solution 230 13.0 7.6
(0.02 m U02304—0.005 m HeSO 34— 6.5 16.8
0.005 m CuS0y) 13.0 8.4
Fuel + 30 ppm Fe3* 220 11.0 10.9
Fuel + 300 ppm Fe?* 225 11.0 9.5

 

 

 

 

 

 

ficient agrees with the average value of 7.4 obtained in numerous static
tests over the high-temperature range. Iron appears to have no effect.

Oxidation state of 1odine at high temperatures and pressures. While 1odate
ion is quite stable at room temperature, at elevated temperatures it de-
composes according to the equilibrium reaction

4105~ 4+ 4HT === 21, 4 502 4 2H 0.

The extent of this decomposition in uranyl-sulfate solutions above 200°C
is not known with certainty, since all observations have been made on
samples that have been withdrawn from the system, cooled, and reduced
in pressure before analysis. Although the iodine in such samples is prin-
cipally elemental, some iodate is always present, possibly because of re-
versul of the iodate decomposition as the temperature drops in the sample
line. Such measurements therefore give an upper limit to the iodate con-
tent of the solution. If periodate is introduced into uranyl-sulfate solution.
at elevated temperatures, it 18 reduced before a sample can be taken to
detect its presence. Jodide similarly disappears if an overpressure of oxy-
gen Is present, although iodide to the extent of 409, of the total 1odine has
been found m the absence of added oxygen [11].

Methods that have been used for determining the iodine/iodate ratio in
fuel solutions are (a) analysis of samples taken from an autoclave at 250°C
at measured intervals after injection of iodine in various states [11],
(b) analysis of samples taken from the liquid in liquid-vapor equilibrium
studies at 260 to 330°C [11], (c¢) rapid sampling from static bombs at
250 to 300°C [12], and (d) continuous injection of iodate-containing fuel
solution into the above described ejector loop at 220°C and determining
651 CORE PROCESSING: IODINE 323

oxidation states in samples withdrawn [11]. The iodine/iodate ratio in
these samples has varied from slightly over 1 to about 70, with no apparent
relation to variations in temperature, oxygen pressure, and total iodine
concentration.

The strongest indication of iodate instability was in the loop experi-
ments, which gave the highest observed iodine/iodate ratio, even though
iodine was continuously introduced into the flowing stream as iodate
and removed by oxygen scrubbing as elemental iodine. The low iodate
content of the samples from these experiments corresponded to a first-
order iodate decomposition rate constant of 6.2 min~! Iodate con-
tents averaging about 10% of the total iodine have been observed in
0.04 m U02804—0.005 m CuSO4—H2804 solution, rapidly sampled from
a static bomb through an ice-cooled titanium sample line. The observed
iodate content was unrelated to whether the free sulfuric acid concentra-
tion was 0.02 or 0.03 m, whether the temperature was 250 or 300°C, and
whether or not the solution was exposed to cobalt gamma radiation at an
intensity of 1.7 watts/kg.

Oxidation state of iodine at low temperatures. At 100°C the iodate de-
composition and iodine oxidation are too slow for equilibrium to be es-
tablished in reasonable periods of time. Thus both states can persist under
similar conditions. In stainless-steel equipment both states are reduced to
iodide, which is oxidized to iodine if oxygen or iodate is present {12].

In a radiation field the iodide is oxidized, iodine is oxidized if sufficient
oxygen 1s present, and iodate is reduced [13]. At the start of irradiation,
iodate is reduced, but in the presence of sufficient oxygen, iodine is later
reoxidized to iodate, probably by radiation-produced hydrogen peroxide
which aceumulates in the solution. Ifinally, a steady state is reached with a
proportion of iodate to total iodine which is independent of total iodine con-
centration from 1079 to 10~ % m and temperatures from 100 to 110°C, but
strongly- dependent on uranium and acid concentrations and on the hydro-
gen/oxygen ratio in the gas phase. When the temperature is increased to
120°C there is a marked decrease in iodate stability under all conditions of
gas and solution composition. Experimental data on the effects of radiation
intensity, temperature, and gas composition for the irradiation of a typical
fuel solution containing 0.04 m T02504—0.01 m H2504—0.005 m CuSO4
are given in Ref. 13. The steady-state iodate percentages are also given in
this reference.

6-5.3 Removal of iodine from aqueous homogeneous reactors. It is
clear that under the operating conditions of a power reactor, iodine in the
the fuel solution is mainly in the volatile elemental state. It can therefore
be removed by sweeping it from the solution into a gas phase, stripping
324 CHEMICAL PROCESSING [cuaep. 6

it from the gas stream by trapping it in a solid absorber or by contacting the
gas with a liquid.

