Pary 111

LIQUID-METAL FUEL REACTORS

 

FrRANK MASLAN,

Editor

Brookhaven N atvonal Laboratory

18. Liquid-Metal Fuel Reactors

19. Reactor Physics for Liquid-Metal Reactor Design
20. Composition and Properties of Liquid-Metal Fuels
21. Materials of Construction—Metallurgy

22. Chemical Processing
23. Engineering Design

24. Liquid-Metal Fuel Reactor Design Study
25. Additional Liquid-Metal Reactors

. BOURDEAU
. B. BRODSKY
. S. BRYNER
. CHERNICK

CONTRIBUTORS

M. JANES

O. F. KAMMERER
C. J. Kramur

R. M. KieuN

. L. MANSFIELD
. A. MEYER

. T. MILES

. RASEMAN

. RoBBA

. G. SCHWEITZER
. V. SHEEHAN

. SUSSKIND

pepmeusowww
PREFACE

This is the most extensive discussion of liquid-metal fuel reactor devel-
opment yet published in the United States. Emphasis has been placed on
the Liquid Metal Fuel Reactor being developed by Brookhaven National
Laboratory and Babcock & Wilcox Co. because it is the most advanced
project. Work on various phases of liquid-metal fuel reactors is being
carried out by Los Alamos Scientific Laboratory, Raytheon Manufacturing
Co., Argonne National Laboratory, Ames Laboratory, and Atomics
International. The editor would like to have given more coverage to work
at the last three locations but was unable to because time was lacking.

The liquid-metal fuel reactor development at Brookhaven started as
an organized program in 1951. Before that, work had been conducted on
bismuth-uranium fuel and other components. In 1954, Babcock & Wilcox
Co., in collaboration with representatives of sixteen other companies,
prepared a reference design and report. In 1956, Babcock & Wilcox con-
tracted with the Atomic Energy Commission to design, build, and operate
a low-power experimental reactor (LMFR Experiment No. 1). Research,
development, and design studies are being carried on concurrently by
B & W and Brookhaven. LMFR Experiment No. 1, on which construc-
tion is scheduled to start in 1960, is intended to demonstrate feasibility
and provide information on the physics, metallurgy, chemistry, and
mechanical aspects of this type of reactor.

The editor expresses appreciation to many of his colleagues at Brook-
haven and Babcock & Wilcox for working with him on these chapters. He
wishes particularly to thank those whose material he drew upon, also
C. Williams, O. E. Dwyer, D. Gurinsky, H. Kouts, F. T. Miles, and T. V.
Sheehan, of Brookhaven National Laboratory; R. T. Schoemer, H. H.
Poor, and J. Happell, of Babcock & Wilcox Co.; R. Rebholz and G. Goring,
of Union Carbide Corp.; D. Hall, of Los Alamos Scientific Laboratory;
and W. Robba, of Raytheon Manufacturing Co. Special appreciation is
due Miss Gloria Ministeri for her laborious and prolonged secretarial
work and Miss Dolores Del Castillo for coming to our aid in emergencies.

Upton, New York Frank Maslan, Edztor
June 1958 | ‘
CHAPTER 18

LIQUID METAL FUEL REACTORS

18-1. BACKGROUND

Liquid metal fuel reactors have received attention since the early days
of reactor technology. The concept of a high-temperature fluid fuel which
could be circulated for both heat exchange and chemical processing has
been an intriguing one [1-4].

This type of reactor was first suggested in 1941 but received little research
and development attention until approximately 1947. At this time the
Nuclear Engineering Department at Brookhaven National Laboratory
began its Liquid Metal Fuel Reactor (LMFR) development. A solution
of uranium in bismuth was suggested because of the low melting point and
low neutron-capture cross section of bismuth. Coupled with these factors
1s the very high boiling point of bismuth, which makes possible the high-
temperature operation of a bismuth-cooled reactor at relatively low
pressures.

Modern steam power plants have a thermodynamic efficiency of approxi-
mately 409,. For a nuclear system to achieve comparable efficiencies, the
working fluid will have to have a reactor outlet temperature in the neigh-
borhood of 500°C. The LMFR is one of the new types of nuclear reactors
having this desirable characteristic. Thus, it is one of the few with poten-
tialities for producing power competitive with the best of the present steam
systems. | '

18-1.1 Work at Brookhaven National Laboratory. In 1948, an appraisal
of various low-melting alloys was made at Brookhaven. Attention was also
given to metallic slurries consisting of uranium in the form of intermetallic
compounds suspended in liquid metal carriers. The uranium-bismuth
system appeared to show considerable promise. Preliminary solubility
studies were completed by 1950 and a start was made on fuel processing
investigations.

Since that time the project has steadily accelerated. Chemical aspects
of the fuel and fuel-processing systems have been and are being investigated
in considerable detail. Metallurgical studies of corrosion, mass transfer,
and stability of fuel systems have advanced from short-time crucible tests
to circulating loops of alloy steel operated for many thousands of hours.
Consideration has also been given to the design of such various reactor
components as pumps, piping, valves, heat exchangers, and instruments.

703
704 LIQUID METAL FUEL REACTORS [cHAP. 18

18-1.2. Work of study groups. In common with other reactor concepts,
the LMFR has been evaluated from time to time as part of the general
Atomic Energy Commission Reactor Development Program. During the
summer of 1953, the LMIR was evaluated under Project Dynamo, and it
was concluded that it was an extremely attractive concept if proven tech-
nically feasible. In 1955 an industrial study group, under the direction of
Babcock & Wilcox, made a detailed appraisal and design of the LMIR
concept [19], and reported that it could be proved technically feasible in
the near future and that it appears attractive from an economic point of
view. In 1957, the Babcock & Wilcox Company re-evaluated the LMFR
and found the outlook as good as indicated previously [21]. Of course, the
development of a new reactor concept of this kind is a long-range program.

Present plans call for a buildup of knowledge through the construction
and operation of several LMFR experiments. The first of these is currently
being designed by Babcock & Wilcox.

18-2. GENERAL CHARACTERISTICS OF LI1QUID METAL
FuerL ReEAacTORs*

18-2.1 Comparison of fluid- and solid-fuel reactors. In order to better
understand the development and characteristics of the Liquid Metal
Fuel Reactor, fluid- and solid-fuel reactors should be compared, and a
distinction should be made between the features of fluid fuels in general
and those of liquid metal fuels in particular.

A reactor using a fluid fuel may have the following advantages over one
with solid-fuel elements:

(1) Simple structure. A fluid fuel can be cooled in an external heat
exchanger separate from the reactor core. Thus the nuclear requirements
(of the core) and the heat flow requirements (of the exchanger) need not
both be satisfied at the same place. This may allow design for very high
specific power. For example, material of high cross section, such as tung-
sten or tantalum, which could not be used in the core, could be used in the
heat exchanger.

(2) Easy fuel handling.

(3) Simplified reprocessing. The reduction to metal, fabrication, canning,
and dissolving steps are eliminated. Because manual steps in refabrication
are unnecessary, decontamination need not be complete. The cooling time
could be made much shorter, resulting in a smaller holdup of fissionable
material.

(4) Simplified waste disposal.

(5) Continuous removal of fission products. The removal of poisons
would improve neutron economy and permit higher burnup. With a lower

*Contributed by F. T. Miles, Brookhaven National Laboratory.
18-2] CHARACTERISTICS OF LIQUID METAL FUEL REACTORS 705

inventory of radioactive material, the potential hazard would be decreased;
this might reduce the size of the exclusion area required for safety.

(6) Inherent safety and ease of control. Any liquid fuel which expands
on heating gives an immediate negative temperature coefficient of re-
activity. This effect is not delayed by any heat-transfer process. The
rate of expansion is limited only by the speed of sound (shockwave) in the
liquid. This instantaneous effect tends to make the reactor self-regulating.
Adjustment of fuel concentration can be used as an operating control.

Disadvantages of fluid fuels are listed below:

(1) Possible fluctuations of reactivity caused by density or concentra-
tion changes in the fuel, e.g., bubbling.

(2) Loss of delayed neutrons in the fuel leaving the core.

(3) External holdup of fissionable material.

(4) Induced activity in pumps and heat exchangers and possible de-
position of fuel and fission products.

(5) Corrosion and erosion problems. Each fuel system has its particular
corrosion problems. These differ greatly from one system to another, but in
every case corrosion is a critical problem which must be solved.

(6) High radiation levels in the reactor and in the component piping
require development of remote maintenance techniques.

18-2.2 Advantages and disadvantages of LMFR. Comparing one liquid
fuel system with another involves relative advantages and disadvantages.
Liquid metal solution systems (in particular, solutions of uraniumn in
bismuth) [5-12] have the following advantages over aqueous systems:

(1) Metals can be operated at high temperatures without high pressures.

(2) Metal solutions are free from radiation damage and do not give off
bubbles. By using liquid metals, therefore, two factors that may limit the
specific power of aqueous systems are avoided.

(3) Liquid metals have better heat-transfer properties than water.

(4) Metal systems do not have inherent moderating properties and can
be used for fast and intermediate reactors as well as for thermal reactors,
provided the critical mass requirements are not excessive.

(5) Liquid metals can be circulated by electromagnetic pumps if desired,
although the efficiency may be poor, as with bismuth.

(6) Some suitable metals, e.g., bismuth, are cheaper than D20.

(7) Polonium, formed from bismuth by neutron capture, may be a
valuable by-product.
~ Liquid-metal systems have the following disadvantages in comparison
with aqueous systems:

(1) The heat capacity is less than with water.

(2) The higher density may be a disadvantage.

(3) Liquid metals are more difficult to pump.
706 LIQUID METAL FUEL REACTORS [cHAP. 18

(4) The absorption cross sections of the best metals (e.g., bismuth
. = 0.032 barn) are inferior to D20, although better than H2O. The cross
section of bismuth may be low enough, however, to allow breeding of U233
from thorium by means of thermal neutrons.

(5) For a thermal reactor, moderator must be supplied.

(6) The limited solubility of uranium in bismuth necessitates the use of
enriched U235 or U233 as fuel. Uranium—-238 or thorium cannot be held in
solution in sufficient concentration to give internal breeding.

(7) Because of items (4) and (5) above, liquid metal fuel reactors are at
least 2 ft in diameter [13] and cannot be scaled down as far as aqueous
reactors can.

(8) The high melting point of most metals makes the startup of a reactor
difficult.

(9) Polonium may represent an additional hazard. However, if the
polonium remains with the fission products, it should not add te the prob-
lems already present.

18-3. Liquip METAL FUuEL REACTOR TYPES

As a solvent for liquid-metal fuels, bismuth is a natural choice because
it dissolves uranium and has a low cross section for thermal neutrons. As a
result, research work at Brookhaven National Laboratory has centered
on bismuth-uranium fuels. Other possible liquid-metal fuels are the Los
Alamos Molten Plutonium System (LAMPRE) [14] and dispersions of
uranium oxide in liquid metals, NaK [15] or bismuth [16]. The limited
solubility of uranium in bismuth is troublesome in some designs. More
concentrated fuels can be obtained by using slurries or dispersions of solid
uranium compounds in bismuth. Among the solids which have been sug-
gested are intermetallic compounds [10] uranium oxide [16], uranium
carbide, and uranium fluoride. Use of a dispersion avoids the limited con-
centration but introduces other problems of concentration control, sta-
bility, and erosion.

Liquid metal fuel reactors would appear to be most useful for large
central station power plants [6,11,17-20] where the integrated chemical
processing, one of the attractive features of an LMFR system, would be
important.

The uranium-bismuth fuel system is flexible and can be used in many
designs. Although other types of liquid-metal systems are certainly possible,
the LMFR at Brookhaven is being designed as a thermal reactor in which
the fuel is dissolved or suspended in a liquid heavy-metal carrier. Ordi-
narily, the liquid metal is bismuth for highest neutron economy, but other
systems such as lead or lead-bismuth eutectic may be used. The moderator
is graphite, although beryllium oxide has also been considered.
18-3] LIQUID METAL FUEL REACTOR TYPES 707

   
 

F————————

D GIND D GEED GIED GEND I D GNP ERED D
(;":‘319"" ) Externally Cooled internally Cooled
ot-Type

ez
i 1

 

 
 

 

 

 

 

 

 

 

 

 

 

 

 

Slightly . U-Bi or Th-Bi or| | Solid Th
é"l -Th Enriched S UI j ?' UO:.-Bi U-Th ThO:-Bi | | Blanket
urry U-Slurry olution Slurry Slurry Slurry || Elements

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

F1c. 18-1. Classification of Liquid Metal Fuel Reactors.

Liquid metal fuel reactors are classified on the basis of their heat-
transfer characteristics (Fig. 18-1) [21]. If heat is transferred within the
core the reactor is said to be internally cooled. If heat is transported by the
fuel to the primary heat exchanger external to the core, the reactor is
externally cooled. The term ‘‘integral reactor”’ implies an externally cooled
system, but one so compact that the reactor and primary heat exchangers
can be placed in the same container.

Externally cooled LMFR’s can be divided into two classes, single-fluid
and two-fluid. In the single-fluid reactor the fissionable and fertile ma-
terials are combined in a single fluid carrier, bismuth. This type of reactor
has no separate blanket, and conversion or breeding takes place within the
core fluid itself. The conversion ratio can be made to approach unity with
the proper choice of such parameters as core size, graphite-to-fuel ratio, and
thorium concentration. However, the most economic design is not neces-
sarily the one having the highest conversion ratio (see Chapter 24). If
no fertile material is mixed with the fuel, the concept reduces to the simple
burner.

The two-fluid externally cooled LMFR (Fig. 18-2) is somewhat more
complex because it has a physically separate core and blanket, but higher
conversion ratios are possible. The blanket can be made in a variety of
ways, making use of either solid or liquid blanket materials. In exploiting
the LMFR concept to the full, a fluid blanket consisting of a slurry of
ThBiz or ThO2 in bismuth is used.

A variety of fuels is also possible. In the two-region reactor, critical
concentrations of uranium in bismuth could be below solubility limits;
708 , LIQUID METAL FUEL REACTORS [cHAP. 18

 

 
 

. / Radioactive
Salt Fuel Bi-U233.F ps Gas Storage

r+ FP's Process
Pump

S(.Zilf—J . 233
U Conch|Bi-U
Excess U233 Confrol

       
 

 

 

 

Bi-Po Steam Plant

'
233 Spuvge ; Tl’Op
l U Gas

fl{ Storage )Pump(s)
Y ’

Waste Bi-Th3Bi5-P0233.U233.Fp’s ;
Storage u233 Blanket
, Bi-Th3Bis
Pa/ Tpy233
Holdup U233 | Thg Bis-Bi

Blanket ~|$
\_ Salt Blanket
+ FP's Process Graphife z

Moderator
SalfTJ I Bi-U233 Fyel

\a“i

 

 

 

 

 

 

 

 

 

 

 

F1c. 18-2. Schematic diagram of LMFR, showing reactor, steam plant, and
chemical processing.

therefore solution fuels are possible. Such a fuel for the single-region re-
actors is possible only for small thorium loadings or for burners. Higher
fuel concentrations can be utilized only through the use of slurries. On the
basis of experiments, a maximum slurry content of 10 w/o (weight percent)
of either uranium or thorium as bismuthide compounds in bismuth can be
assumed. If an oxide slurry is used, approximately 20 w/o can be carried
by the bismuth. So far only fuels of U233 and U235 have been investigated
in the LMFR program.

18-4. LMFR ProGgraMm

In the following chapters detailed discussions of the liquid metal fuels
research, development, and engineering work are given. Practically all
the LMFR work is in the research and development stage. In the first
group of chapters, the physics, chemistry, and engineering design of the
LMFR are discussed. In the last chapters, several liquid metal fuel re-
actor designs, based on current research and development, are presented.
It should be understood that these are design studies and it is expected that
more than one liquid metal fuel experimental reactor will have to be built
and operated before a final commercial design is evolved.
REFERENCES 709

REFERENCES

1. H. HaLBaN and L. Kowarskr, Cambridge University, England, Cavendish
Laboratory, 1941. Unpublished.

2. M. E. Leg, Fairchild Engine & Airplane Corp., NEPA Division, 1950.
Unpublished.

3. E. P. WIGNER et al., Argonne National Laboratory, 1944. Unpublished.

4. G. Young, Oulline of a Liquid Metal Pile, USAEC Report MonP-264, Oak
Ridge National Laboratory, Mar. 5, 1947.

5. O. E. DwyEr, Heat Transfer in a Liquid-Metal-Fuel Reactor for Power, in
Chemical Engineering Progress Symposium Series, Vol. 50, No. 11. New York:
American Institute of Chemical Engineers, 1954. (pp. 75-91)

6. C. WiLLiams and F. T. MiLes, Liquid Metal Fuel Reactor Systems for
Power, ibid., No. 11. (pp. 244-252)

7. J. E. ATHERTON et al., Studies in the Uranium-Bismuth Fuel System, ibid.,
No. 12. (p. 23)

8. C. J. RasemaN and J. WEIsMAN, Liquid-Metal-Fuel Reactor Processing
Loops, ibid., No. 12. (p. 153)

9. D. W. Bareis et al., Processing of Liquid Bismuth Alloys by Fused Salts,
ibid., No. 12. (p. 228)

10. R. J. TEiTeEL et al., Liquid-Metal Fuels and Liquid-Metal Breeder Blan-
kets, ibid., No. 13. (p. 11)

11. NucLEAR ENGINEERING DEPARTMENT, BROOKHAVEN NATIONAL LABORA-
ToRY, Liquid Metal Fuel Reactor Systems, a collection of seven papers, Nucleonics
12(7), 11-12 (1954).

12. O. E. Dwykr et al., Liguid Bismuth As a Fuel Solvent and Heat Transport
Medium for Nuclear Reactors, paper presented at the Nuclear Engineering and
Science Congress at Cleveland, Ohio, Dec. 12-16, 1955. (Preprint 50)

13. J. CHERNICK, Small Liquid Metal Fueled Reactor Systems, Nuclear Scu.
and Eng. 1, 135-155 (1956).

14. R. M. KieuN, A Molten Plutonium Reactor Concept—LAMPRE, USAEC
Report LA-2112, Los Alamos Scientific Laboratory, January 1957: Los Alamos
Molten Plutonium Reactor Equipment (LAMPRE), Nucleonics 14(2), 14
(February 1956); Molten Plutonium Reactors, in Radialion Safety and Major
Activilies in the Atomic Energy Programs, July—December 1956, U. S. Atomic
Energy Commission. Washington, D. C.: Government Printing Office, January
1957. (p. 43)

15. B. M. ABraHaM et al.,, UO2-NaK Slurry Studies in Loops to 600°C,
Nuclear Sct. and Eng. 2, 501-512 (1951).

16. J. K. DavipsoN et al., A UOgz-Liquid Metal Slurry for Economic Power,
paper presented before the American Nuclear Society at Washington, D. C.,
Dec. 10-12, 1956.

17. F. T. MiLes and C. WiLLiams, Liquid Metal Fuel Reactor, in Proceedings
of the International Conference on the Peaceful Uses of Atomic Emergy, Vol. 3.
New York: United Nations, 1956. (P/494, p. 125)
710 LIQUID METAL FUEL REACTORS [cHAP. 18

18. D. J. SExGsTAKEN and E. Durnawm, Liquid Metal Fuel Reactor for Central
Station Power, paper presented at the Nuclear Engineering and Science Congress
at Cleveland, Ohio, Dec. 12-16, 1955. (Preprint 39)

19. Bascock & Wircox Co., Liquid Metal Fuel Reactor; Technical Feasibility
Report, USAEC Report BAW-2(Del.), June 30, 1955.

20. D. Magrs et al., Preliminary Design of an LMFR Power Plant, Nuclear
Sct. and Eng., in preparation.

21. BaBcock & WiLcox Co., 1958. Unpublished.
CHAPTER 19
REACTOR PHYSICS FOR LIQUID METAL REACTOR DESIGN*

The flexibility of liquid metal fuel systems is such that they range over
several different reactor categories. Liquid metal reactors may be designed
as fast, intermediate, or thermal systems, with either circulating or static
fuel systems. The reactor core components consist of a fuel carrier such as
molten bismuth or lead, and a moderator such as graphite or beryllium, if
the neutrons within the reactor core are to be thermalized. If the fuel is
stationary, a second fluid is required as the reactor coolant.

In the simplest system, a high-temperature liquid-metal solution or
slurry would be pumped through an externally moderated reactor core. For
such a reactor, the neutron physics problems would be similar to those of
aqueous homogeneous systems. The chief difference would lie in the neu-
tron spectrum, which would be higher because of weaker moderation and
higher operating temperatures.

The liquid-metal system that has received the greatest emphasis to date
is of the heterogeneous, circulating fuel type. This reactor, known as the
Liquid Metal Fuel Reactor (LMFR), has as its fuel a dilute solution of
enriched uranium in liquid bismuth, and graphite is used as both moderator
and reflector. With U233 as the fuel and Th232 as the fertile material, the
reactor can be designed as a thermal breeder. Consideration is restricted
here to this reactor type but, wherever possible, information of a general
nature is included.

19-1. LMFR PARAMETERS

19-1.1 Cross sections. Most of the cross sections required for neutron
physics studies of the LMFR can be obtained from BNL-325. The fol-
lowing exceptions should be noted. The 2200 m/sec value of the absorption
cross section of graphite is given as 3.2 4+ 0.2 mb. The best experimental
value however is 3.6 mb after correcting for the presence of such impurities
as B, N2, etc. Graphite of density 1.65 to 1.70 g/cm3 is obtainable with
an absorption cross section of about 4 mb, including impurities. Graphite
of density 1.8 g/cm3 or higher is becoming available, but the purity of
this high-density graphite has not been well established.

The 2200 m /sec value of the absorption cross section of Bi?%9 is 32 + 2 mb.
Two isomeric states of Bi2!0 are formed, one of which decays by 8-emission
with a half-life of 5 days into Po210.

*Contributed by J. Chernick, Brookhaven National Laboratory.
711
712 PHYSICS FOR LIQUID METAL REACTOR DESIGN [cHAP. 19

TaBLE 19-1

PARAMETERS oF B1202 RESONANCES

 

 

 

Eo(ev) agol’, barn-ev I, ev
810 9400 5.84+0.3
2370 7660 17 +1.5

 

 

 

 

 

Bismuth has prominent resonances at 810 ev and 2370 ev, largely due
to scattering. Breit-Wigner parameters obtained by Bollinger et al. at
Argonne National Laboratory are listed in Table 19-1. To determine neu-
tron capture within these resonances, it is necessary to estimate the value
of the level width, I'y. One method is to use the value of 0.5 b obtained
by Langsdorf (ANL—-4342) for the resonance integral, which implies that
I'y is about 150 mv. An analysis of Bollinger’s data indicates that a more
likely value is about 50 mv.

High-energy cross sections of bismuth and lead are of secondary interest
in well-moderated liquid-metal reactors, but would become of prime in-
terest in fast- or intermediate-energy reactors. On the basis of the known
levels and spin assignments for bismuth and lead, Oleksa of Brookhaven
National Laboratory has calculated cross sections that are in good agree-
ment with experimental data. The (n, p) and (n, @) cross sections are neg-
ligible. The threshold for the (n, 2n) cross section in bismuth is high,
7.5 Mev. At 1.0 and 4.3 Mev the transport cross sections of bismuth are
calculated as 4.3 b and 4.2 b, respectively. The capture cross section at
1 Mev is 3.4 mb.

Inelastic scattering in bismuth is important in considering fission-energy
neutrons. The results of Oleksa’s studies are presented in Table 19-2.
The lowest levels in Bi?% occur at 0.9, 1.6, 3.35 Mev, respectively. At
energies up to 2.6 Mev, Oleksa finds that the cross sections for scattering
into the individual levels are in good agreement with calculations based
on the Hauser-Feshbach model.

In a U-ueled liquid-metal system, the cross sections of the higher iso-
topes or uranium are of considerable importance in determining equilib-
rium concentrations of these isotopes and the time required to approach
their equilibrium. These equilibrium conditions require study because of
solubility limitations in a liquid-metal fuel reactor. The chain starts with
either U235 or U233, depending on whether a converter or breeder reactor
i1s under consideration, and ends with U237 because of its short half-life.
In addition, some U?3® may be present in the fuel. Thermal cross sections
are given in BNL-325.
19-1] LMFR PARAMETERS 713

Other absorption cross sections of importance to high-power, high-fuel-
burnup reactors are those of the long-lived fission products and, in a U233
breeder, that of Pa233. Despite a number of comprehensive studies of these
effects, accurate values may not be known until such reactors have been
in operation for some time. Fuel-processing studies for the LMFR, how-
ever, indicate that the poisoning effect can economically be maintained at
a few percent.

Although the LMFR is a heterogeneous reactor, the fuel and moderator
arrangements that have been proposed yield a core which is nearly homo-
geneous from the neutron physics viewpoint. The preferred core is an
impermeable graphite structure perforated with holes of about 2 in. di-
ameter for passage of the liquid-metal fuel. The moderator volume is about
equal to that of the liquid metal, bismuth, which contains about 0.1 w/o
enriched uranium. Actually, the size of the fuel channels could be consider-
ably increased without seriously increasing the flux disadvantage factor
and, hence, the critical mass of the reactor core.

19-1.2 Neutron age and diffusion length. The following formulas, ap-
propriate for mixtures, have been used to obtain the diffusion area, L?,
and neutron age, 7, of graphite-bismuth LMFR cores:

 

 

 

[2=13.%. (19-1)
S e + R) , (19-2)
(£24)Bi (Ztr)Bi
1+ (EZo |1+ e 8]

where £ is the logarithmic energy decrement,
2, 1s the macroscopic scattering cross section,
2. 1s the macroscopic absorption cross section,
2+ 1s the macroscopic transport cross section,
the subscripts Bi and C indicate the macroscopic cross section for
the respective materials, and R is the bismuth-to-graphite volume
ratio.

19-1.3 Reactivity effects. A problem unique to circulating fuel reactors
is the loss of delayed neutrons in the external circuit. Since the time spent
by the delayed-neutron emitters outside the reactor core is generally greater
than that spent within the core, a considerable fraction of the delayed neu-
trons may be wasted. In addition, since most of the delayed-neutron emit-
ters are produced as gases, they may be carried off during degassing opera-
tions. For U233, the delayed neutron fraction in thermal fission is only
0.249,. Thus prompt critical may, in some cases, be as little as 0.19,
excess reactivity.
714 PHYSICS FOR LIQUID METAL REACTOR DESIGN [cHAP. 19

TaBLE 19-2

INELASTIC SCATTERING CROSS SECTION OF BI

 

 

 

E, Mev onBi, barns
0.9 0
1.0 0.1
1.5 0.4
2.0 0.7
3.0 1.4
4.0 2.0
5.0 2.4
6.0 2.6
7.0 2.6
8.0 2.5
10.0 1.5

 

 

 

 

Coupled with this problem is the fact that the prompt temperature
coefficient (due to liquid metal expansion) in the LMFR system under
consideration is of the order of —5 X 107 5/°C. Thus, ignoring temperature
overshoots, which are discussed later, the magnitude of rapid reactivity
changes must be limited to avoid large metal temperature changes. The
total temperature coefficient of the LMFR runs about —1.5 X 1074/°C,
the delayed coefficient resulting primarily from increased neutron leakage
due to the heating of the graphite structure. While the slow response of
the graphite to power changes thus limits the size of the prompt tempera-
ture coefficient, it aids in stabilizing the system against small oscillations
at high power output.

19-1.4 Breeding. The LMFR can be operated as a breeder on the
U233-Th232 cycle. The possible breeding gain is not large, since the value
of n for U233 is about 2.3. The theoretical gain is at most 0.3, but a value
of 0.10 is about the maximum possible in a practical system. In fact,
optimization based on economic considerations would probably reduce the
gain to zero in any power breeder built in the near future. The gain is
reduced by competitive neutron capture in the core and blanket, and by
neutron leakage from the blanket and from the ends of the reactor core.
A problem not yet solved is that of a leakproof, weakly absorbing con-
tainer that will separate the core and blanket. It is hoped that beryllium
or an impermeable graphite will provide such a container for the LMFR.
Croloy steel or tantalum containers about 1/4-in thick appear satisfactory
19-2] LMFR STATICS 715

from the mechanical and metallurgical standpoint but effectively wipe out
the potential breeding gain because of their absorption cross section.

A number of studies of the so-called immoderation principle have been
carried out in an attempt to reduce the neutron losses to the container.
By removing the bulk of the moderator from a small region on both sides
of the container wall the thermal neutron losses can be greatly reduced.
Several feasible mechanical designs embodying this principle have been
worked out by the Babcock & Wilcox Company.

19-2. LMFR STaTIiCS

19-2.1 Core. The standard LMFR is predominantly thermal, nearly
homogeneous, and moderated by graphite. Thus age-diffusion theory is
applicable, and therefore the following formula can be used for a critical
system:

koe ™8 =1+ L2B2, (19-3)
where B2 is the buckling of the system, and

ko = 1, (19-4)

where 7 is the number of fast neutrons produced per thermal neutron
captured in the fuel, and f is the thermal utilization factor. The product
of the fast fission effect, €, and resonance escape probability, p, is assumed
equal to unity.

In view of the uncertainty in the value of I', for bismuth, the validity of
neglecting resonance capture is still uncertain, and Monte Carlo studies
are planned at BNL to obtain lower limits for p as a function of channel
size and lattice pitch. For small channels, the homogeneous formula for f
is adequate, since consideration of self-shielding of the fuel reduces f in a
typical core by about 2%.

Studies have yielded for buckling the typical values given in Table 19-3 [1]
for both U233 and U235 as the fuel in a graphite moderator at an average
core temperature of 475°C.

10-2.2 Reflector. In order to apply the above results to reflected re-
actors, it is necessary to determine the reflector savings, which can be
obtained from conventional two-group theory. This method could also
be used to estimate the critical size of the reactor but, for small cores, two-
group theory underestimates the size of graphite-moderated reactors.

Two-group results obtained for typical reflectors are given in Table
19—4 [1] for a cylindrical reactor system surrounded by a large reflector.
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

716 PHYSICS FOR LIQUID METAL REACTOR DESIGN [cHAP. 19
TABLE 19-3
BuckLING OoF GRAPHITE-MoODERATED LMFR CorEks
VBi/VC Ny/Ng;=0.6 X 10~3 Ny/Ngj=1X 10~3 Ny/Ng;j=15X%X 10-3
U233 Fuel
0.25 6.64 X 10~ 4cm—2 9.51 X 10~ 4cm—2 11.85 X 10~ 4cm—2
0.50 8.20 10.70 12.50
1.00 8.09 9.82 10.95
1.50 7.33 8.60 9.43
2.00 6.59 7.60 8.22
U235 Fuel
0.25 6.05 8.65 10.72
0.50 7.43 9.63 11.19
1.00 7.25 8.74 9.71
1.50 6.53 7.64 8.30
2.00 5.68 6.72 7.23
TABLE 194

REFLECTOR SAVINGS OF U235-Br CorRES MODERATED BY (GRAPHITE

 

 

 

 

 

 

 

 

 

Core
Reflector
Vei/Veo= 1 2 1 2
Reflector Ny/Npi=[0.6X1073]0.6X 1073{1.5x 1073|1x 103
Graphite 1.71 ft 2.10 ft 1.66 ft 2.10 ft
909,C-109, Void 1.62 2.00 1.63 2.01
Ve-Vai—Vrn:
85— 5- 0 1.69 2.04 1.65 2.04
8515~ 0 1.66 1.97 1.59 2.01
85-10—- 5 0.70 0.82 0.78 0.87
70-20-10 0.58 0.68 0.65 0.73

 

 

 
19-2] LMFR STATICS 717

19-2.3 Critical mass. The results of age-diffusion theory are in good
agreement with multigroup calculations for predominantly thermal LMFR
reactors. At higher fuel concentrations, however, the age theory over-
estimates the critical mass, as shown in Table 19-5 [1]. The differences in
critical mass estimates are large only for weakly moderated reactors.

"TABLE 19-5

CriTicaL. Mass AND DiaAMETER orF U235-FurrLep LMFR
SPHERES WITH A 90-cM GRAPHITE REFLECTOR

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Age—Diffusion Multigroup
Nuy/Nsi | VeiU/V¢ :
Diameter, ft | Mass, kg | Diameter, ft | Mass, kg
Graphite-moderated
1x 103 0.25 4.38 2.73 4.52 3.02
1.0 3.88 4.77 3.81 4.53
2.0 4.15 7.79 4.04 7.15
1x 10-2 0.25 2.57 5.50 2.28 3.89
1.0 ' 2.79 17.69 2.24 9.20
Beryllium-moderated
1x 103 0.25 3.86 1.87 3.88 1.92
1.0 3.08 2.37 2.94 2.07
2.0 3.29 3.86 3.04 3.05
1X 10~2 0.25 1.90 2.22 1.66 1.49
1.0 2.11 7.70 1.73 4.19
2.0 2.43 15.55 1.94 7.87

 

 

 

 

 

 

 

19-2.4 Breeding. The conversion ratio obtainable in liquid metal sys-
tems depends on a number of variables, such as the fuel and fertile material
concentrations, the fission-product processing methods, losses to the core
container, etc. In a feasibility study of the LMFR conducted by the
Babcock & Wilcox Company, currently practical reactor designs were
reported (BAW-2) with conversion ratios ranging from 0.8 to 0.9, de-
pending on whether an oxide slagging or fused salt method was used for
nonvolatile fission-product processing. The U/Bi atomic ratio was low
(0.6 X 10~3) and a 23% Cr-1% Mo steel core container was used, both
718 PHYSICS FOR LIQUID METAL REACTOR DESIGN ~ [cHAP. 19

choices tending to reduce the possible breeding ratio. The estimates of
the neutron balance are given in Table 19-6 [6].