Numerous experiments have shown that silver supported on alundum is
a very effective reagent for removing iodine from gas or vapor systems,
although its efficiency is considerably reduced at temperatures below
150°C. Silver-plated Yorkmesh packing is very effective for removing
iodine from vapor streams in the range 100 to 120°C. In one in-pile ex-
periment [14] 909% of the fission-product iodine was concentrated in a
silvered-alundum pellet suspended in the vapor above a uranyl-sulfate
solution. This method of using a solid 1odine absorber, however, would
present difhicult engineering problems, since xenon resulting from iodine
decay would be expected to leave the absorber and return to the core unless
the absorbers were isolated after short periods of use and remotely replaced.

Iodine removal by gas stripping requires a continuous fuel letdown. In
case this is not desirable, the vapor can be stripped of iodine in the high-
pressure system by contacting with a small volume of liquid which is sub-
sequently discharged. Liquids considered include water and aqueous
solutions of alkali, sodium sulfite, or silver sulfate [11]. Although the so-
lutions are much more effective iodine strippers than pure water, their use
requires elaborate provision for preventing entrainment in the gas and sub-
sequent contamination of the fuel solution. Thus most of the effort in
design of iodine-removal systems is based on stripping by pure heavy
water.

One possible iodine-removal scheme uses Oz or Oz 4+ D2 stripping [15].
The iodine is serubbed from the fuel solution by the gas in one contactor
and then stripped from the gas by heavy water in a second contactor. This
water would then be let down to low pressure and stored for decay or proc-
essed to remove 1odine.

In most homogeneous reactors some of the fuel solution is evaporated
to provide condensate for purge of the circulating pump and pressurizer.
Since iodine is stripped from the fuel by this evaporation this operation can
be used for iodine removal. This method, which is illustrated in Fig. 6-12,
has been proposed for the HRE~3 [16]. Here a stream of the fuel solution
is scrubbed with oxygen in the pressurizer. The steam is condensed and the
oxygen recycled. The condensate is distilled to concentrate the 1odine into
such a small volume that its letdown does not complicate reactor operation.

Lodine removal in the HRE-2. lodine adsorption on the platinized alu-
mina recombination catalyst, such as that used in the HRE-2, poisons the
catalyst severely [10]. Although the catalyst can be restored by operation
at 650°C, this would not be feasible in HRE-2 operation. A method for
removing iodine from the gas stream by contact with alundum or York-
mesh coated with silver was developed in the HRT mockup. Iodine was
introduced into the system and vapor from the letdown stream and dump
6-5] CORE PROCESSING: IODINE 325

Oxygen

     
   
 
 
   
 
 
    

Condenser

 

7.53 b/ min
Condenser

Pressurizer

Holdup Tank

To High Pressure
D,OStorage 7.46 Ibfmin

 

 

 

.

I Core Solution
20 gal/min 536°F To Low Pressure
To Reactor Core 596°F System 0,07 Ib/min

Fia, 6-12. Todine removal system proposed for HRE-3.

tank was passed through a silvered alundum bed and the recombiner, and
then to a condenser. Condensate was returned to the high-pressure loop
through a pressurizer and the circulating pump. After injection, the 1odine
concentration of the high-pressure loop dropped from 1.8 mg/liter to
0.1 mg literin 2 hr. In similar experiments with silvered Yorkmesh, 1odine
levels in the condensate and pressurizer were cven lower relative to the
high-pressure loop. The Yorkmesh efficiency depended strongly on how
densely 1t was packed. The iodine removal efliciencies calculated from
these experiments and others are given in Table 6-8. In laboratory ex-
periments with a I-in.-diameter bed which could not be tightly packed,
Yorkmesh efliciencies were consistently poorer than those of silvered
alundum.

The ability of a bed of silver-plated Yorkmesh to remove iodine from the
reactor system was apparently confirmed during the initial operating
period of the HRE-2 [17]. Here the iodine activity in the reactor fuel
appeared to be even lower than cxpected when iodine was removed at
the same fractional rate as fuel solution was let down from the high-
pressure system. Less than 3% of the iodine produced during 40 Mwh of
operation was found in the fuel solution. Experience with the HRT
mockup indicates that the iodine not in solution was held on the silvered
bed.
326 CHEMICAL PROCESSING [cHAP. 6

TaBLE 6-8

Iopine REmovaL Erriciency orF SiLverep Beps 1N HRT Mockup

 

 

 

Absorber Bed helght, Temp:arature, Efficiency,
in. C Yo
Silvered alundum 8 150 97.7
rings 8 120 81.0
5 110 64.0
Yorkmesh, 22 1b/ft3 10 120 97.0
Yorkmesh, 29 Ib/ft3 6 120 99.6