TABLE 19-6

NEUTRON BALANCE oF Tu232, U233 BREEDER

 

 

 

 

 

Scheme A Scheme B
Oxide slagging Fused-salt process

Production per U233 absorption 2.31 2.31
Losses: Absorption in

U233 1.00 1.00

Bi 0.13 0.13

C 0.05 0.05

Fission products 0.12 0.03

Higher isotopes 0.02 0.02

Croloy structure 0.12 0.12

Th 0.80 0.89

Pa 0.02 0.02

Leakage 0.05 0.05

 

 

 

19-2.5 Control. Because of its prompt temperature coefficient, the
LMFR is expected to be stable. Nevertheless, it represents a completely
new and untested system. There are a number of ways in which the
reactivity of the system can change, for example, with changes in inlet
temperature, concentration, or velocity of the fuel, and changes in xenon
concentration, delayed neutron emitter concentration, and blanket com-
position. Most of these changes are expected to be gradual, but they can
be sufficiently large to require the use of control rods. Inherent stability
has not been demonstrated in operating reactors except over a limited
range in reactivity and power output. In a reactor with a high-velocity
coolant there may occur sudden changes of reactivity which are too fast
for conventional control. Thus both inherent stability against sudden
reactivity changes and control rods for large but gradual reactivity changes
are needed until considerable experience has been gained in operation of
the reactor.

Studies have been carried out at BNL on control requirements for an
LMFR experiment. The control requirements depend not only on the
choice of operating temperatures, the passible xenon and fission-product
poisoning, etc., but also on conceivable emergency situations such as errors
in fuel concentration control. In a reactor with a full breeding blanket, the
control requirements may have to include the effect of complete loss of
the breeder fluid.
19-3] LMFR XINETICS 719

For a 5-mw experiment, control of 15% reactivity appears to be ample
and can be obtained with four 23% Cr-1% Mo steel rods of about 2-in.
diameter. Blacker rods containing boron could, of course, be used to in-
crease reactivity control. A study of various arrangements of identical
rods in a ring around a central rod indicates that the optimum position of
the ring occurs at about 1/4 of the distance from the reactor center to the
(extrapolated) radius of the reactor core.

It would be highly desirable to use sheaths for control rods in order to
eliminate the problem of rod insertion through a heavy liquid metal. Steel
sheaths are not satisfactory, since they reduce the breeding ratio in a
liquid-metal power breeder and reduce the over-all thermal flux in an
experimental reactor. The solution to the problem may lie in the develop-
ment of structurally sound beryllium sheaths.

19-2.6 Shielding. Shielding of an LMFR is compli¢ated by the necessity
of shielding an external circuit in which the delayed neutron emitters and
fission products decay.

Calculations by K. Spinney at BNL indicate that even for a 5-Mw
experimental reactor, about 5.5 ft of concrete are required as a neutron
shield around the reactor cell. Gamma shielding of the cell requires about
8.5 ft of ordinary concrete or 4.5 ft of BNL concrete (70% Fe). For this
reason, it has been proposed that heavy concrete be used as the shield for
the 5-Mw reactor. For the rest of the circuit, including the degasser,
pumps, héat exchanger, etc., the advantage of using BNL concrete is
less evident.

19-3. LMFR KiNETICS

A number of fundamental studies of the kinetics of circulating fuel
reactors have been carried out at ORNL and by Babcock & Wilcox Com-
pany. A review of the subject has been given by Welton [2]. At low power,
the equations governing the system are linear and complicated chiefly by
the feedback of delayed neutrons. General results for the in-hour relation
have been obtained by Fleck [3] for U233- and U235-fueled reactors. At
high power, the kinetics are much more complicated and there is a real
question whether the response of a complex reactor can be accurately pre-
dicted in advance of its operation. Bethe [6] has strongly recommended
the use of oscillator experiments to determine reactor transfer functions.
Despite such experiments, however, the mechanism responsible for the
resonances observed in EBR-I has, to date, not been satisfactorily ex-
plained.

There are two methods of treating the kinetics of a reactor. In the open-
loop method, the inlet temperature is taken as constant. The justification
for this procedure is that this condition generally prevails during rapid
720 PHYSICS FOR LIQUID METAL REACTOR DESIGN [cHAP. 19

transients, the feedback of information through the external system being
slow by comparison. The method, however, suffers from the defect that it
cannot reveal instabilities associated with the entire circuit. In the closed-
loop method, the external system, or a reasonable facsimile, is coupled to
the reactor system. The representation of the reactor, however, is generally
oversimplified because of the complexity of the over-all system.

Although the set of kinetic equations that include temperature effects
are nonlinear, the linearized equations are satisfactory for the investigation
of stability and the qualitative transient behavior. A large subset of equa-
tions is required to properly treat the effect of the delayed neutron emitters.
Again, however, lumping the delayed neutrons into a single group, or
neglecting them altogether, always appears to lead to qualitatively, if not
quantitatively, correct results.

A study of the temperature-dependent open-loop kinetics of the LMFR
has been carried out by Fleck [4]. The effect of delayed neutrons and the
delayed moderator temperature coefficient were neglected. Under these
conditions, Fleck found that the reactor responded rapidly and with little
overshoot in temperature when subjected to the largest permissible re-
activity excursions.

Using a method developed at the Oak Ridge National Laboratory
(ORNL-CF1-56-4-183) for homogeneous systems, the Babcock & Wilcox
Company has studied the stability of the LMFR against small oscillations.
The results show that the LMFR models under study are stable up to
power densities 100 to 1000 times greater than the nominal design level.

Fleck has also examined the transient pressures in LMFR cores by treat-
ing the bismuth as a frictionless, compressible fluid. He found that the
maximum pressures developed during conceivable transients were quite
small. The assumption sometimes made, that the fluid external to the core
can be represented as an incompressible slug, was found to overestimate
the transient pressures.

In general, heterogeneous reactors possessing both a small prompt
(positive or negative) fuel temperature coefficient and a large delayed
negative moderator temperature coefficient can be expected to exhibit
oscillatory instability at sufficiently high power. However, elementary
models indicate that power levels high enough to cause such instability
are not achievable in present reactors. Further study of the complex heat-
transfer transients in reactor systems is still required before reactor sta-
bility can be assured.
REFERENCES 721

REFERENCES

1. J. CHERNICK, Small Liquid Metal Fueled Reactor Systems, Nuclear Scs.
and Eng. 1(2), 135 (1956).

2. T. A. WErroN, Kinetics of Stationary Reactor Systems, in Proceedings of
the International Conference on the Peaceful Uses of Atomic Energy, Vol. 5. New
York: United Nations, 1956. (P/610, p. 377)

3. J. A. FLECK, JR., Theory of Low Power Kinetics of Circulating Fuel Reactors
with Several Groups of Delayed Neutrons, USAEC Report BNL-334, Brookhaven
National Laboratory, April 1955.

4. J. A. FLECK, JR., The Temperature Dependent Kinetics of Circulating Fuel
Reactors, USAEC Report BNL-357, Brookhaven National Laboratory, July 1955.

5. G.T. TramMELL, Oak Ridge National Laboratory, 1955. ORNL-1893,1955.

6. Bacock & WiLcox Co., Liquid Metal Fuel Reactor; Technical Feastbility Re-
port, USAEC Report BAW-2(Del.), June 30, 1955.
CHAPTER 20
COMPOSITION AND PROPERTIES OF LIQUID-METAL FUELS*

20-1. Core FuerL CoMPOSITION

In Chapter 18, the advantages and disadvantages of liquid metal fuels
were discussed in a general way. The point was made that a liquid-metal
fuel has no theoretical limitation of burnup, suffers no radiation damage,
and is easily handled for fission-product poison removal. In this chapter,
the results of research and development on various liquid-metal fuels are
presented. This work has been largely concentrated on uranium dissolved
in bismuth.

At the contemplated operating temperatures of approximately 500°C,
it was found that uranium has adequate solubility in bismuth when present
by itself. However, as the work progressed, it soon became evident that
other materials would have to be added to the solution in order to obtain a
usable fuel. The present fuel system contains uranium as the fuel, zir-
conium as a corrosion inhibitor, and magnesium as an oxygen getter.

An LMFR operating on the contemplated Th?32 to U233 breeding cycle
can be designed with an initial U233 concentration of 700 to 1000 ppm in
bismuth. The actual figure, of course, is dependent upon the specific de-
sign and materials used. In Chapter 24, in the design studies, such figures
are given. The concentrations of zirconium and magnesium are each ap-
proximately 300 ppm. It is contemplated that these concentrations will
have to be varied depending upon desired operating conditions. In their
use as corrosion inhibitor and antioxidant there is enough leeway for this
purpose. |

The fuel described in the previous paragraph is the clean fuel which would
be charged initially. During reactor operation, however, fission products
will build up in the fuel and would be maintained at a level dictated by the
economics of the chemical reprocessing system used. It has been found
that the fission products and other additives to the bismuth have an im-
portant effect on the solubility of uranium in bismuth. These have been
carefully investigated in order to permit selection of reactor temperatures
that will ensure that all the uranium remains in solution during reactor
operation. Likewise, the solubility of steel corrosion products has been in-
vestigated to determine their effect on uranium solubility in bismuth.

*Based on contributions by D. H. Gurinsky, D. G. Schweitzer, J. R. Weeks,
J. S. Bryner, M. B. Brodsky, C. J. Klamut, J. G. Y. Chow, R. A. Meyer, R. Bour-
deau, and O. F. Kammerer, Brookhaven National Laboratory.

722
20~-2] SOLUBILITIES IN BISMUTH 723

635° C 560 496 441 394 352 315
5.0 | | | | | | J

 

 
   

¥
1

  

 

 

 

1.0 1.1 1.2 1.3 1.4 1.5 1.6 1.7
1000/T °K

Fia. 20-1. Solubility of uranium in bismuth.

It is important to note that although the basic fuel is a simple one, the
uranium used for liquid metal fuel reactors using the Th—-U233 cycle must
be almost completely enriched 233 or 235 in the initial charge. Further,
since the concentrations are measured in parts per million by weight, it is
not an easy matter to maintain a strict accounting of all fuel. When deal-
ing with such small amounts, losses due to reaction of uranium with
carbon and adsorption of uranium on steel and graphite walls can be sig-
nificant.

The fuel for the LMFR is still under extensive study. At present, most
of the major information for the design of an LMFR experiment is at hand.
This information is primarily solubility data and other fuel information,
presented in the following pages.

20-2. SOLUBILITIES IN BismMuTH

20-2.1 Uranium. The experimental techniques used to measure solu-
bilities in liquid bismuth have been described previously [1,2]. Several
workers [3-7] have investigated the solubility of uranium and bismuth.
Recently, with improvements in analytical techniques, redetermination of
the solubility curve has been undertaken. The latest results are at variance
with the older work of Bareis [5], as shown in Fig. 20-1. It can be seen
that the recent data obtained at Brookhaven National Laboratory are, at
some temperatures, as much as 20 to 259, lower than those obtained some
years ago.

This variance in solubility determinations may be due to several factors,
but it is believed that the improved techniques are more reliable, and that
the newer values are consequently more precise. The presence of such
724 PROPERTIES OF LIQUID-METAL FUELS [cHAP. 20

Temperature, °C

 

 

 

 

 

1200 1000 800 700 600 500 400 300
eyttt | ' |
10.0
o 1:00
3
>
3
2
<
v
100
010
6 8 10 12 14 16 18
1/T °K x 104

F1G. 20-2. Solubility of uranium and thorium in bismuth.

727 636 560 494 441 394 352 315 283

 

     
 

 

 

 

10% T T T T T T 1

O
E ].Oo/o - pu—
3

=

5 Los Alamos

lo- 0.1 °/o e ot
a .
C01% _»l : | | | ] 1 | | ‘L

092 10 1.1 1.2 1.3 . 1.4 1.5 1.6 1.7 1.8 1.9

1000/T °K
Fig. 20-3. Solubility of plutonium and uranium in bismuth.
20-2] SOLUBILITIES IN BISMUTH 725

other materials as nickel, copper, manganese, etc., in the bismuth in quan-
tities large enough to affect the uranium solubility still remains to be in-
vestigated. IFor example, nickel has been shown to markedly reduce the
uranium solubility in bismuth [1].

It 1s obvious that even slight variations of the solubility of uranium in
bismuth might be of considerable importance in LMFR reactor design.
The solubility of uranium, according to the preferred data (the solid curve
in Fig. 20-1), allows a rather small leeway in uranium conecentration in the
reactor cycle when the lowest temperature of 400°C in the heat exchangers
is taken into account.

20-2.2 Thorium and plutonium. The solubility of thorium in bismuth,
as determined by Bryner, is compared with the solubility of uranium
in Fig. 20-2. In the temperature range 400 to 500°C, the solubility of
thorium is markedly lower than that of uranium. In fact, it is so low that
a breeding cycle using only thorium in solution with bismuth cannot be
carried out.

To fill out the information on fissionable fuel solubility in bismuth,
Fig. 20-3 shows the solubility of plutonium in bismuth, as determined at
the Los Alamos National Laboratory. In comparing plutonium with
uranium, it is seen that plutonium is significantly more soluble.

20-2.3 Fission-product solubility. The solubilities of most of the impor-
tant fission products have been determined, and are shown in Fig. 20-4.
In general, all the fission products are soluble enough so that they will stay
in solution throughout the reactor cycle. This is not true of molybdenum
however. Attempts at determining the solubility of Mo have indicated
that it is less than 1 ppm (the limit of detection) at temperatures below
800°C. Since a fair amount of the Mo is produced by fission, this means
that a sludge might form during reactor operation. (Beryllium presents
similar difficulties, since at temperatures below 800°C the solubility of
Be has been shown to be less than 10 ppm.)

20-2.4 Magnesium and zirconium. - The solubility of magnesium in
bismuth in the temperature range 400 to 500°C is approximately 5 wt.9,
which is considerably higher than the amounts of magnesium being con-
sidered in this work (300 ppm). Little work has been done on this partic-
ular determination at Brookhaven.

The solubility of zirconium in bismuth has been determined and is shown
in Fig. 20-5. This information is important in showing that the saturation
solubility of zirconium is very close to the amounts desired for corrosion
inhibition in the temperature range 400 to 500°C.
726 PROPERTIES OF LIQUID-METAL FUELS [cHAP. 20

T, °C
560 495 440 394 352 315
100 T 1 | T T T

N
o
w

 

L L 1 i1l

 
  

llllllll

      
 

llllll|

Zr

Atomic %

| |||H||

b

o
N
|
=
c
/

o
o
O fary
T ||H|||
L L1l

 

 

 

1.1 1.2 13 1.4 1.5 1.6 1.7 1.8
1000/T, °K

F1c. 20-4. Solubility of fission products in bismuth.

20-2.5 Solubility of corrosion products in bismuth. An alloy steel is
contemplated as the tube material for containing the circulating fuel in
the LMFR. Hence it has been pertinent to determine the solubility of alloy
steel constituents in bismuth. Figure 20-6 shows the solubilities of iron,
chromium, nickel, and manganese, all of whose solubilities are fdirly high
from a corrosion point of view. Nickel and manganese are particularly
high.

The solubility of titanium is shown in Fig. 20-7. It has been shown [8]
that titanium will reduce the mass-transfer corrosion of steels by liquid
bismuth.

20-2.6 Solubilities of combination of elements in bismuth. The effect
of Zr on the U solubility. The mutual solubilities of uranium and zirconium
in bismuth have been measured over the temperature range 325 to 700°C.
The data are plotted in Fig. 20-5. When bismuth is saturated with zir-
conium, the uranium solubility is appreciably decreased. On the other
20-2] SOLUBILITIES IN BISMUTH 727

636 560 496 441 394 352°C

 

   
  
 

(Bareis) Uin Bi _

(D.G.S. & JRW) UinBi -

 

U in Bi saturated with Zr

Uor Zr in Bi, Wt %

 

 

 

1.0 1.1 1.2 1.3 1.4 1.5 1.6 1.7
1000/7, °K

F1a. 20-5. Mutual solubility of uranium and zirconium in bismuth.

hand, only a slight decrease is noted in the Zr solubility. The addition of
1000 ppm magnesium had no effect on either the uranium or zirconium
solubility. This, of course, is in considerable excess of the quantity of mag-
nesium contemplated for use in the fuel.

The mutual solubility effects were further studied by determining the
ternary system U-Zr-Bi at three temperatures, 375, 400, and 425°C.
These are shown in Fig. 20-8.

T'he effect of fission products on the solubility of U-Bi. Considerable work
has been done on determining the mutual solubility effect, of fission prod-
ucts on uranium and bismuth. A good typical example is shown in Fig.
20-9, which shows that the solubility of uranium and bismuth is affected
by 250 ppm Zr, 350 Mg, 60 Nd, 15 Sm, 15 Sr, 10 Cs, and 8 Ru. There is
little doubt that this small amount of fission products, 120 ppm, has a
small but definite effect on uranium solubility.