 

 

 

 

 

 

6—6. URANYL SULFATE BLANKET PRoCEsSING*

6-6.1 Introduction. The uranyl sulfate blanket solution of a plutonium
producer is processed to remove plutonium and to control the neutron
poisoning by corrosion and fission products. Although a modified Purex
solvent extraction process can be used for plutonium removal, the method
shown schematically in Fig. 6-3, based on the low solubility of plutonium
in uranyl sulfate solution at 250°C, appears more attractive. A hydroclone
similar to that used for reactor core processing is used to produce a con-
centrated suspension of PuO2 along with solid corrosion and fission prod-
ucts. The small volume of blanket solution carrying the plutonium is
evaporated to recover the heavy water and the solids are dissolved in
nitric acid. After storage to allow Np?3? to deecay, plutonium is decon-
taminated by solvent extraction.

6—6.2 Plutonium chemistry in uranyl sulfate solution. The amount of
plutonium remaining dissolved in 1.4 m U050, at 250°C is dependent
on a number of variables, including solution acidity, plutonium valence,
and initial plutonium concentration. Under properly controlled condi-
tions, less than 3 mg/kg H20O has been obtained. Since plutonium is re-
moved from solution by hydrolysis to PuQOg, solubilities are increased by
increasing the acidity. Table 6-9 summarizes data on the solubility be-
havior of plutonium for various acidities.

*Contribution from R. E. Leuze.
6-6] URANYL SULFATE BLANKET PROCESSING 327

TABLE 6-9

SOLUBILITY OF TETRAVALENT PLUTONIUM
IN 1.4 m UO2804 AT 250°C

 

Excess sulfuric acid, Pu(IV) solubility,
m mg/kg H20

 

 

3.7
17
39
68

105

oo
= 0 D

 

 

 

 

Plutonium behavior is difficult to predict because of its complex valence
pattern. In the absence of irradiation, plutonium dissolved in 1.4 m U02804
under a stoichiometric mixture of hydrogen and oxygen at 250°C exists in
the tetrapositive state. However, when dissolved chromium is present or
when an overpressure of pure oxygen is used, part of the plutonium is oxi-
dized to the hexapositive state. Experiments indicate that in the presence
of Co% gamma irradiation [18], reducing conditions prevail even under an
oxygen pressure and plutonium is held in the tetrapositive state. The
valence behavior discussed here is somewhat in question, since actual
valence measurements were made at room temperature immediately after
cooling from 250°C. It is known that tetrapositive plutonium will dispro-
portionate upon heating [19]. The disproportionation in a sulfate system
is depressed by the sulfate complex formation with tetrapositive plu-
tonium. These results indicate that plutomium in a reactor will be pre-
dominantly in the tetrapositive state.

When the plutonium concentration exceeds the solubility limit, plu-
tonium will hydrolyze to form small particles of PuOs about 0.5 micron
in diameter and in pyrex, quartz, or gold equipment forms a loose preci-
pitate with negligible amounts adsorbed on the walls. However, if these
solutions are contained in type-347 stainless steel, titanium, or Zirealoy,
a large fraction of the PuOg adsorbs on and becomes incorporated within
the oxide corrosion film. Attempts to saturate these metal surfaces with
plutonium in small-scale laboratory experiments were unsuccessful even
though plutonium adsorption was as much as 1 mg/em?.

6-6.3 Neptunium chemistry in uranyl sulfate solution. Neptunium dis-
solved in 1.4 m UO2S804 at 250°C under air, stoichiometric mixture hy-
drogen and oxygen, or oxygen is stable in an oxidized valence state, prob-
328 CHEMICAL PROCESSING [cHAP. 6

ably Np(V). The solubility is not known, but it is greater than 200 mg/kg
H20. Since the equilibrium concentration is only about 50 mg/kg H0,
for a 1.4 m UO2304 blanket with an average flux of 1.8 X 10'* neu-
trons/(em?)(sec), all the neptunium should remain in solution in most
reactor designs.

6-6.4 Plutonium behavior under simulated reactor conditions. Pluto-
nium behavior in actual uranyl sulfate blanket systems has not been
studied; however, small-scale static experiments with 100 ml of solution
and circulating loop experiments with 12 liters of solution have been car-
ried out in the absence of irradiation under conditions similar to those ex-
pected in an actual reactor.