Effects of additives on solubility of corrosion products in liquid bismuth.
The ordinary concentrations of zirconium (250 to 300 ppm) do not affect,
the equilibrium iron solubility at temperatures from 500 to 700°C. For
728 PROPERTIES OF LIQUID-METAL FUELS [cHAP. 20

Temperature , °C
727 636 560 496 44 394 352 315

 

   
 
    

 

 

 

 

 

 

 

 

100 I I I I 1000
- Mn Bi stable ] .
- (B Ni - Bi) -
o Mn in Bi i
=
2 10 — 100 £
$ - ] G
C ~ o
. - —{ 50
@ _ _ &
£ S
- — — 30 S
2 i _ E
s 5
z o
0.1 i— —4 10 <«
= = O
- 3 @
- ~ -
N
. ~o i
p \ e
09 1.0 1.1 1.2 1.3 1.4 1.5 1.6 1.7 1.8
1000/T °K
F1a. 20-6. Solubility of Fe, Cr, Ni, and Mn in bismuth.
T, °C
727 637 561 497 442 395 353 317
> 1 1 1 T ]
03 }—
0.10 |—
5 [
$ 007
= |
0.05 —
0.03 |—
10 1.1 12 13 14 15 16 17

1000 /T, °K
F1g. 20-7. Solubility of titanium in bismuth.
20-2]

SOLUBILITIES IN BISMUTH

729

 

 

 

  
 

 

 

340 l 1 T | | I I [
Saturation Zr in Bi {425°C)
500 = o —
460 \\ —~—
420 |~ \ —
Saturation Zr in Bi (400°C) \
380 b= \ _
340 = e ———— \ ]
E 300 |- \ —
Q.
% ko Saturation Zr in Bi (375°C) \ _
N e —— \(425°C)
220 \ —
180 “(400°C) \ |
140 - \\ _
\ .
- 1 \\
20 ] 1 I L 2 ] 1 L1
700 900 1100 1300 1500 1700 1900 2100 2300 2500

F1a. 20-8. The U-Zr-Bi ternary system:

Ul Wf %

U, ppm

liquidus curves at 375, 400, and 425°C.

 

   
 

 

 

 

636°C 560 496 441 394 352
3.0
N |
N
1.0 — —]
0.6 |— . _
. UinBi 4 250 Zr ppm —
0.3 |— UinBi- 250 Zr ppm
+ 350 Mg ppm —+
120 ppm Mixed
Fission Products
0. | 1 |
1.0 1.2 1.4 1.6

1000/T °K

FIG.‘/ 20-9. Solubility of U in Bi+ 250 ppm Zr, and in Bi+ 250 ppm Zr + 350
ppm Mg + 120 ppm mixed fission products. Original alloys 3.9% U and 3% U,
respectively.
730 PROPERTIES OF LIQUID-METAL FUELS [cHAP. 20

 

   
  
 
    

 

 

 

T, °C
5 727 637 561 497 442 395 353 317
10 1 T T T T 3
C \\ ]
- Product of Normal Fe and i
104 L \‘\/Normul Cr Solubility i
N Apparent Solubility Product -
t X of Fe and Cr in Bi A
a
S 103 E 3
E N 2
& I \\ 4 300
g INON |
1102 =N\ — 100
5 N 370
¥ " \ 4 E
- 43 §
" i S
10! - 10 &
- 37 &
N r Normal ] 3
- Normal Fe Solubility \Cr Solubility |
1.0 1.1 1.2 1.3 1.4 1.5 1.6 1.7

1000/T. °K

Fia. 20-10. Effect of Cr on the solubility of Fe in Bi.

higher concentrations (above 700 ppm zirconium), the iron solubility is
increased in this same temperature range.

Zirconium in all concentrations up to saturation does not affect the sol-
ubility of chromium in bismuth.

Uranium, with magnesium additions up to 2000 ppm, does not affect
the solubility of iron in bismuth. The possible effects on chromium solu-
bility are not known at this time.

Chromium has a marked effect on the solubility of iron, whereas the
chromium solubility itself is not affected. An apparent solubility product
is observed as is shown in Fig. 20-10 by the line titled ‘‘Apparent solu-
bility product. Below 450°C, the iron solubility appears to be increased by
saturating the solution with chromium. Above that temperature, the iron
solubility is markedly reduced by chromium.

Titanium, at concentrations greater than 100 ppm, has been found to
reduce the iron solubility in the temperature range 475 to 685°C [9].

20-2.7 Salts. In some of the contemplated chemical fuel processing
methods the liquid bismuth fuel will be brought in contact with chloride
20-5] FUEL STABILITY 731

and fluoride salts. A typical chloride salt is the eutectic mixture of
NaCl-KCl-MgCls,. It is important that none of the salts dissolve in the
bismuth and get carried over into the core, since chlorine is a neutron
poison. Preliminary investigations at BNL indicate that the solubility of
these chloride salts is less than the detectable amount, 1 ppm.

20-3. PaYSICAL PROPERTIES OF SOLUTIONS

20-3.1 Bismuth properties. The physical properties of bismuth are
listed in Table 23-1.

20-3.2 Solution properties. Little work has been done on determining
physical properties of the solutions. The available results indicate that
the small amount of dissolved material does not appreciably affect the
physical properties of density, viscosity, heat capacity, and vapor pressure.
For design purposes, the properties of pure bismuth can probably be used
with safety.

20-3.3 Gas solubilities in bismuth. The question of the solubility of the
fission-product gases xenon and krypton in bismuth is of extreme im-
portance. In particular, Xe!35, a strong neutron poison, must be removed
from the system as fast as it is formed in order to have a good neutron
economy.

Attempts at measuring and calculating the solubility of these gases in
bismuth have proved extremely difficult, because of the extremely small
solubilities. Mitra and Bonilla [10] have measured the solubility of xenon
in bismuth at 492°C as 8 X 10~7 atom fraction at atmospheric pressure.
On the other hand, McMillan [11] has calculated the solubility as 10712
atom fraction at 300°C. It is probable that the amount of gases produced
in the reactor lies between these two determinations. At present, the ques-
tion of xenon and krypton solubility in bismuth is open to more intensive
research.

20-4. FueEL PREPARATION

Fuel has been prepared at BNL by simply dissolving the solid uranium,
magnesium, and zirconium in molten bismuth. The solids are usually in
the form of small chips and are placed in a small metal basket which is
then suspended in the bismuth.

20-5. FuUEL STABILITY

It is essential to maintain a homogeneous fuel and to prevent the uranium
from concentrating in any particular region of the reactor. Stability tests
have been conducted to determine conditions necessary for keeping the
732 PROPERTIES OF LIQUID-METAL FUELS [cHAP. 20

uranium in solution by preventing its reaction with the steel and graphite
of the system. Measurements have also been made of the rate of oxidation
of uranium in the liquid fuel stream. This study indicates the effect of an
accidental air leak during the reactor operation.

20-5.1 Losses of uranium from bismuth by reaction with container ma-
terials. Early attempts to make up uranium-bismuth solutions resulted
in about a 50% loss of uranium even though very high-purity bismuth
(99.99%) was used. Apparently the uranium reacted with the few im-
purities in bismuth or adsorbed on the walls of steel containers. Sand-
blasting and acid-pickling of the container walls, deoxidizing the bismuth
by hydrogen firing, and adding 250 ppm Zr and 350 ppm Mg before intro-
duction of U reduced this loss to about 5%. It is possible that even this
5% loss may not be real, but is attributable to analytical and sampling
techniques.

Only small decreases in the zirconium and magnesium concentrations
have been observed, and in tests where titanium was used as an oxygen
scavenger, no loss of U was observed.

When the fuel solution is brought in contact with graphite, usually 10 to
15% of the uranium is lost from the solution. Apparently it reacts with the
graphite or impurities present in the graphite. Research on this is under
way at present. However, it is proving to be extremely difficult since the
amounts of materials involved are so small.

Since zirconium reacts with graphite to form zirconium carbide in
preference to uranium forming uranium carbide, addition of zirconium to
the solution should help prevent loss of uranium. This effect has been
observed.

Generally it has been found that zirconium concentration will initially
drop and then maintain a constant level throughout the exposure of the
fuel solution to graphite.

20-5.2 Reaction of fuel solution with air. Should an air leak occur in
the LMFR, the uranium, magnesium, and zirconium will all tend to oxi-
dize in preference to the bismuth. Figure 20-11 shows the results of an
experiment in which air was bubbled through fuel kept at a temperature
of 405°C. These results indicate that the preference of oxidation is in the
order magnesium, uranium, zirconium.

The reaction data indicate that the uranium oxidation rate is one-half
order dependent on the UO2 present. The magnesium oxidation rate, in
general, is first order with respect to magnesium concentration. Other
experiments show that if additional amounts of magnesium are added to
the solution after the oxidation, most of the UO2 can be reduced back to
uranium. These data are given in Table 20-1.
20-5] FUEL STABILITY 733

 

100
IZirconium I I

    

O
o

I

~N
o

O
o

50 —
Magnesium

40 —

Percent of Original U, Mg, or Zr Remaining in Solution

W
(=
I
I

 

 

20 I | I I I
0 4 8 12 16 20 24
Time in Minutes _
Fie. 20-11. Concurrent oxidation of U, Zr, and Mg from Bi containing 750 ppm
U, 284 ppm Mg, and 280 ppm Zr.

 

TaAaBLE 20-1

RepuctioN oF UO, BY Ma 1IN Br

 

 

 

 

 

 

 

 

 

U (ppm)
U (ppm) :
as UOq presegt n Time after .U (pp n'1) *
7°C | before ad- solution | Mg (ppm) Mg addi- in solution | 9, UO2
e before ad- added : after Mg | reduced
dition of . tion
Mg dition of added
Mg
405 960 10 6600 25 min 710 75-100
400 150 530 5000 48 hr 660 90
360 510 310 2500 10 hr 460 30
360 550 10 1000 48 hr 290 50

 

 

 

 

*The values listed as 9, UO2 reduced are probably lower than equilibrium
values, since the samples were taken at arbitrary times after the Mg was added.
734 PROPERTIES OF LIQUID-METAL FUELS [cHAP. 20

Work on fuel stability is obviously of great importance, and is being
continued. Very little has been done so far on observation of stability
under neutron bombardment. A program is getting underway for the study
of radiation effects on the fuel concurrently with a study of corrosion
effects. For this purpose the Brookhaven Pile will be used together with
Radiation Effects Loop No. 1.

20-6. THoRIUM BIiSMUTHIDE BLANKET SLURRY

20-6.1 Status of development. In developing a blanket system for the
LMPFR, it has seemed logical to select one which is as similar as possible to
the core fuel. After considerable evaluation the principal emphasis has
been placed upon a bismuth fluid containing thorium bismuthide in the
form of very small particles. This is commonly called the thorium bis-
muthide slurry system.

Since this fluid has practically the same physical properties as that of
the core, it would be possible to balance pressures across the graphite wall
separating the blanket from the core and, in the event of mixing the core
and blanket fluids, no violent reactions would ensue. Furthermore, from a
chemical processing point of view, an all-metallic blanket system offers
considerable advantage when pyrometallurgical processing techniques
are used.

This does not mean that other types of blankets are not being studied.
Work is concurrently under way on thorium oxide-bismuth slurries. Also,
thorium carbide, thorium fluoride, and thorium sulphide slurries are under
consideration.

At the Ames Laboratory (Iowa State College) the solution of thorium
in magnesium has received considerable attention in the past few years.
This is a true solution, and certainly offers another possibility for a blanket
fluid. However, unless an absolute method for keeping the magnesium
solution separate from the core bismuth solution is found, this system
would be hazardous when used with the contemplated uranium-bismuth
core fluid, since magnesium and bismuth will react violently and cause
a marked temperature rise.

20-6.2 Chemical composition of thorium bismuthide. The thorium
bismuthide intermetallic compound discussed in this section has the chemi-
cal formula ThBi2. This compound is 35.7 w/o thorium. A second com-
pound, ThsBi4, also can exist and has been observed in alloys containing
greater than 50 to 55 w/o thorium.

20-6.3 Crystal chemistry of thorium bismuthide. ThBis has a tetragonal
crystal structure (with ap = 4.942 A, and ¢o = 9.559 A) containing two tho-
rium atoms and two bismuth atoms per unit cell. The density as determined
20-6] THORIUM BISMUTHIDE BLANKET SLURRY 735

by x-ray measurement, is 11.50 g/cc at 25°C. It is estimated to be approxi-
mately 11.4 g/cc at 550°C.

Th3Bis has a body-centered cubic structure (ap = 9.559 A) containing
12 thorium atoms and 16 bismuth atoms in the unit cell. The density is
11.65 g/cec.

Ordinarily, when thorium bismuthide is prepared at 500°C, very small
equiaxed particles (less than 0.5 micron) are formed. These equiaxed
particles grow until they reach the average size of 50 to 60 microns, and
under certain conditions they can grow to considerably larger dimensions.

When a 5 to 10 w/o thorium bismuthide slurry is cooled from a tempera-
ture of complete solution (above 1000°C), ThBis precipitates in the
form of platelets having diameter-to-thickness ratios greater than 50:1.
The plane of the platelet is parallel to the 001 plane of the crystal. Platelet
diameters up to 1 ecm have been observed in alloys cooled at moderate
- rates. The diameters can be decreased by increasing the cooling rate.

Whereas equiaxed particles tend to grow equally in all three dimensions,
it has been found that the platelets, when heated isothermally at tempera-
tures above 300°C, tend to grow faster in thickness than in diameter. The
solid particles thus tend to approach an equiaxed shape. The rate of ap-
proach to equiaxiality increases as the temperature of isothermic treat-
ment is increased.

Considerable work has been carried out on control of crystal structure
and size. The addition of tolerable amounts of Li, Be, Mg, Al, Si, Ca, Ti,
Cr, Mn, Fe, Co, Ni, Cu, Zn, Zr, Mo, Pd, Ag, Sn, Sb, Te, Pa, La, Ce, Tr,
Nd, Ta, W, Pt, Pb, and U has little effect on the mode of thorium bis-
muthide when it is precipitated. It has been found, however, that tellurium
inhibits the thorium bismuthide particle growth, agglomeration, and de-
position during thermal cycling. The platelet mode of bismuthide precipi-
tation is not modified by addition of tellurium. The amount of tellurium
used in these experiments has been 0.10 w/o.

The mechanisms by which tellurium additions inhibit ThBig particle
growth, agglomeration, and deposition are as yet uncertain. Although
- additions of tellurium in larger concentrations decrease the solubility of
thorium in bismuth markedly, the concentration of tellurium required for
inhibition decreases the solubility only slightly. These small amounts of
tellurium appear to be associated with the solid phase rather than the
liquid phase. They do not appear to alter the crystal structure.

It has been observed that under certain conditions ThBis particles
suspended in liquid bismuth can be pressure-welded to one another and to
container materials by the forces of impact. This pressure-welding phe-
nomenon has been observed at 525°C and higher temperatures. Since this
phenomenon might cause plugging by agglomeration at points of high
impact, it will be necessary to take this factor into account in the design
of slurry circulation systems.
736 PROPERTIES OF LIQUID-METAL FUELS - [cHAP. 20

20-6.4 Thorium-bismuth slurry preparation. Dispersions of small equi-
axed particles of ThBiz in bismuth can be prepared by heating finely
divided thorium, in the form of powder or chips, in contact with liquid bis-
muth at 500 to 600°C under an inert atmosphere. The intermetallic com-
pound is formed by an exothermic reaction at the thorium-bismuth
interface, when the convex radius of curvature of the thorium surface is
suitably small. The compound exfoliates into the liquid as agglomerates of
very small particles (less than 0.5 micron). These small particles grow
very rapidly, the larger at the expense of the smaller, as equiaxed single
crystals of ThBis. Rapid growth ceases when the maximum crystal di-
mensions approach approximately 50 to 60 microns. The time necessary
for complete reaction varies with the dimensions of the thorium. For
example, 325-mesh thorium powder reacts completely in 5 min at 500°C,
thorium chips 1/2” X 1/8” X 0.010” require 2 hr at 500°C, and thorium
chips 3/4” X 3/16” X 0.020"" require 13 hr at 500°C. The thorium dimen-
sions have only a slight effect upon the ultimate particle size. The reaction
can be accelerated by raising the temperature. Higher temperatures,
however, increase both the particle size and the tendency to form sintered
agglomerates rather than single crystals.

If thorium powder is added to the liquid bismuth surface at the reaction
temperature, it is necessary to stir the thorium into the liquid. Otherwise
a crust of intermetallic compound forms on the surface which is rigid enough
to support subsequent additions, thus preventing contact between the
thorium and the bismuth. ‘

During the reaction, evolution of an unidentified gas (possibly hydrogen
from thorium hydride) has been observed. It is necessary to stir the slurry
under vacuum to remove the undesirable trapped bubbles of this gas.