In the static experiments, plutonium was added batchwise to 1.4 m
U02804 at a rate of about 6 mg/kg HoO/day. The solution was heated
overnight in a pyrex-lined autoclave at 250°C under 200 psi hydrogen and
100 ps1 oxygen. The solution was cooled to room temperature for analysis
and for adding more plutonium. This was repeated until a total of 140 mg
of plutonium per kilogram of water was added. Small disks of type-347
stainless steel were suspended 1n the solution throughout the experiment to
determine the amount of plutonium adsorption. The behavior of plu-
tonium for a stainless-steel surface area/solution volume ratio of 0.6 cm?2/ml
is shown in Fig. 6-13. As the plutonium concentration was gradually in-
creased to 45 mg/kg H20, essentially all the plutonium remained in solu-
tion as Pu(VI). There was a small amount of adsorption, but no precipita-
tion. During the next few additions the amount of plutonium in solution
decreased rapidly to about 5 mg/kg Hz0. At the same time there was a
rapid increase in plutonium adsorption and in the formation of a loose
PuO, precipitate. All plutonium added after this was either adsorbed or
precipitated.

Other experiments were made with surface/volume ratios of 0.2 and
0.4 em?/ml. In all cases, the plutonium remaining in solution and the plu-
tonium adsorption per square centimeter were essentially the same as
that shown in Fig. 6-13. Thus, by decreasing the surface/volume ratio,
it is possible to increase the amount of plutonium in the loose precipitate.
For example, when the total plutonium addition was 130 mg/kg H20, 40%
of the plutomium was as a loose precipitate for a surface/volume ratio of
0.6 em?/ml, 609% for a ratio of 0.4 em?/ml, and 689 for a ratio of
0.2 em?/ml.

Plutonium behavior under dynamic conditions was studied by injecting
dissolved plutonium sulfate and preformed PuQOg into a circulating stream
of 12 liters of 1.4 m UO2S04 at 250°C under 350 psi oxygen. This solution
was contained 1n a type—347 stainless steel loop equipped with a canned
rotor pump, a hydroclone, metal adsorption coupon holders, and a small
ti—t3] TRANYL SULFATE BLANKET PROCESSING 329

60

 

| l I | ! I (

\ 1.4 m U0,80, at 250°C
100 psiO4 4+ 200 psiH

2 2 9
Surface to Volume 0.6 cm “/ml

p:- S
o
|

\
\

In Solution,
mg/kg HyO

3
W

\

i ]
\

i

 

0 | | 1 ! ] T T
1000 | ] 1 { | | {

 

 

100 |—

Adsorbed,
ng/cm?
o
[

 

0.1
60

 

 

40 |—

20

In Precipitate,
mg/kg H20

 

 

 

| | l l | | i
0 20 40 60 80 100 120 140 160
Total Plutonium Added, mg/kg H20

 

Fia. 6-13. Plutonium behavior in uranyl sulfate solution contained in type-347
stainless steel.

pressure vessel that could be connected and removed while the loop was in
operation. Plutonium was added and ecirculating-solution samples were
taken through this vessel. Tetrapositive plutonium added to the circulating
solution was completely oxidized to hexapositive i less than 5 min. When
45 mg/kg Ho0 of dissolved plutonium was added every 8 hr, the amount
of plutonium circulating i solution increased to a maximum of about
150 mg/keg Ho0. As more plutonium was added, it was rapidly adsorbed
on the loop walls. After the last addition of plutonium, the loop was
operated at 250°C for several days. Twelve hours after the last addition
the plutonium concentration had decreased to 100 mg/ kg H»>0, and about
40 hr later the amount of plutonium in solution had dropped to an ap-
parent equilibrium value of 60 mg/kg H»0. Essentially all the plutonium
removed from solution was adsorbed on equipment walls uniformly
throughout the loop. Less than 0.19, of the plutonium was removed in
330 CHEMICAL PROCESSING [caAP. 6

the hydroclone underflow, and no precipitated solids were circulating.
Even when 850 mg of plutonium as preformed PuQ» was injected into the
loop, no circulating solids were detected 5 min later. Only 209, of this plu-
tonium was removed by the hydroclone, 359 was adsorbed on the stainless
steel, and the rest was distributed throughout the horizontal sections of
the loop as loose solids. The hydroclone was effective for removing solids
that reached it, but the loop walls and low veclocity in horizontal pipes
were effective traps for PuO..

There are several differences in conditions between the loop runs and an
actual reactor, the most important of which are probably the presence of
radiation, the lower surface/volume ratio (0.4 compared with 0.8 cm?/ml
for the loop), the slower rate of plutonium growth in the reactor (12 to
15 mg/kg H20/day) and the probability that a plutonium producer
would have to be constructed of titanium and Zircaloy to contain the con-
centrated uranyl-sulfate solution. Based on these laboratory results, how-
ever, 1t appears that plutonium adsorption on metal walls may be a serious
obstacle to processing for removal of precipitated PuQs.

6-6.5 Alternate process methods. Because of the problem of plutonium
adsorption on metal walls, removal methods based on plutonium concen-
trations well below the solubility limit have been considered. In a full-
scale reactor plutonium will be formed at the rate of up to 12 to 15
mg/kg H20/day. In order to keep the plutonium concentration below
3 mg/kg H20, the entire blanket solution must be processed at least four
to five times a day. By adding 0.4 m excess H2S04 (see Table 6-9), the
plutonium solubility is increased to greater than 100 mg/kg H.O and the
blanket processing rate can be decreased to once every 3 or 4 days. Slightly
longer processing cycles can be used if part of the plutonium is removed as
neptunium before it decays.

Of the various alternate processes considered, ion exchange and ad-
sorption methods show the most promise. Dowex—-50 resin, a strongly
acidic sulfonic acid resin loaded with UO2" 1, completely removed tetra-
positive plutonium from 1.4 m UO2504 containing 20 mg of plutonium
per liter [20]. The resin capacity under these conditions, however, has not
been determined. Because of the high radiation level it may not be feasible
to use organic resins. Sorption of plutonium on inorganic materials shows
some possibilities as a processing method {21]. Although rather low plu-
tonium /adsorber ratios have been obtained, indications are that capacities
will be significantly higher at higher plutonium concentrations. Special
preparation of the adsorbers should alsc increase capacities. Attempts to
coprecipitate plutonium with tri- or tetrapositive iodates, sulfates, oxalates,
and arsenates were not successful, owing to the high solubilities of these
materials in 1.4 m UO28504.
6-6] URANYL SULFATE BLANKET PROCESSING 331

r—=To Off Gas T
-+

   

Irridiated
Tharium

  

Scrubber
Additive Condenser

Dissolvent —

 

 

7
}

 

1

 

Acid | e——Acid
Catch Fractionator

 

 

 

 

| 1
Dissolver > Waste

 

e e ——————

§

| To Solvent

| ,

| | Extraction .

] Reboiler
te——Feed

Adjustment
Tank

 

 

 

 

 

 

Recovered

l
f
LACId Tank

 

 

 

 

 

Fi1a. 6~14. Thorex process, feed preparation flowsheet.

% - --—— First Cycle +———— Second Cycle — \-i
‘ I

Aqueous Agueous Agueous Aqueous Aqueous
Scrub Strip Scrub Strip Strip

 

 

 

 

 

 

 

 

 

 

 

 

   
 

   
   

To Orgonic

     

From Feed
Preparation

  

Solvent To
Th Th Recovery First
Cycle

 
 

Th, U

  

 

 

 
 

Extraction
|
Stripping
Extraction
Partitioning
Stripping

 

  
 
 

 

Reductant

 

       
 
 

  

 

 

 

 

 

 

 

- Th iTh
U Organic Organic | | Uranium
x Extract- Scrub I Isolation
Waste ont
Sioc:cr:ge Evaporator Evaporator Il.'..-,. Th Product

 

Waste

F1c. 6-15. Thorex process, solvent extraction co-decontamination flowsheet.
332 CHEMICAL PROCESSING [cHAP. 6

-Third Cycle ——————————=

e isolation —— - ———————
! 1

Elutriant Elutriant Agueous
- Scrub

           

    

From
Solvel_fl Effluent T
Extraction to Waste ;
Solvent
Recovery

  
 

Third Cycle
Feed
Preparation

Organic
Extractant

 
     
 
  

U Sorption Column
Extraction Column

 
 
 

Th Removal Column

  

   
 

 
 

Waste v
or
Storage
To Storage Shigronen'r

Fic. 6-16. Thorex process, uranium isolation and third cycle flowsheet.

6-7. TuoriuM OXIDE BLANKET PROCESSING

6-7.1 Introduction. At the present the only practical method available
for processing irradiated thorium-oxide slurry is to convert the oxide to a
natural water-thorium nitrate solution and treat by the Thorex process.
Although this method is adequate, it is expensive unless one plant can be
built to process thorium oxide from several full-scale power reactors.
Therefore methods for ThO2 reprocessing which could be economically in-
corporated into the design and operation of a single power station have
been considered. Alternate methods that have been subjected to only brief
scouting-type experimentation are discussed in Article 6-7.3.

6-7.2 Thorex process.* The Thorex process has been developed to sep-
arate thorium, U233, fission product activities, and Pa?33; to recover the
thorium and uranium as aqueous products suitable for further direct
handling; and to recover isotopically pure U2 after decay-storage of the
Pa233, The flowsheet includes two solvent-extraction cycles for thorium
and three solvent-extraction cycles plus lon exchange for the uranium.
Although only irradiated thorium metal has been processed, the process is
expected to be satisfactory for recovery of thorium and uranium from
homogeneous reactor tuels.