A photomicrograph of a typical slurry produced by the exfoliation
method is shown in Fig. 20-12. The dark ThBi2 particles appear 1n a
white matrix of solidified bismuth. The method has been used to prepare
90-1b batches of slurry and may readily be adapted to tonnage-scale prepa-
ration. The method is suitable for preparation of the initial blanket charge,
but would probably not be used for slurry reconstitution during subsequent
blanket processing. |

A modification of this method has been studied in which finely divided
thorium from a supernatant mixture of fused chlorides is electrolytically
deposited on a molten bismuth cathode at the desired temperature [13].
The thorium must be stirred through the interface. Slurries that are satis-
factory with respect to thorium content and particle size and shape have
been produced by the electrolytic method in batches of up to 10 Ib. No
evolution of gas has been detected during the thorium-bismuth reaction.
Unfortunately, the necessary stirring introduces chloride inclusions which
are difficult to remove completely. Since these inclusions would decrease
20-6] THORIUM BISMUTHIDE BLANKET SLURRY 737

     

By B

Fia. 20-12. 5 w/o Th-95 w/o0 Bi. Dispersion of equiaxed ThBiz particles in Bi
Produced by heating Th chips in Bi at 500°C for 2 hr. (150X)

the efficiency of neutron utilization in a breeder blanket because of the
high cross section of chlorine, the electrolytic method of slurry preparation
must, at present, be considered unsatisfactory.

Another preparation method for thorium-bismuth slurry is by quenching
and heat treatment. In this method a solution of thorium, for example
5 w/o, is very rapidly cooled from about 1000°C down to about 600°C.
This can be accomplished by pouring a hot solution into a container having
a sufficiently high heat capacity or by pouring the hot solution into an
equal volume of liquid bismuth heated just above the melting point. When
this is done tiny platelets are formed.

As will be discussed in the following section, the platelet form of crystal
is unsatisfactory from a fluidity point of view. When these fine platelets
are heat-treated for 20 min at 800°C, or for 5 min at 900°C, dispersions of
ThBis particles having maximum dimensions less than 100 microns and
diameter-to-thickness ratios equal to or less than 5 to 1 are produced.
Platelet formation during cooling is avoided by agitating the slurry to
suspend the particles.

Figure 20-13 shows the fine platelets produced by the quenching and
Fig. 20-14 shows the larger particles produced from these fine platelets
by the heat treatment at 800°C for 20 min. Such a slurry exhibits high
fluidity after concentration to 10 w/o thorium by removal of excess liquid
phase, and is suitable for use in the reactor blanket.

Other possible ways for reconstituting a satisfactory slurry after heating
to complete solution involve the use of ultrasonic energy [14]. It has been
demonstrated that application of ultrasonic energy to a thorium-bismuth
solution during cooling causes the formation of essentially equiaxed particles
rather than platelets. It has also been demonstrated that application of
738 PROPERTIES OF LIQUID-METAL FUELS [cHAP. 20

 

     

27 W
& €

Fia. 20-13. 5 w/o Th-95 w/o Bi. Dispersion of ThBi; platelets in Bi. Alloy
heated to 1000°C and quenched by pouring into graphite crucible at 25°C. (150%)

 

Fia. 20-14. 10 w/o Th-90 w/o Bi. Dispersion of reconstituted ThBis particles
in Bi. Produced by heating fine-platelet dispersion to 800°C for 20 min. (150%)

ultrasonic energy to platelet dispersions causes the platelets to break up
into essentially equiaxed fragments.

20-6.5 Engineering studies of slurries. The intermetallic compound
ThBi2 is quite soft, having a Rockwell 15-T hardness of approximately
60 at room temperature. It is brittle at room temperature but appears to
exhibit some ductility at 400°C. The compound is pyrophoric and must be
protected against oxidation.

When slurries of equiaxed bismuthide in bismuth are prepared, they are
fluid at temperatures above the melting point of bismuth, 271°C. In
these slurries the solid phase is in thermodynamic equilibrium with the
20-6] THORIUM BISMUTHIDE BLANKET SLURRY 739

liquid phase and is perfectly wetted by it. At the proposed reactor tem-
peratures (350 to 550°C) practically all the thorium in the slurry appears
in the solid phase, since the solubility in the liquid is very low.

The ideal slurry composition represents a balance between a desire for a
high thermal neutron utilization factor (i.e., a high thorium content) and
the necessity for high fluidity. Fluidity studies have shown that the upper
limit of thorium concentration for high fluidity at reactor temperatures is
approximately 10 w/o of thorium. This corresponds to 24.9% by volume
of ThBi., and a thermal neutron utilization factor of 0.957. Although the
viscosity of Th-Bi slurries has not been measured, calculations based on
the viscosity of liquid bismuth and the behavior of similar systems indicate
that at 550°C the viscosity of a 10 w/o Th suspension of 50-micron, equi-
axed ThBis particles should be approximately 2.5 centipoises. It has been
observed that increasing the thorium content beyond 10 w/o Th causes a
disproportionately large increase in the viscosity, so that the consistency
approaches that of a mud or paste. The maximum thorium concentration
for high fluidity decreases when the ThBi. particle shape departs signifi-
cantly from an equiaxed shape.

The density of liquid bismuth varies from 9.97 at 350°C to 9.72 at 550°C,
and should not be changed appreciably by the small amount of thorium
dissolved at these temperatures. Therefore the solid particles should sink
in the liquid. Although settling rates have not been measured, the mag-
nitude of expected settling rates can be calculated. The settling rate for
100-micron spheres at 550°C, as calculated by Stokes’ Law, is 0.030 fps.
The settling rate in a 10 w/o Th-Bi dispersion of 100-micron spheres at
550°C, as calculated by the hindered settling equation, is 0.0026 fps.

It has been observed in small systems that equiaxed ThBiz particles
settle to a relatively stable layer in which the thorium concentration is
approximately 15 w/o Th. Such layers can be redispersed by mild agita-
tion of the supernatant liquid. When the thorium concentration in the
settled layer is increased to 18 to 20 w/o Th (by centrifugation, for ex-
ample), the viscosity of the layer is so high that mechanical agitation of
the layer itself is necessary to redisperse the particles.

Experiments have shown that the viscosity of a 10 w/o0 Th slurry, using
platelets of 50- to 100-micron size, is so high as to make the slurry com-
pletely unsuitable for use.

Slurry behavior under conditions of reactor blanket operation. It 1s antici-
pated that the slurry would be contained in a low-permeability graphite
within the reactor blanket. For heat removal and processing, the slurry
would be circulated externally through pipes and heat exchangers fabri-
cated of low-chromium steel or comparable material. During circulation
for heat removal, the slurry would be subjected to thermal cycling between
a probable maximum temperature of 550°C in the blanket and a possible
740 PROPERTIES OF LIQUID-METAL FUELS [cHAP. 20

minimum of 350°C in the heat exchangers. Capsule and pumped-loop ex-
periments have been carried out to study the behavior of the slurry under
conditions of thermal cycling and flow.

In the capsule experiments, small specimens of slurry are caused to flow
back and forth at 6 cycles/min in periodically tilted tubes fabricated of
the container material under test. The tubes, which are sealed under
vacuum, are heated to a higher temperature at one end than at the other.
When specimens of slurry containing 10 w/o Th, with and without addi-
tions of 0.025 w/o zirconium, were cycled between 350 and 550°C in 219,
Cr-19, Mo steel tubes, nearly all the ThBi2 was deposited-in the cooler
ends of the tubes in less than 500 hr. Examination of the deposits disclosed
that a deposit due to mass transfer of the steel had formed on the tube
walls prior to deposition of the ThBis. This suggested that mass transfer
of the steel may have been instrumental in starting the ThBi2 deposition,
perhaps by roughening the walls or perhaps by altering the composition
of the tube surface.

Specimens of 5 w/o Th slurries have been cycled for 500 hr between 350
and 580°C in graphite tubes with no evidence of plug formation. In these
experiments, a relatively rapid increase in ThBis particle size (from 50 to
225 microns in 500 hr) was observed. This increase was due to particle
agglomeration rather than growth of single crystals. No evidence of graph-
ite erosion was observed.

Specimens of slurries containing 10 w/o thorium and 0.10 w/o tellurium
have been cycled between 350 and 580°C in graphite, and between 350
and 580°C in 239, Cr-19, Mo steel for 500 hr with no evidence of ThBi-
plug formation or mass transfer of the steel. The specimens showed no
increase in the maximum particle dimension and no particle agglomeration.
When a specimen of slurry containing 10 w/o Th, 0.10 w/o Te was cycled
at higher temperatures in a 239, Cr-19, Mo steel tube, mass transfer of
steel and deposition of ThBiz in the cooler end were observed after less
than 100 hr.

Slurries containing up to 7 w/o Th and minor additions of zirconium
have been circulated through small 219, Cr-19, Mo steel loops by means
of a propeller pump. Isothermal circulation at 450°C has been carried
out for more than 450 hr at velocities between 0.3 and 1.5 fps, with no
difficulty in circulation or maintaining suspension. Attempts to circulate
these slurries through a temperature differential, however, have resulted
in the formation of ThBis2 deposits in the coldest section of the loop. In
a modified loop containing a graphite liner in the finned-cooler section,
isothermal circulation was maintained without difficulty. ThBig, however,
agaln deposited in the finned-cooler section when a temperature dlfferentlal
was applied.

When a slurry containing 7 w/o Th. 0.025 w/o Zr, and 0.10 w/o Te was
20-7] THORIUM COMPOUND SLURRIES 741

circulated in a 219 Cr-19,, Mo steel loop through a temperature differ-
ential, ThBis deposited in the finned-cooler section. The rate of buildup
of the deposit was markedly less than in the case of slurries containing no
tellurium.

The problem of ThBiz deposition during circulation through a tempera-
ture differential is one which must be solved before the Th-Bi slurry is
acceptable as a fluid breeder-blanket material. The favorable results ob-
tained by tellurium additions in the capsule experiments offer hope that
the problem can be solved.

20-7. Tuorium COMPOUND SLURRIES

20-7.1 Thorium oxide. Probably the best blanket material, next to the
thorium bismuthide slurry, is the suspension of thorium oxide in bismuth.
The thorium-oxide slurry should be compatible with the graphite and steel
in the reactor structure. Experiments have shown that ThO: is wetted
by the liquid bismuth if some zirconium or thorium is dissolved in the bis-
muth. Slurries of 10 w/o thorium oxide have been prepared.

The separation of thorium oxide from the liquid bismuth for processing
could be achieved by mechanical means, and the oxide could then be
processed by the existing Thorex process.

The thorium-oxide blanket slurry is gaining increased attention. A
loop of several pounds per minute capacity has been completed for forced
circulation of the oxide slurries at BNL and an 800 Ib/min loop is ready
at Babcock & Wilcox.

20-7.2 Other thorium compounds. A small amount of attention has
been directed toward ThCs, ThS, and ThF4 slurries in bismuth. How-
ever, the major effort is on the thorium bismuthide and thorium-oxide
slurries.
742 PROPERTIES OF LIQUID-METAL FUELS [crHAP. 20

REFERENCES

1. J. R. WEEKs et al., Corrosion Problems with Bismuth-Uranium Fuels, in
Proceedings of the First International Conference on the Peaceful Uses of Atomic
Energy, Vol. 9. New York: United Nations, 1956. (P/118, pp. 341-355); D. H.
Gurinsky and G. J. D1ENs, Nuclear Fuels. Princeton, N. J.: D. Van Nostrand
Co., Inc., 1956. (Chap. XIII); J. R. WEEKs, Metallurgical Studies on Liquid
Bismuth and Bismuth Alloys for Reactor Fuels or Coolants, in Progress tn Nuclear
Energy, Series IV, Technology and Engineering, Vol. I. New York: Pergamon
Press, 1956. (pp. 378-408)

2. J. E. ATHERTON et al., Studies in the Uranium-Bismuth Fuel System, in
Chemical Engineering Progress Symposium Series, Vol. 50, No. 12. New York:
American Institute of Chemical Engineers, 1954. (pp. 23-37); Nucleonics 4(7),
4042 (1954).

3. D. H. Aumany and R. R. BawpwiN, The Uranium-Bismuth System,
USAEC Report CT-2961, Iowa State College, 1945.

4. MASSACHUSETTS INSTITUTE OF TBCHNOLOGY, Progress Report for the Month
of October 1946, USAEC Report CT-3718. |
5. D. W. Barets, Liquid Reactor Fuels: Bismuth-Uranium System, USAEC

Report BNL-75, Brookhaven National Laboratory, 1950.

6. R. J. Terrer, Uranium-Bismuth System, J. Metals 9, 131-136 (1957).

7. G. W. GREENWOOD, personal communication to J. R. Weeks, Aug. 29, 1957.

8. O0.J. Ergert and C. J. EcaN, Dynamic Corrosion of Steel by Liquid Bismuth,
USAEC Report MTA-12, California Research and Development Co., 1953.

9. J. R. WeEks and D. H. Gurinsky, Solid Metal-Liquid Metal Reactions
in Bismuth and Sodium, in ASM Symposium on Liquid Metals and Solidification,
ed. by B. Chalmers. Cleveland, Ohio: The American Society for Metals, 1958.

10. C. R. Mirra and C. F. BoniLra, Solubility and Stripping of Rare Gases
in Molien Metals, USAEC Report BNL-3337, Columbia University Department
of Chemical Engineering, June 30, 1955.

11. W. G. McmMmILLAN, Estimates of the Solubility and Diffusion Constant of
Xenon in Liquid Bismuth, USAEC Report BNL-353, Brookhaven National
Laboratory, June 1955.

12. M. E. SkBERT, Investigation of Methods for Preparation of Thorwum
Bismuthide Dispersions in Liquid Bismuth, Final Progress Report, Horizons,
Inc., Oct. 31, 1956.

13. AEROPROJECTS, INC., Applications of Ulirasonic Energy, Progress Report
No. 4, USAEC Report NYO-7918, 1957.
CHAPTER 21
MATERIALS OF CONSTRUCTION—METALLURGY*

21-1. LMFR MATERIALS

21-1.1 Metals. Alloy steel. For maximum power production, it is de-
sirable to operate an LMFR at the highest possible temperature consistent
with the mechanical properties and corrosion resistance of the materials
of construction. A maximum temperature of 500°C or higher is deemed
desirable for economically attractive operation of the reactor. No ma-
terials have yet been found that are mechanically strong at these tempera-
tures, readily fabricable, and also completely resistant to corrosion by the
U-Bi fuel.

This does not mean that there is no hope for obtaining a good material
for holding bismuth fuel. On the contrary, very significant advances have
been made in the past few years. It must be realized that before work was
started on liquid metal fuel reactors, very little was known about the
solubility and corrosion characteristics of liquid bismuth with reference
to containing materials. There is general optimism that continuing research
and development will lead to suitable materials for containing the U-Bi
fuel system. ,

The low-alloy steels offer a good compromise for use in the heat ex-
changer, piping, and reactor vessel, particularly since their corrosion re-
sistance can be greatly improved by the addition to the fuel of Zr + Mg
as corrosion inhibitors. Nickel-containing stainless steels cannot be used,
despite their good high-temperature mechanical properties, because of
the high solubility of Ni in Bi, and the greatly lowered U solubility in the
presence of this dissolved Ni. Extensive engineering and fundamental
studies have been made on the corrosion of the low alloy steels by inhibited
U-Bj, as well as the mechanism of corrosion inhibition. Radiation effects
are currently being investigated. |

Of course, besides steels, there are other materials, notably the rarer
metals, which have characteristics making them suitable for certain uses
in a liquid-metal system. However, unless the cost and ability to fabricate
these materials can be improved significantly, heavy dependence will have
to be placed upon alloy steels for the main containment problem.

*Based on contributions by D. H. Gurinsky, D. G. Schweitzer, J. R. Weeks,
J. S. Bryner, M. B. Brodsky, C. J. Klamut, J. G. Y. Chow, R. A. Meyer, R. Bour-
deau, O. F. Kammerer, all of Brookhaven National Laboratory; L. Green, United
Engineers & Constructors, Ine., Philadelphia, Pa.; and W. P. Eatherly, M. Janes,
and R. L. Mansfield, National Carbon Company, Cleveland, Ohio.

743
744 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

21-1.2 Graphite. In the LMFR, graphite is considered as the principal
choice for the moderating material because of its availability, cost, and
knowledge of its characteristics under radiation. However, there are addi-
tional special requirements for the graphite in the LMFR system. It not
only is the moderator, but is also the container material for the U-B1
solution in the reactor. Hence it should be impervious to the liquid metal
and mechanically strong.

Experimental work at BNL has shown that graphite can be used directly
in contact with the fuel stream without danger of corrosion. By preferen-
tially reacting to form ZrC at the fuel-graphite interface, the Zr corrosion
inhibitor also prevents reaction of the U and fission products with the
graphite. Special grades of graphite are being developed that appear to
have the desired mechanical strength and low porosity required for use as
moderator and reflector in the reactor. Reactions of graphite with the
fuel, and the possible effects of pile radiation on these reactions, are de-
scribed in the following sections.

21-2. STEELS

21-2.1 Static tests. In order to attack the steel corrosion problem in a
basic manner, solubilities of the various components and combinations
have been determined. Most of these solubilities are given in Chapter 20.
However, more solubility work, important from a corrosion point of view,
is discussed here.