The Thorex process may be divided into three parts: feed preparation,

*(Contribution from W. T. McDuffee.
6-7] THORIUM OXIDE BLANKET PROCESSING 333

solvent extraction, and product concentration and purification. These
three divisions are shown in Figs. 6-14, 6-15, and 6-16.

In the feed preparation step, uranyl sulfate solution from the reactor core
and thorium oxide from the blanket system, freed of D»0 and suspended in
ordinary water, are fed into the dissolver tank. The dissolvent is 13 N
nitric acid to which has been added catalytic amounts (0.04 N) of sodium
fluoride. When short-cooled thorium is being processed, potassium iodide
is added continuously to the dissolver to provide for isotopic dilution of
the large amount of fission-produced 1'3! which is present. The dissolver
solution is continuously sparged with air, and the volatilized iodine is re-
moved from the off-gases in a caustic scrubber.

The dissolver solution is transferred to the feed adjustment tank where
aluminum nitrate is added, excess nitric acid recovered, and the resultant
solution made slightly acid-deficient by evaporating until a temperature
of 155°C is reached. During digestion in the feed adjustment tank any
silica present is converted to a form that will not cause emulsion problems
in pulse columng, and fission products generally are converted to forms
less likely to be extracted by the solvent (4297 TBP in Amsco).

In the solvent extraction step thorium and uranium are co-extracted in
the first cycle; subsequent partitioning of thorium and uranium in the
second cycle gives two decontaminating cycles to both products while
using only five columns. For short-decayed thorium a reductant, sodium
hydrogen sulfite, is continuously added to the feed streams of both cycles
to decrease the effect of nitrite formed by irradiation. Without the sulfite
addition, the nitrite formed by radiation decomposition of nitrates con-
verts ruthenium to a solvent-extractable form. Aecid deficiency in the
second cycle feed is achieved by adding dibasic aluminum nitrate (diban).

The spent organic from the second cycle is recycled to the first cycle as
the organic extractant. The spent solvent from the first cycle is processed
through a solvent-recovery system and reused as the organic extractant in
the second cycle.

In the uranium product concentration and purification step (Fig. 6-16),
uranium is isolated by ion exchange, using upflow sorption and downflow
elution. In this way & concentrated uranium solution in 6 N HNO3 is ob-
tained. This solution is stable enough for storage or is suitable as a feed
for the third uranium extraction cycle. The third uranium cycle is a
standard extraction-stripping solvent-extraction system using 15% TBP-
Amsco as the organic extractant. Although installed as a part of the com-
plete Thorex flowsheet, the third cycle may be used separately for re-
processing long-stored uranium to free it of objectionable decay daughters
of U232, When used as an integral part of the Thorex scheme, additional
decontamination of the uranium is achieved and the nitrate product is
well adapted for extended storage or future reprocessing.
TaBLE 6-10

AVERAGE DECONTAMINATION FacTors rFor THorRIUM AND URANIUM PRODUCTS
IN THE THOREX Piror PLANT

Thorium irradiated to 3500 grams of mass—-233 per ton, two complete cycles for both uranium and thorium,

one additional uranium cycle for material decayed only 30 days.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Decontamination factors
Gross Pa, Ru Zr-Nb Total rare earths I
Thorium
400 days decayed 1 x 10° I x 104 4 x 103 3% 105 2% 108 —
30 days decayed 4 x 10* 7 X 10% 200 3 X 104 2 % 108 9 x 108
Uranium-233
400 days decayed 3 X 105 3 X 105 2 X 10° 8 x 10% 9 x 108 —
30 days decayed 5 x 107 5% 1010 4 x 108 7 X 1(® 3 X 108 3 x 107

 

 

¥Ee

DNISSTI0dd TVOIRAHD

g "d¥HO]
6-7] THORIUM OXIDE BLANKET PROCESSING 335

For return to an aqueous homogeneous reactor the decontaminated
uranium would probably be precipitated as the peroxide, washed free of
nitrate, and then dissolved in D2SO4 and D20O. Product thorium would
be converted to thorium oxide by methods described in Section 4-3.

The adaptability of the Thorex flowsheet just described to processing
thorium irradiated to contain larger amounts of U23% per ton and decayed
a short time has been demonstrated in the Thorex Pilot Plant at Oak Ridge
National Laboratory [22]. Fifteen hundred pounds of thorium irradiated
to 3500 grams of U233 per ton and decayed 30 days was processed through
two thorium cyecles and three uranium cycles. The decontamination fac-
tors for various elements achieved with short-decayed material are com-
pared in Table 6-10 with results obtained with longer-decayed material.
While the decontamination factors obtained with the short-decayed ma-
terial compare favorably with the factors for the long-decayed material,
the initial activity in the short-decayed thorium was 1000 times greater
than in the long-decayed. Therefore, while the thorium and uranium
products did not meet tentative specifications after two complete cycles,
the uranium product did meet those specifications after the third uranium
cycle. Since the chemical operations necessary to convert these materials
to forms suitable for use in a homogeneous reactor can be carried out re-
motely, the products are satisfactory for return to a homogeneous reactor
after two cycles.