Solubility of steel components and inhibiting additives in liquid Bz. Iron.
The solubilities of iron in Bi, Bi+0.1%Mg, Bi+ 0.2% U+ 0.1%Mg, and
Bi+ 0.19%Mg+saturation Zr are given in Fig. 21-1. Uranium and Mg,
in the quantities added, have no effect on the iron solubility over the
temperature range 400 to 700°C. Zirconium increases iron solubility
slightly at temperatures above 500°C. Titanium (which might be present
as a corrosion inhibitor) has been found to decrease the iron solubility at
temperatures above 450°C, the extent of this decrease being proportional
(but not linearly) to the amount of Ti in the liquid. Below 400°C, there
appears to be a considerable increase in the iron solubility. For example,
Bi containing 1600 ppm Ti dissolved only 30% as much iron as pure Bi at
690°C, while Bi containing 300 ppm Ti (saturation) at 350°C dissolved
more than ten times as much iron as pure Bi.

Zircontum. The solubility of Zr in Bi is given in Fig. 20-5. This appears
to be unaffected by the presence of Mg, Cr, or Fe in the liquid metal.

Chromium. The solubility of Cr in Bi is given in Fig. 20-6. This also
appears to be unaffected by the presence of Mg, Zr or Fe in the liquid
bismuth. However, the presence of Cr in Bi causes a marked reduction in
the iron solubility.
21-2] STEELS 745

727 637 561 497 442 395 353
300 I | | | l

 

|

Fe in Bi+Mg+U

100

50

T TTTIT]
RN

|
I

[

I

Fe, ppm

CTTTTT
| d L]

?)

~w

—

I
l

 

 

 

] | | I | | ,
1.0 1.1 1.2 1.3 1.4 1.5 1.6
103/T°K

F1ac. 21-1. Solubility of Fe in Bi alloys.

 

Maiscellaneous data. The Fe—Zr intermetallic compound ZrFez appears
to decompose when added to Bi, Zr dissolving approximately to its normal
saturation and Fe somewhat in excess of its normal solubility in the presence
of Zr. The amount of excess Fe present in the liquid metal can possibly
be attributed to a finite solubility of the undissociated intermetallic com-
pound ZrFes.

The solubility of Ta in Bi is estimated to be less than 0.01 ppm (detec-
tion limit) at 500°C.

The solubility of Ni in Bi is close to 5% at 500°C and probably greater
than 1% at 400°C. |

The solubility of Mg in Bi is close to 4% at 500°C and 2% at 400°C.

Surface reactions. Experimental evidence has shown that the corrosion
resistance of steels in Bi is in part due to the formation of insoluble films
on the steel surfaces. The effect of these films on the corrosion behavior of
different steels is not readily determined by thermal convection loop experi-
ments because of the relatively low temperatures (400 to 550°C) and long
times associated with such tests. The comparative behavior of different
746 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

steels and different films is more easily obtained from high-temperature
(600 to 850°C), short-time, static contact tests.

Steel specimens approximately 1/2 in. wide, 2 in. long, and 1/8 in. thick
are cleaned and given various surface treatments, such as sandblasting,
chemical etches, polishes, etc. Six to ten different materials are then placed
in a vacuum furnace, heat-treated as desired, and immersed in a Bi alloy
containing the desired additives. The crucible used to contain the liquid
metal is either a material inert to Bi, such as Mo or graphite, or the same
material as the specimen. After contacting, the samples are removed from
the solution at temperature and allowed to cool in He or in vacuum. The
adherent Bi is removed from the steel by immersing in Hg at 200°C in a
vacuum or inert atmosphere. After rinsing, the residual adherent Hg is
completely removed by vacuum distillation at 100 to 200°C. The cleaned
surfaces are examined by x-ray reflection techniques, utilizing a North
American Phillips High Angle Diffractometer.

Surface reaction of zirconium, titanium, and magnesium. When pure iron
was contacted with bismuth containing radioactive zirconium tracer for
1 hr at 450°C, a Langmuir type adsorption of the zirconium on the iron
crucible surface was obtained. Increasing the temperature to 520°C and
the contact time as much as 24 hr showed an increased amount of reaction.
The structure of this deposit is not known. On the other hand, when pure
iron is contacted in saturated solutions of zirconium in bismuth for times
ranging from 100 to 300 hours at 500 to 750°C neither corrosion nor x-ray
detectable surface deposits occur. At concentrations of zirconium below
saturation value, pure iron is extensively attacked.

A tightly adherent, thick, uniform, metallic deposit was found on the
surfaces of pure Fe dipsticks contacted with liquid Bi saturated with Ti
at 650 to 790°C. In all cases the x-ray patterns were the same but could
not be identified. The 15- to 25-micron layers were carefully scraped off
and chemically analyzed. The results corresponded to a compound having
the composition FeTi4Bis.

Pure Fe and 239, Cr-19, Mo steel samples contacted with 2.5 w/o Mg
in Bi at 700°C for 250 hr showed no deposit detectable by x-ray diffraction,
Slight uniform intergranular attack was observed on all the samples.
Pure Fe samples contacted with Bi solutions containing 0.569, Mg
+ 170 ppm Zr, and 0.239, Mg 4 325 ppm Zr at 700°C were not attacked
and did not have detectable surface films. These solutions acted similarly
to those saturated with Zr.

Reactions of steels with UBt solutions. Uranium nitride (UN) deposits
have been identified on the surfaces of 5%, Cr-1/29, Mo, 219, Cr-19, Mo,
Bessemer, and mild steels, after these samples were contacted with Bi
solutions containing U or U 4 Mg. Extensive attack always accompanied
UN formation, indicating that this film is not protective. Nitrogen analyses
21-2] STEELS 747
made on these contacted specimens show that depletion of the N in the
steel 1s much more rapid than it is when the same steels are contacted with
solutions containing Zr.

Reactions of steels with Bt solutions containing combinations of Zr, Mg,
U, Th, and Ti. Deposits of ZrN, ZrC, and mixtures of the two have been
identified on many different steels contacted with Bi solutions containing
Zr with or without combination of Mg, U, and Th. No corrosion has ever
been observed on such samples contacted at 600 to 850°C for 20 to 550 hr,
nor have films other than ZrN or ZrC been found. When a mild steel was
contacted with Bi containing 1000 ppm Zr and 200 ppm Ti at 650°C, x-ray
examination showed strong lines for TiN and a less intense pattern of TiC.

Considerable difficulty was experienced in establishing the correct unit
cell dimension for the nitrides and carbides of Zr and Ti. Many different
values may be found in the literature. The inconsistency in the data
probably can be attributed to the existence of varying amounts of C, O,
or N in the samples. Table 21-1 gives the parameters determined by a
number of investigators. The values of ag used in this research were those
given by Duwez and Odell [1]. These compared favorably with the values
found on test specimens, powdered compact samples, and ZrN prepared
by heating Zr in purified N2 at 1000°C for 20 hr.

A nondestructive x-ray method of measuring film thickness has been de-
veloped for this research [2]. The x-rays pass through the film and are
diffracted by the substrate back to a counter. The intensity is reduced by
the absorption of the film. Unknown conditions of the substrate are
eliminated by measuring the intensity of two orders of reflection or by
measuring the intensity of a reflection using two different radiations. The
method is accurate to about 209.

TaBLE 21-1

PuBLisHED X-RAY PARAMETERS FOR THE UNIT CELLS OF
ZrC, TiC, ZrN, anp TiN
(CuBic, NACl-TYPE)

 

 

 

 

 

 

 

 

 

Becker and | Van Arkel | Kovalskii and | Dawihl and | Duwez and

Ebert [20] [21] Umanskii [22] Rix [23] Odell [24]
ZrC 4.76 4.73 4.6734 4 .685
TiC 4.60 4.26 4.4442 4 .31 4.32
ZrN 4.63 4.61 4 .567
TiN 4.40 4.23 4.234 4.236 4.237

 

 
748 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

TaABLE 21-2

ORIGINAL ANALYSES AND FiLMs FORMED ON SPECIAL STEELS
UseEp IN StaTic TESTS

 

 

 

. 9% Al %N % N as Film
Material o) | (moty | EHN® | 'EHN | formed

5Cr-4Mo 0.016 0.023 0.0002 1.0 ZrN
21Cr-1Mo 0.003 0.042 0.0001 0 )
21Cr-1Mo 0.055 | 0.050 — — ”
21Cr-1Mo 0.003 0.01 — — ”
21Cr-1Mo 0.06 0.047 0.0003 1.0 »
24Cr-1Mo 0.009 0.013 0.0001 1.0 ”
Bessemer 0.003 0.009 0.0002 2.0 »
Carbon 0.007 0.005 0.0001 2.0 ”
21Cr-1Mo 0.015 | 0.015 100 7xC
21Cr-1Mo 0.44 0.054 0.025 50 »
21Cr-1Mo 0.014 0.013 0.009 70 ?
21Cr-1Mo 0.022 0.015 0.010 70 »
21Cr-1Mo 0.02 0.015 0.011 75 »
1:Cr-1Mo 0.02 0.014 0.010. 70 »
RH 1081 (0.3 Ti) ”

 

 

 

 

 

 

 

 

*EHN: Ester-halogen insoluble nitrogen. This is believed to be an indication
of the nitrogen combined as AIN or TiN in steels [26].

Effect of steel composition and heat treatment. It has been found experi-
mentally that some steels with very similar over-all compositions behave
quite differently in the same static corrosion tests. Films that form on
these materials range from pure ZrN to pure ZrC. Table 21-2 gives typical
analyses selected from the more than 100 steels run in static corrosion tests,
and identifies the surface films. After contacting, the only changes in
analyses were found in the total nitrogen remaining and the amount of
ester halogen insoluble nitrogen (EHN) present in the steels. The only
significant difference in analyses between nitride-formers and carbide-
formers in Table 21-2 is found in the relative amounts of EHN. The
carbide-formers have more than 509, of the total nitrogen combined as
EHN, while the nitride-formers have only a few percent of the total nitro-
gen combined. At present, the relationship between the N, Al, Cr, and the
Mo contents of the steels and their film-forming properties is not obvious.
Some excellent nitride-formers have very low nitrogen content, while some
carbide-formers have high nitrogen content. The same holds true for the

\d
21-2] STEELS 749

Al Cr, and Mo contents of the steels. The EHN content of a steel can be
readily changed by short-time heat treatment at.700°C and higher [3], so
that this variable is controllable within limits.

To a first approximation, the corrosion resistance of a particular steel is
enhanced by high “inhibitor” concentrations and/or the presence of in-
soluble adherent films formed on the steel surface. The first of these con-
ditions is neither desirable nor practical in a solution-type fuel reactor be-
cause of the adverse effect of Zr on the U solubility. At present, work is
being done to measure quantitatively the effects of different alloying con-
stituents on the activities of N and C in steels. Consider the following
reactions:

Zr i) + N stee) == ZrN (im), (21-1)
ZrBi) + Cstee) === ZrC(1m). (21-2)
Assuming that the films are insoluble in Bi, then at equilibrium

1 1
K(ZrN) = M’ and K(ZrC) —_ W' (21—3)

If the products of the Zr activity in the Bi with the activities of the N and
C in the steel are not sufficient to satisfy the respective equilibrium con-
stants, the reactions will not occur, and the steel will not form ZrC or
ZrN films. If the activity products are greater than the constants, Kz:c)
or K (z:n), the reactions will proceed until the activities are lowered to these
values. Thus, for a fixed Zr activity, the activities of N and C in the steel
determine whether the carbide and nitride film-producing reactions should
occur. The excess of N or C above these equilibrium values should be a
measure of the driving force of reactions (21-1) and (21-2) to the right.

Solutzon rate tests. The solution rates of Fe into Bi, and Bi + Zr and Mg,
were measured in crucibles of a carbon steel, a 23% Cr-1% Mo, a 5%
Cr-1/2% Mo, and an AISI type—410 steel. The crucible, Bi, and additives
were equilibrated at 400 to 425°C, the temperature rapidly raised to 600°C,
and the concentration of Fe in solution measured as a function of time.
Results are shown in Fig. 21-2. In the presence of Zr+ Mg, the 5%
Cr-1/2% Mo and the AISI type—410 steels dissolved at approximately
the same rate, while the 23% Cr-1% Mo steel dissolved more slowly. No
detectable dissolution of Fe from the carbon steel was measured in 44 hr
at 610°C. These results are parallel to the thermal convection loop results,
and consistent with the film-formation studies in that the measured solu-
tion rates are inversely proportional to the ability and rate at which the
steels form ZrN films. At present no data are available on rates of solution
for ZrC-forming steels.
750 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

I
70 | | T | | T T 1T I I T

 

Solubility (Temp Cycle)
60 |— /5°/o Cr Steel into Pure Bi / JI }

12% Cr Steel into Pure Bi

........... en e
I .......
.t
o’
.
.

 

 

 

 

 

/ 12% Cr Steel into / —
. Bi + 0.01% Mg ) ) 5% Cr Steel into
/ ................... peeneees Bi + 0.1% Zr + 0.1% Mg
.o.-"'...“% - - ]
o ~"%<12% Cr Steel into Bi + 0.05% Zr
. / +0.04% Mg = __—-=—""
/ ________ l I' -
ool T T——12.1/4% Cr Steel into Bi + 0.05% Zr
—————— '~ /1020 Steel into Bi + 0.04% Zr +0.12% Mg
|-
| | l | | | I I |
2 3 4 5 6 7 8'2 40 6 80

Time in Hours

Fig. 21-2. Dissolution of Fe into Bi (plus additives) at 600°C from steel crucible.

Rates of precipitation. The rate of precipitation of iron from bismuth in
a pure iron steel crucible is very rapid. Iron precipitated from bismuth,
saturated at 615°C, as rapidly as the temperature could be lowered to
425°C. The addition of Zr plus Mg to liquid metal did not change the rapid
precipitation of most of the iron from the bismuth under these same condi-
tions, but produced a marked delay in the precipitation of the last amount
of iron in excess of equilibrium solubility. An apparently stable super-
saturation ratio of 2.0 was observed for more than 7 hr at 425°C in a
pure iron crucible containing Bi 4+ 1000 ppm Mg 4 500 ppm Zr, and 1.7
for more than 48 hr at 450°C. In a 5% Cr steel crucible, a supersatura-
tion ratio of iron in Bi+ Mg+ Zr of 2.9 was observed after 24 hr at 425°C.
This phenomenon may be due to the ability of the formed surface deposits
to poison the effectiveness of the iron surface as a nucleation promotor or
catalyst, the different supersaturations observed being due to the relative
abilities of a Zr-Fe intermetallic compound or of ZrN to promote nuclea-
tion of iron. This observed supersaturation suggests that mass transfer
should be nearly eliminated in a circulating system in which the solu-
bility ratio due to the temperature gradient does not exceed the meas-
ured ‘“‘stable supersaturation’ at the cold-leg temperature.

Precipitation rate experiments made in AISI type—410 steel crucibles
show that Zr + Mg stabilize Cr supersaturations of 2.0 to 3.0 for more
than 24 hr. However, no Cr supersaturation was found during precipita-
tion rate experiments made in pure Cr crucibles when Zr-4 Mg were
present in the melt [4]. The measured supersaturations should therefore
be due to the films present on the steel surfaces.
21-21 STEELS | 751

21-2.2 Corrosion testing on steels. The research effort on materials for
containment of the LMFR has been concerned mainly with low-alloy
steels having constituents which have low solubilities in Bi, such as C, Cr,
and Mo. Although the solubilities of Fe and Cr are only 28 and 80 ppm
respectively at the intended maximum temperature of operation, severe
corrosion and mass transfer are encountered when pure Bi or a U-Bi solu-
tion is circulated through a temperature differential in a steel loop. This
results from the continuous solution of the pipe material in the hot portion
of the system and subsequent precipitation from the supersaturated solu-
tion in the colder portions. Zirconium additions to U-Bi greatly reduce
this corrosion and mass transfer.

The behavior of steels in U-Bi is studied in three types of tests. Thermal
convection loops are used to test materials under dynamic conditions. In
these, the fuel solution is continuously circulated through a temperature
differential in a closed loop of pipe. Variables such as material composi-
tion, maximum temperature, temperature differential, and additive con-
centrations are studied in this test. More than sixty such loops have now
been run at BNL. The principal limitation in these tests is that the veloc-
ities obtained by thermal pumping are extremely low when compared with
the LMFR design conditions.

Forced circulation loops are used to study materials under environments
more closely approximating LMFR conditions. Three such loops are now
in operation at BNL and two more are under construction. A very large
loop (4 in. ID) which will circulate U-Bi at 360 U.S. gpm and transfer
about 23 X 10° watts of heat, is now under construction and is expected to
go into operation late this year.

Static tests, as discussed previously, in which steels are isothermally im-
mersed in high-temperature U-Bi containing various additives, are used
to study their corrosion resistance and the inhibition process as a function
of additive concentration and steel composition. Most of the tests have
been performed on a 2% Cr-19% Mo steel (Table 21-3). However, some
tests have also been made with higher Cr steels, 11% Cr-1/2% Mo,
1/2% Cr-1/2% Mo, and carbon steels.

21-2.3 Thermal convection loop tests at BNL. A typical thermal con-
vection loop that has been used at BNL is shown in Fig. 21-3. The loop
is provided with a double-valve air lock at the top of the vertical section
which permits taking liquid metal samples while the loop is running without,
contaminating the protective atmosphere. The hot leg is insulated and
heat is supplied to that section of the loop while the cold leg is exposed and
two small blowers are utilized to extract heat. The hottest point in the
loop is at the “‘tee’” at the upper end of the insulated section, and the coldest,
in the bottom of the exposed section. The total height of the loop proper is
[cHAP. 21

OF CONSTRUCTION—METALLURGY

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Fig. 21-3. Thermal convection loop. A. Air lock. B. Hot leg. C. Cold leg.
D. Fans. E. “Tee” connection. F.Melt tank with AISI type—410 steel filter
bottom.

approximately 15 in. and the total length of the loop is approximately
40 in. With this configuration, the flow rate is approximately 0.05 fps
when BI is circulated with a 100°C temperature differential. Temperature
differentials ranging from 40 to 150°C can be conveniently applied to the
loop. Radiographic inspection of the loop while in operation is periodically
made to monitor it for corrosion at the hottest section and deposition at
the coldest section. The inside of the steel pipe for the loop is either acid-
cleaned or grit-blasted. The pipe is then cold-bent to the desired shape,
and welded at the “‘tee’” by the inert-gas shielded-arc process.

The general procedure for running the loop is as follows: (1) Solid Bi
is charged into the melt tank. (2) The entire system is leak-checked with
a mass spectrometer. (3) The Bi is melted and introduced into the uni-
formly heated (550°C), fully insulated loop through a 35-micron AISI
[cHAP. 21

MATERIALS OF CONSTRUCTION—METALLURGY

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[cuAP. 21

OF CONSTRUCTION—METALLURGY

MATERIALS

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21-2] STEELS 157

type—410 stainless-steel filter. (4) Zr and Mg are introduced into the loop
through the air lock. (5) The Bi in the loop is sampled, using a graphite
sample extractor, to check for additive concentration. (6) Uranium is
added if additive concentrations are as desired; if necessary, additive
concentrations are adjusted prior to U addition. (7) The temperature dif-
ferential is obtained by removing the insulation from the cold leg and
starting the fans. The temperature differential is usually applied in two
steps, first 40°C and then the differential at which the loop is to operate.
(8) The entire loop is radiographed every 750 hr. (9) The liquid metal is
sampled at regular intervals. (10) After completion of the test the entire
loop is sectioned longitudinally and transversely for metallographic ex-
amination.

In Table 214, data from 29 thermal convection loop experiments at
BNL are summarized. The first three loops were fabricated from 239,
Cr-19, Mo steel pipe and the U-Bi solution was not inhibited. Loops No. 4
to 17 inclusive were made with 219, Cr-19, Mo steel and inhibited with
Mg and Zr. Loops No. 18 to 25 inclusive were fabricated from various
types of steels, ranging from Bessemer to 189, Cr-89%, N1 austenitic steels.
Loops No. 26 to 29 were made from 239, Cr-19, Mo steel pipe. The pur-
pose of these tests was to study the effectiveness of Ca, Th, and Ti as in-
hibitors.

It can be seen from the first three tests that deposition in the cold legs of
uninhibited loops after a few hundred hours is sufficient to stop flow. These
tests show conclusively that uninhibited U-Bi solution causes serious mass
transfer of this steel even at a temperature differential as low as 40°C.
Metallographic examination of the hot legs of these loops showed a gen-
eralized intergranular attack.

The data from loops inhibited with 350 ppm of Mg and 250 to 350 ppm
Zr (loops No. 4 to 17) show that mass-transfer rate can be decreased con-
siderably by the introduction of these additives. This effect is attributed
to the formation of a ZrN film on the steel surface [4,5]. The data demon-
strate, however, that this film is not completely protective in 2%, Cr-1%
‘Mo steel system. Test No. 4 indicates that for a differential in the order of
90°C, the film was sufficiently protective to prevent corrosion or deposition
in 13,550 hr of test. For temperature differentials of 100°C or higher, in-
cipient corrosion can be expected in about 5000 to 6000 hr. Some of the
21 Cr-1 Mo loop sections were joined with 5 Cr-1/2 Mo welding rod. This
higher chromium material has lower resistance to inhibited U-Bi than the
pipe, so that corrosion generally starts at these welds. Increasing the
temperature of the hot leg seems to increase the rate of corrosion, as il-
lustrated by the results of loops No. 5 to 9 inclusive, which were all tested
at 100°C temperature differential.
758 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

Loop No. 10, which was normalized from 954°C and tempered at 732°C
after welding, stood up poorly when compared with other tests. The heat-
treating was done in an argon atmosphere, and no subsequent alteration
was made to the surface left by the heat treatment. This heat treatment
was thought, from some observations in the pumped loops, to improve cor-
rosion resistance. It is believed that the poor results of test No. 10 are due
to alteration of the surface during the heat treatment and not to the
metallurgical structure of the steel. Loops No. 11 and 12 had Be and graph-
ite inserts in the hot leg to study their effect on mass transfer and also the
stability of U, Mg, and Zr concentrations. No detrimental effects on either
have been observed.

Loop No. 15, made of a 23 Cr—1 Mo steel that was internally nitrided to
a depth of 0.015 inch, appears to be standing up much better than other
21 Cr-1 Mo loops tested at a 125°C temperature differential. The added
nitrogen in the steel probably promoted the formation of the ZrN film. A
slight loss in Zr has been observed in this test, but other additive concen-
trations have remained constant.

Results of tests No. 18 to 25 show that higher-alloyed steels are inferior
to the carbon steels in inhibited U-Bi. Loop No. 18, fabricated from
Bessemer carbon steel, showed exceptional resistance to U-Bi. Metal-
lographic examination of this loop indicated no evidence of corrosion or
deposition after 12,356 hr of operation at 135°C temperature differential.
X-ray diffraction studies of a polished insert in this loop indicated that a
ZrN film was present; however, no film was detected on the pipe surface.
Considerable evidence of structural instability in the form of grain coarsen-
ing and graphitization of the Bessemer steel was found in the metal-
lographic examination of the loop. A medium carbon steel containing Ti,
Rh-1081, also exhibited good resistance to U-Bi corrosion. The lower
chromium steels, such as 1/29, Cr-1/29, Mo and 139, Cr-1/29, Mo
(tests No. 20, 21, and 22), also seem to be standing up well to inhibited
U-Bi. The testing of these steels will be increased.

Loops No. 26 to 29 were run to study the effectiveness of Th, Th 4 Zr,
Ti+ Mg, and Ca + Zr as inhibitors. It can be seen from these tests that
these inhibitors were much inferior to Mg and Zr combinations. In tests
No. 27 and 29 it was found that a slow but continuous loss of Zr, Th, and
Ca occurred. Horsley [6] also ran Bi loops containing Ca and Zr as in-
hibitors. He reported similarly that the loops plugged, and that Ca and Zr
were lost from the melt. The results of loop No. 28 indicate that Mg and
Ti provide some inhibition, but are not nearly as effective as Mg and Zr.
Metallographic examination of this loop shows that the corrosion was uni-
form and that most of the attack took place 4 to 6 in. downstream from the
“tee,” In an area which is normally somewhat lower in temperature than
the “‘tee.”
21-2] STEELS 759

Metallographic examinations of 239, Cr-19;, Mo loops inhibited by Mg
and Zr show that corrosion starts in the form of a pit. After the pit has
penetrated about 0.020 or 0.025 in. into.the pipe, the progress of corrosion
generally proceeds laterally on the pipe, widening the pit rather than
deepening it. A typical pitted area formed by U-Zr-Mg-Bi in 219, Cr-19,
Mo steel is shown in Fig. 21-4. The attack is transgranular throughout
the pitted area. Horsley [6] reported intergranular attack at the bottom
of pits formed in a similar steel by Zr-Ca—-Bi. Metallographic examination
of plugged thermal convection loops shows the deposit to be very flaky and
not tightly adherent to the loop walls. Chemical analysis of the deposition
in a 219, Cr-19, Mo steel loop indicated it to be about 959, Fe and 29, Cr.
ZrN films have been positively identified by x-ray reflection in only two
thermal convection loops: Bessemer steel loop, and one 239, Cr-19;, Mo
steel loop. The protective film is possibly so thin that it can be identified
only under ideal conditions.

21-2.4 High-velocity tests. Although thermal convection loops are
convenient for extensive corrosion testing under dynamic conditions, the
velocity is not large enough to give design data for actual operating condi-
tions. These data must be obtained by operating loops in which the bis-
muth solution is pumped at considerably higher linear velocities through
suitably designed test sections.

Attainment of higher flow velocities complicates corrosion testing meth-
ods. Large heat inputs are necessary to obtain temperature differentials
comparable to those readily attained in thermal convection loops. Special
equipment is required to measure flow. Pumps must be designed leaktight
to maintain absolute system purity. The U-Bi must be prevented from
freezing in the piping. |

Some important features of the BNL loops are (1) the systems are com-
pletely sealed and operate under a purified inert-gas blanket, (2) samples
of the liquid metal can be taken at any time during operation without
contamination, (3) radiography permits nondestructive examination of
test samples so that long runs are possible, and (4) the safety control system
is designed to prevent the freezing of the U-Bi (and subsequent bursting)
in the piping. '

In corrosion loops (HVL I and II) at BNL [7], no valves are used to
hold the fluid up in the system, and flow is measured with a submerged
orifice located in the sample tank. Piping in these loops is 3/4-in. schedule-
40 except in the test sections. A GE G—6 electromagnetic pump is used in
HVL I, while Callery 25-20 electromagnetic pump is used in HVL II.
Flows in the order of 1 to 2 gpm are achieved with both pumps. An inter-
mediate heat exchanger does about 50% of the heating and cooling. Re-
sistance furnaces provide the heat. The head developed is sufficient to
760 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

 

F1c. 21-4. Pitted area at ‘‘tee” in Loop #12. (a) Macrograph. (b) Micrograph
(original 250%) of pitted areas.
21-2] STEELS 761

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Fia. 21-5. High-velocity pump test loop.

allow the use of short small-diameter test sections in which velocities up to
8 fps are attained.

A third corrosion loop (loop G) uses a canned centrifugal pump to cir-
culate the U-Bi, permitting flows up to 4.8 gpm and velocities up to 14 fps
to be attained. Engineering devices such as bellows-sealed valves and
pressure transmitters are being tested in this loop, as well as materials for
high-velocity corrosion resistance. Heating and cooling is the same as in
HVL I and II. The loop is fabricated for the most part of 23% Cr-1% Mo
steel, 1-in. schedule—40 piping.

Two more loops (HVL III and IV), similar to loop G, are now under
construction. These are shown in Fig. 21-5. Flow rates up to 12 gpm,
velocities up to 25 fps, and temperature differentials of 150°C will be
attainable in these loops. HVL III will be fabricated of 23% Cr-1% Mo
steel, while HVL IV will be of 1% Cr-1/2% Mo steel.

While the basic material of construction of these loops is a Cr—Mo steel,
test sections are usually made up of a variety of steels and weldments.
All welds are made by the inert arc process. Cleaning, for the most part,
is done by grit-blasting.
[cHAP. 21

MATERIALS OF CONSTRUCTION—METALLURGY

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21-2] STEELS 763

 

Fig. 21-6. Precipitation of Fe-Cr alloy in high-velocity cold-leg test section
(HVL I-Run 1).

Results obtained thus far with the pumped loops are shown in Table
21-5. In most cases a run consists of a test on a particular test section rather
than on the entire loop. The exception is Run 4 of HVL I, which was
continued until the loop plugged, at which time the entire loop was dis-
mantled and refabricated of 13% Cr-1/2% Mo steel. The film temperature
reported is the calculated liquid-steel interface temperature. In all tests
Zr and Mg were used as inhibitors.

The purpose of Run 1 on HVL I was to study the effect of velocity on the
material deposited in the cold portions of the system. It was felt that a
high flow rate might dislodge the loose precipitated crystals and circulate
them to the hot sections, where they would redissolve in preference to the
pipe wall. The high-velocity section was placed at the region of the coldest
bulk liquid. All test specimens were fabricated of 219 Cr-1% Mo steel
and welds were performed with a 5% Cr-1/2% Mo bare filler rod.

Examination of the test sections after 1026 hr of test showed that no
corrosion had taken place in the hot sections. A one-grain, adherent layer
of precipitated Fe-Cr alloy was found in the high-velocity cold section
(Fig. 21-6).

To test the effect of impact on the inhibiting film, a right-angle, high-
velocity section was inserted in the hot leg of HVL I-Run 2. The section
was about 5 in. long, 0.355-in. ID and contained a graphite insert specimen
and an annealed and hardened 2% Cr-1% Mo steel specimen. All welds
were made with 5% Cr-1/2% Mo bare filler rod. The flow through the
764 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

  
  
 

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Cavitation !
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Cavitation
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Fia. 21-7. Cavitation-erosion on downstream side of right-angle bend in HVL
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F1a. 21-8. Preferential attack on 5%, Cr-49, Mo Weld in HVL I-Run 4.
21-2] STEELS 765

small diameter was 5 fps. Physically, the right-angle section was located
just outside the furnace. Polished bushings and tab samples were also
placed in the furnace leg. Temperature conditions were identical to Run 1.

This run was terminated after 1026 hr of AT operation, and samples
were removed for metallographic examination. No corrosion was detected
on the polished samples located in the furnace. No erosion or corrosion
was observed on any of the graphite samples. Again a very small amount of
deposition was found in the cold section, this time between the pipe wall and
the cold sample bushing inserts. Large pits, some about 1/2-in. diameter
and 1/8-in. deep, were found on the steel samples located in the exit side of a,
right-angle bend (Fig. 21-7). If a vortexing of the fluid (therefore a locally
increased velocity) is assumed to have occurred after the change of direction
in the right-angle bend, then a condition for cavitation may have existed,
i.e., the static head at this point (7 psi gauge) may have been exceeded by
the dynamic head.

Run 3 of HVL I was a duplicate of Run 2 except that the static head at
the right-angle bend was doubled. Examination of the specimen after the
1000-hr run showed that the cavitation pitting was eliminated.

Run 4 of HVL I was intended to test the inhibiting film under an in-
creased temperature gradient (150°C) with the maximum temperature
raised to 550°C. The test section consisted of samples of 23% Cr-1% Mo
steel (both annealed, and normalized and tempered), mild steel, and several
grades of graphite. All welds were made with 5% Cr-1 /2% Mo bare filler
rod. This run was terminated after 2591 hr of operation because of ex-
tensive pitting in the hot section and plugging in the finned cooler section,
as revealed by radiographs. Attempts to drain the loop completely were
unsuccessful. Upon sectioning, localized masses of intermetallic compounds
were found, which probably developed when the system was cooled. These
formations did not redissolve on heating because of the lack of good mixing,
and were viscous enough to prevent the bismuth from draining.

The entire loop was sectioned and examined. A severe pit-type corrosion
was found in the hot sections; welds (Fig. 21-8) as well as parent material
(Fig. 21-9) were grossly attacked; 2z% Cr-1% Mo steel in the normalized
and tempered condition was less corroded than the same material in the
annealed condition; carbon steel samples showed little or no attack.
Corrosion occurred in crevices between the samples and pipe walls. The
mass-transfer plugs occurred in the region of the coldest film temperature
rather than the coldest bulk temperature. Velocity did not sweep away
the deposited material, but rather a plug of high density was produced
(Fig. 21-10).

The minimum film temperature for this run may possibly have been as
much as 25°C lower than the 400°C reported because of an error involved in
selecting the actual cooling area.
766 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

 

Fia. 21-9. Attack on annealed 249, Cr-19%, Mo hot test specimen in HVL
I-Run 4.

 

Fig. 21-10. Cross section of deposited Fe-Cr alloy in finned-cooler section of
HVL I-Run 4.
21-2] | STEELS 767

Loop G-Run 1, had a hot-leg test section consisting of specimens of
2107 Cr-19, Mo, 139, Cr-1/2%, Mo, AISI type-410, and Bessemer steels,
joined with welds made with 239, Cr-19, Mo, 13% Cr-1/2% Mo, 5%
Cr-1,29%, Mo, AISI type—410, and mild steel bare filler rods. This run was
terminated after 938 hr of operation at a 75°C film gradient and a flow of 4
fps in the test section. Examination of the test section showed that welds
made with a 59, Cr-1/29, Mo rod were severely attacked, while those
made with a 219, Cr-19, Mo rod and a 119, Cr-1/29%, Mo rod were not
attacked. The only other corrosion observed in this run was some slight
attack on welds and base material of AISI type—410 steel. Cavitation pits
were also observed on the pump impeller.

Loop G—-Run 2, HVL I-Run 5, and HVL II-Run 1 are still under test.
In these loops are specimens of 229, Cr-19, Mo and 13% Cr-1/29% Mo
steels having various heat treatments, AISI type—410 and Bessemer steels,
and welds made with 119, Cr-1/29, Mo, 219, Cr-19%, Mo, 5% Cr-1/2%
Mo, AISI type—410 and mild steel bare filler rods. No corrosion has been
detected radiographically in the loop G test section after 2500 hr of opera-
tion. No corrosion has been detected in the HVL I-Run 5 test section;
however, slight deposition has been detected in the finned cooler. This
precipitate was first observed after 1400 hr of operation but is still not seri-
ous after 4000 hr.