6-7.3 Alternate processing method.* Attempts to leach protactinium
and uranium produced in ThO» particles by neutron irradiation [23] in-
dicate that both are rather uniformly distributed throughout the mass of
the ThOs particle, and migration of such ions at temperatures up to 300°C
is extremely slow. Since calculations show that the recoil energy of frag-
ments from U233 fission is sufficiently large to eject most of them from a
particle of ThO2 not larger than 10 microns in diameter, this offers the
possibility of separating fission and corrosion products from a slurry of
ThO. without destroying the oxide particles. Such a separation, however,
depends on the ability to remove the elements that are subsequently ad-
sorbed on the surface of the ThOs. Adsorption of various cations on ThO»
and methods for their removal are discussed in the following paragraphs.

Trace quantities of such nuclides as Zr® Nd'#7, Y°!, and Ru'®® when
added to a slurry of ThO»s in water at 250°C are rapidly adsorbed on the
oxide particles, leaving less than 10749 of the nuelides in solution. The
tracer thus adsorbed cannot be eluted with hot dilute nitric or sulfuric acid.
The adsorption of macroscopic amounts of uranium or neodymium on
ThO; at 250°C is less for oxide fired to 1600°C than for 650°C-fired oxide,

*Contribution from R. E. Leuze.
336 CHEMICAL PROCESSING [cHAP. 6

TABLE 6-1]

ErrFect OoF CALCINATION TEMPERATURE ON
URANIUM AND NEODYMIUM ADSORPTION ON THQs aT 250°C

0.5 g of ThO; slurried at 250°C in 10 ml of 0.005 m Nd(NO3)3
or 0.0bm U02804—0.05 m HzSO4.

 

 

 

 

1 *

Calcination temperature, Adsorption, mg/g Th
°C

U Nd

650 3.3-4.4 7.4

850 1.9-2.4 6.1

1000 0.72-1.10 2.4

1100 0.08-0.19 0.5

1600 0.06-0.12 0.3

 

 

 

 

 

*Single numbers represent data from single experiments. In other cases the range
for several experiments is given.

TABLE 6-12

Use or PBO 10 DECREASE CATION ADSORPTION ON THO»

0.2 g of ThO; plus various amounts of PbO coslurried in 10 ml of solution at
250°C for 8 hr.

 

 

 

. PhO/ThO, . Cation adsorbed on ThOs,
Solids . Solution

wt. ratio ppm
ThO» 0.002 m UO2804 3100
PbO + ThO- 0.2 (0.002 m UO2804 220
ThOz 0.001m CE(N03)3 6200
PbO + ThO. 0.2 0.001 m Ce(NO3)s 10
ThO, (.01 m Nd tartrate 2700
PbO + ThO- 0.4 0.01 m Nd tartrate 10

 

 

 

 

 

 
6-71 THORIUM OXIDE BLANKET PROCESSING 337

as illustrated in Table 6-11. This change in amount of adsorption may be
almost entirely due to decrease in surface area of ThO; with increased
firing temperature. The surface area of 1600°C-fired ThO3 is only 1 m?/g
ThO3,, while the 650°C-fired ThO» has a surface area of 35 m2/g ThO,.

The cation adsorption on ThO;z can be decreased by coslurrying some
other oxide with the ThO2. The added oxide must adsorb fission products
much more strongly than ThO, and be easily separable from ThOz. The
effectiveness of PbO in decreasing cation adsorption on ThO; is shown in
Table 6-12. When PbO; was used, more than 999 of the cations added
to the ThO2-PbO slurry was adsorbed on the PbOa. However, cations
adsorbed on ThOgz were not transferred to PbO2 when it was added to
slurry in which the cations were already adsorbed on the ThO, particles.
Addition of dilute nitric acid to the ThO2-PbO coslurry completely dis-
solved the PbO and the cations adsorbed on it without disturbing the ThOs-.

In all cases, cations adsorbed on ThO: at 250°C are so tightly held that
dilute nitric or sulfuric acid, even at boiling temperature, will not remove
the adsorbed material. However, the adsorbed ions can be desorbed by
refluxing the ThO» in suitable reagents under such conditions that only a
small amount of 1600°C-fired ThOs is dissolved. Under the same treat-
ment ThO» fired to only 650°C would be 90% dissolved.
338 CHEMICAL PROCESSING [cHAP. 6

REFERENCES

1. A. T. Gresky and E. D. Ar~NoLp, Poisoning of the Core of the Two-region
Homogeneous Thermal Breeder: Study No. 2, USAEC Report ORNL CF-54-2-208,
Oak Ridge National Laboratory, 1954.

2. E. D. Arnvouwp and A. T. Gresky, Relative Biological Hazards of Radiations
Ezpected in Homogeneous Reactors TBR and HPR, USAEC Report ORNL-1982,
Oak Ridge National Laboratory, 1955.

3. R. A. MoNErs and S. PETERsoN, in Homogeneous Reactor Project Quarterly
Progress Report for the Period Ending July 81, 1955, USAEC Report ORNL-1943,
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4. D. E. Fercuson et al., in Homogeneous Reactor Project Quarterly Progress
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5. P. A. Haas, Hydraulic Cyclones for Applicatton to Homogeneous Reactor
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6. W. D. BurcH et al., in Homogeneous Reactor Project Quarterly Progress
Report for the Period Ending July 31, 1957, USAEC Report ORNL-2379, Oak
Ridge National Laboratory, 1957. (p. 26)

7. R. A. McNEeEs ct al,, Oak Ridge National Laboratory, in Homogeneous
Reactor Project Quarterly Progress Report, USAEC Reports ORNT-2379, 1957
(p. 137); ORNL-2432, 1957 (p. 147); ORNL-2493, 1958.

8. W. E. Browning and C. C. Bovta, Measurement and Analysis of Holdup
of Gas Mixtures by Charcoal Adsorption Traps, USAEC Report ORNL-2116,
Oak Ridge National Laboratory, 1956.

9. W. D. BurcH et al.,, Oak Ridge National Laboratory, in Homogeneous
Reactor Project Quarterly Progress Report, USAEC Report ORNL-2432, 1957
(p. 23); ORNL-2493, 1958.

10. P. H. Harrey, HRT Mock-up Iodine Removal and Recombiner Tests,
USAEC Report CF-58-1-138, Oak Ridge National Laboratory, 1958.

11. R. A. KxELER et al., Vitro Laboratories, 1957. Unpulblished.

12. S. PerERrsoN, unpublished experiments.

13. D. E. FErRGUsON et al,, Oak Ridge National Laboratory, in Homogeneous
Reactor Project Quarterly Progress Report, USAEC Reports ORNL-2272, 1957
(pp. 133-135); ORNL-2331, 1957 (pp. 142-143, 148-149); ORNL-2379, 1957
(pp. 138-139).

14. R. A. McNEEs et al., in Homogeneous Reactor Project Quarterly Progress
Report for the Period Ending Apr. 30, 1955, USAEC Report ORNL-1895, Oak
Ridge National Laboratory, 1955. (pp. 175-176)

15. D. E. FErauson, Removal of Todine from Homogeneous Reactors, USAEC
Report CF-56-2-81, Oak Ridge National Laboratory, 1956; Preliminary Design
of an Todine Removal System for a 460-Mw Thorium Breeder Reactor, USAEC
Report CF-56-7-12, Oak Ridge National Laboratory, 1956.

16. H. O. WEEREN, Preliminary Design of HRE-3 Iodine Removal System
#1, USAEC Report CF-58-2-66 Oak Ridge National Laboratory, 1958.

17. S. PETERSON, Behavior of Iodine in the HRT, USAEC Report CF-58-3-75,
Oak Ridge National Laboratorys 1958.
REFERENCES 339

18. R. E. Leuze et al.,, in Homogeneous Reactor Project Quarterly Progress
Report for the Period Ending Apr. 30, 1957, USAEC Report ORNL-2331, Oak
Ridge National Laboratory, 1957. (p. 151)

19. R. E. CoxnNickg, in The Actinide Elements, National Nuclear Energy Series,
Division 1V, Volume 14A. New York: McGraw-Hill Book Co., Inc., 1954,
(pp. 221, 238-241)

20. 8. 8. Kigrsuis, Oak Ridge National Laboratory. Unpublished.

21. R. E. Levuze and 8. 8. Krrsuis, Oak Ridge National Laboratory, 1957.
Unpublished.

22, W. T. McDvurrer and O. O. Yarsro, Oak Ridge National Laboratory,
1958. Unpublished.

23. D. E. Ferguson et al., in Homogeneous Reactor Project Quarterly Progress
Report for the Period Ending July 31, 1955, USAEC Report ORNL-1943, Oak
Ridge National Laboratory, 1955. (p. 221)