Pitting of welds made with 5%, Cr-1/29, Mo rod and possible corrosion
of a 219, Cr-19%, Mo weld have been detected in Run 1 of HVL II. This
loop first operated 2500 hr with a film AT of 75°C (500 to 425°C film) with
no radiographically detectable corrosion and little or no mass transfer.
The temperature differential was then increased to 75°C bulk (522 to
427°C film). After 100 hr at this new differential, deposition in the cooler
was observed. Pitting of the 59, Cr-1/2%, Mo weld was detected after
2000 hr at the new differential. However, after a total of 7400 hr of tem-
perature gradient operation, the amount of pitting and transferred ma-
terial was not serious enough to stop loop operation.

The effect of velocity on corrosion and mass transfer is not apparent at
this time. This is mainly due to a lack of tests in which velocity is the only
variable. Correlation between the pump loops and thermal convection
loops suggests that velocity has little effect on mass transfer other than on
the type of plug formed, provided conditions for cavitation do not exist.
However, results from tests now under way should definitely evaluate this
variable.

21-2.5 Rapid oxidation of 239, Cr-19, Mo steel. In a pumped Bi loop
containing Mg-+Zr, the 219, Cr-19, Mo steel adjacent to a pinhole leak was
found to be severely oxidized. The appearance of the oxide scale was very
similar to that reported by Leslie and Fontana [8]. These investigators
768 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

found that rapid oxidation of a 169, Cr-259, Ni—69, Mo steel will occur in a
stagnant atmosphere, and that MoOs3; catalyzed the rapid oxidation of this
steel. They also reported that Bi2O3 could produce a similar effect.

A test simulating a leak in a loop was made to study this phenomenon.
A 239, Cr-19, Mo steel pipe was filled with Bi containing 1000 ppm U,
350 ppm Mg, and 250 ppm Zr. A 1/32-in. hole was drilled in the pipe below
the Bi level. A patch of asbestos tape 2 in. in diameter was put over the
hole. The tube was then heated to 594°C and pressurized to 5 psi to force
out a small amount of Bi. The pressure was dropped as soon as some Bi had
leaked out and the tube was then heated at 594°C for 1025 hr. The appear-
ance of the pipe underneath the asbestos tape patch and a cross section of
the oxidized area are shown in Fig. 21-11. A complete oxidation through
the pipe wall has occurred in the area adjacent to the leak.

Tests run at 738 and 816°C in covered crucibles show that chemically
pure Bi2O3 can promote rapid oxidation in 239, Cr-19, Mo, 119, Cr-1/29,
Mo, and carbon steels. Tube tests and crucible tests are being continued to
determine the minimum temperature at which rapid oxidation is a problem.

21-2.6 Radiation effects on steels. Since steel will be in direct contact
with the liquid U—-Bi fuel, it is important to determine the effects of fission
recoils and fast neutrons on the rate of corrosion or erosion. If corrosion
inhibition is achieved by a layer of ZrN between the Bi and the steel, there
is concern that fission recoil particles might destroy this film. Local heating,
resulting from the stopping of the particles, may cause a differential
expansion between the layer and the steel. Consequently, the layer may
break away from the steel, leaving the surface exposed to Bi attack. On
the other hand, neutrons should not be detrimental to a thin ZrN layer,
per se.

Effects of neutrons must also be determined on the mechanical properties
of the steels at reactor temperatures. Radiation-induced increases in tensile
strength and elastic modulus may not anneal out at LMFR operating
temperatures. A decrease in the impact strength is not considered too
probable at these temperatures, although the possibility must be
investigated.

To study the-radiation effects on materials, a capsule test has been de-
veloped. The capsules, in which samples are inserted in a highly enriched
U-DB1 solution containing Mg and Zr inhibitors, are exposed in the BNL
reactor and irradiated to the desired level. The test has the advantage of
attaining high temperatures (700°C) and a high fission recoil density.

Samples of 139, Cr-1/29%, Mo and 219, Cr-19,, Mo have been placed in
one of these capsules with Bi containing 4000 ppm U, 2500 ppm Zr, and
3500 ppm Mg. The exposure was approximately 2.2 X 10'° nvt. Sectioning
of this capsule has shown no corrosion of the samples. A second capsule
21-2] . STEELS 769

 

Fi1a. 21-11. Rapid oxidation of 239, G-19, Mo steel pipe after 1025 hr at 594°C
at pinhole leak.

has been put in the BNL pile containing samples of 139, Cr-1/29, Mo and
mild steels. Results are not yet available.

Dynamic tests of the reaction between Bi and steel in the presence of a
radiation field must be completed before a final selection can be made of
materials for the LMFR. The effect of velocity on corrosion is not certain
from the out-of-pile studies, so that no exact analogy can be made between
out-of-pile forced circulation loops and in-pile capsules. There has been
limited work done at Harwell [6] with thermal convection loops in and out
of a radiation field. These loops had no U but did contain Ca and Zr
inhibitors. The data suggest that pile radiation may have induced some
acceleration of mass transfer.
770 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

An in-pile, forced-circulation loop has been built at BNL and two others
at Babcock & Wilcox Research Laboratory to test the corrosion stability
of LMFR materials under conditions to be expected in the reactor experi-
ment. In this loop, Bi containing approximately 1500 ppm of U235, 180 ppm
Zr, and 350 ppm Mg will be pumped at a rate of 56 to 7 gpm. The bulk
AT will be approximately 75°C, with a maximum temperature of 500°C.
There are three sample sections in the loop: one, containing samples of
129, Cr-1/29, Mo steel, 239, Cr-19; Mo steel, Be, and graphite, will be at
the center of the reactor and will be in a flux of approximately 3 X 103
thermal neutrons; one within the shield will see delayed neutrons at a
temperature of 500°C; and the third section will be outside the reactor at a
temperature of 425°C. This test is presently being assembled, and will be
operating late in 1958.

21-3. NONFERROUS METALS

Several nonferrous metals are receiving attention as container materials,
principally because of their low solubility in bismuth. At this time only a
small amount of work has been done on testing these metals. The following
is a brief discussion of some of the most important of them.

21-3.1 Beryllium. In the Liquid Metals Handbook, beryllium has been
reported to have a good resistance to attack by liquid bismuth at 500°C
and probably also at 1000°C. Recent work at Harwell has shown that
beryllium has good resistance to mass transfer by liquid bismuth circulating
through a temperature gradient of 500°C, with a base temperature of
300°C. Recent advances in the production technology of beryllium metal
have improved its mechanical properties sufficiently so that it is now pos-
sible to consider this metal for reactor core vessel or moderator. However,
no data on reaction between beryllium and uranium additives and fission
products dissolved in a liquid metal fuel are available. Since uranium is
known to form a stable intermetallic compound with beryllium, the possi-
bility of such a reaction must be investigated before beryllium can be
recommended as a moderator in contact with a liquid-metal fuel. Static
tests on the resistance of Be to attack by U-Bi at 550 and 650°C showed
some (but not conclusive) evidence of a slight attack (less than 1 mil/yr
average penetration). A slight attack on Be was noted by Brasunas by a
29, U-Bi alloy after 4 hr at 1000°C.

21-3.2 Tantalum. Recent work at the Ames Laboratory indicates that
tantalum is an excellent material to contain U-Bi. A 5 w/o U in Bi solution
was circulated through a 3/4-in. OD by 0.030-in. wall tantalum loop at a
rate of 800 Ib/min. The U-Bi was circulated with an electromagnetic pump
and the temperature differential was 100°C (950 to 850°C).
-

21-4] BEARING MATERIALS 771

Liquid metal samples taken during operation showed that the Ta con-
centration never exceeded 6 ppm. After 5250 hr of operation, the loop was
shut down and examined metallographically. No transferred material was
detected in the cold leg and the corrosion was less than 1 mil. The tantalum
remained shiny and ductile throughout the experiment.

Even though the above results are hopeful, the use of tantalum as a
container material for a U-Bi reactor system is limited by (1) its poor air
oxidation resistance, (2) difficulty of fabrication, (3) cost.

21-3.3 Molybdenum. Molybdenum is also known to be highly resistant
to U-Bi. However, its use as a container material is hampered by its poor
oxidation resistance and difficult fabrication.

Some success has been obtained, however, with Mo applied as a coating
on low-chrome steel. The process was developed by the Vitro Corporation
and is described in detail in KLLX-10009. Essentially, the coating is applied
by first electrophoretically depositing 809, Mo + 209, MoOs on the steel.
The coating is then pressed, reduced in a Ho atmosphere, repressed and
sintered in an Hg + HCI atmosphere.

These specimens have been subjected to static Bi which was temperature
cycled at 550 to 400°C. Filtered liquid metal samples of the Bi were taken
and analyzed for the presence of Fe and Cr. The Bi was allowed to stay at
the 550°C temperature overnight before a liquid metal sample was taken.
After 4300 hr at temperature and some 17,000 cycles, the concentration
of Fe was 6.9 ppm and Cr 2 ppm. By comparison, the solubilities of Fe
and Cr at 550°C were 30 ppm and 80 ppm respectively.

On the basis of their low solubility in U—Bi such materials as tungsten,
tantalum, molybdenum, and beryllium could be classified as container
materials. Their practical use, however, is governed by their poor air
oxidation resistance, fabrication difficulties, availability, and cost.

21-4. BEARING MATERIALS

The relative bearing properties of materials for use as valve seats, disks,
and guides are being determined in Bi containing Mg and Zr. Testing is
done under a boundary lubrication condition.

The test apparatus consists of a main test chamber and a sump tank to
facilitate easy handling of the liquid Bi and to permit periodic sampling of
the liquid metal. The main test chamber, containing the samples and the
sample loading mechanism, is shown in the cutaway of Fig. 21-12. The
method of transmitting the load with air pistons through flexible bellows
can be clearly seen. A second chamber mounted on top of the main test
chamber contains a pc Thymotrol motor which rotates the cylindrical
specimen. The power input versus torque characteristics of the motor when
772 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

Motor Shaft Bellows Seal

Crank

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F1e. 21-12, Bearing materials test apparatus.

operated in a helium atmosphere using “‘high-altitude” aircraft brushes
has been determined.

A test consists of contacting the rotating cylindrical specimen with a
flat specimen under a constant force for a set period of time under set condi-
tions. After the contact run, the Bi is removed from the specimen. The
degree and kind of scoring, galling, material transfer, and depth of wear on
the surface were noted. Surface roughness measurements are made with a
profilometer. Hardness measurements are made. Coefficients of friction
are calculated from the measured torque data by the equation

T
=2 (21-4)

where f= the coefficient of friction, P = the applied load (lb), r = the
radius of sleeve (ft), and T = the torque (ft-lb).

In general, the hard-to-hard material combinations have shown good
wear resistance except for some scoring. The best hard material tested
thus far i1s AloO3 flame-coated on AIST 4130 steel. When this material was
contacted against itself, no wear or scoring could be detected. This ma-
21-5] SALT CORROSION 773

terial will be thermally cycled and exposed to inhibited U-Bi for long times
to evaluate its utility. Stellite 90 and Rex AA also behaved well. Contacts
made with common die steels and low-alloy steels have exhibited severe
scoring and wear. Corrosion has also been detected on these samples. Of
the cemented carbides, only TiC with either a mild steel or 239, Cr-19, Mo
binder has been tested. This material did not show good wear properties
and also exhibited some pitting corrosion.

Graphitar versus tool steel and Mo versus Rex AA or Stellite 90 have
shown the best results of the hard versus soft combinations tested. The
results have been good in that the wear has been very smooth; however,
the wear has been excessive. The use of these combinations would be limited
to very low-load applications.

21-5. SALT CORROSION

In earlier chapters, it was pointed out that one of the chief advantages
of the LMFR lies in the possibility of easy chemical processing. Several
processing techniques have been studied, most of which are based on
pyrometallurgical processes. The two chief pyrometallurgical methods
under consideration are the chloride process, in which the bismuth fuel is
contacted with a ternary mixture of molten chloride salts, and the fluoride
process, in which the bismuth fuel is contacted with molten fluoride salts
containing hydrogen fluoride. As may be imagined, the construction
material problem for these plants is very difficult.

A corrosion test program is actively under way at BNL and Argonne
National Laboratory on the chloride and fluoride processes respectively.
At BNL, these tests have consisted principally of rocking furnace and tab
exposure tests.

In the rocking furnace test, a piece of tubing approximately 12 in. long
and 1/2 in. ID, containing a charge of either salt or a mixture of salt and
bismuth, is placed on a rocking rack in a furnace. This rack alternately
tilts to one end for a period of 1 min and then to the other end for a like
amount of time. The two ends of the furnace are kept at 450 and 500°C
in order to give a temperature differential and thus induce mass-transfer
corrosion. The standard test period has been 1000 hr. These tests are part
of the initial screening program. When they are completed, the metals
which have given the best performance will be further evaluated in test
loops and pilot-plant equipment.

At present, only molybdenum has been satisfactorily tested against a
mixture of salt and bismuth fuel. However, the results are definitely en-
couraging. It has been found that the ternary salt, MgCl:—NaCl-KCl,
with or without zirconium and uranium chlorides, can be contained fairly
well in austenitic stainless steels, particularly 347 stainless steel. When a
774 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

mixture of bismuth fuel and the ternary salt containing less than 1% BiClj
was tested, the ferritic stainless steels were the best materials. These
include 410, 430, and 446 stainless steels. Probably the best of the ferritics
is the 239 Cr-19% Mo stainless steel.

During one step in the chloride chemical process, it is necessary to have
the ternary salt, containing more than 1% BiCls, in contact with bismuth
fuel. For this mixture only molybdenum has been satisfactory. However,
considerably more testing is required before this can be considered a
satisfactory material. |

The experience in handling salt with larger-sized equipment is quite
limited. A small loop built of 347 stainless steel has been operated satis-
factorily for a fairly short time. A much larger loop, loop “N,” is now
being censtructed at BNL. This will contact the chloride salt and the
bismuth fuel. The salt part of the loop is constructed of 347 stainless steel.
The bismuth fuel section of the unit is constructed of 23% Cr-1% Mo
steel. The actual contacting units are constructed of both 347 and the
low-chrome steels. This pilot plant, when placed in operation, should
furnish considerable information on the corrosion characteristics of the

- molten chloride salt.

' The fluoride process also presents difficulties with materials of construc-
tion. The mixture of the molten fluoride salts, containing HF, with the
bismuth fuel is extremely corrosive. Pure nickel has been found to stand
up fairly well to the molten fluoride salts alone. However, the combination
of the three materials has proved to be very corrosive even to nickel. The
extensive development program to investigate the materials of construction
for the fluoride process is continuing.

21-6. (GRAPHITE

21-6.1 Mechanical properties. In the proposed LMFR system, the
moderator, graphite, is also employed as the container material. Therefore,
the graphite should have good physical properties such as strength, hard-
ness, and resistance to shock. Since graphite is to be the container material
for the bismuth solution, it should theoretically be completely impervious
to the solution. For this reason, special graphites, more impervious than
the usual reactor grades, have been developed and are under development.
Physical properties of typical examples of these graphites are given in
Table 21-6. In comparison with the usual reactor grade, AGOT, having
a compressive strength of 6000 psi, these impervious grades have a strength
of 6500 to 9700 psi.

Another special requirement for the graphite is that it withstand erosion
or pitting by the flowing fuel. Test sections of accurately bored graphite
were placed in test loops where the flow velocity of bismuth was 6 to 8 fps.
No observable effect was noted after 1000 hr of test at 550°C.
21-6] GRAPHITE 775

Although tests so far have been on rather small samples, the mechanical
properties of these improved graphites appear sufficiently good for use in
LMFR systems. These new graphites must be manufactured in large sizes
in order to conveniently make up the core of an LMFR. The graphite
industry in the United States is now developing manufacturing techniques
for making such large sizes.

21-6.2 Graphite-to-metal seals. Leaktight joints of steel to graphite are
required at several places in the core of an LMFR. These seals must with-
stand an average of 125 psi at approximately 550°C. This is done by joining
finely machined steel and graphite surfaces under sufficient spring loading
to prevent bismuth leaking across the seal.

Tests were run by Markert at the Babcock & Wilcox Research Center
to evaluate such pressure seals. Three-inch and six-inch steel pipes
(24% Cr-19% Mo) with machined ends were pressed against a flat surface
of a block of MH4LM graphite (Great Lakes Carbon Co., density 1.9 g/cc).
The graphite surface had been prepared by sanding and polishing with
No. 000 emery paper. A seal was effected against Biat 438° with a pressure
differential across the seal of 100 psi, and with 1500 psi stress between the
graphite and the steel. The minimum stress that may be used without
visible Bi leakage at this pressure differential was found to be as low as
600 psi. It was not necessary to resort to complicated interface configura-
tions to obtain a seal. These initial results are very encouraging, and
further development work is being directed toward more complicated seals.

21-6.3 Graphite reactions. If graphite is to be in direct contact with the
U-Bi fuel, it should be inert to the various fuel constituents and also to
fission products and corrosion products. Work has been done at various
locations on these reactions. Thermodynamic data on chemical equilib-
rium, when available, have proved to be extremely valuable in guiding
the experiments.

Uranium-graphate reactions. The reaction between uranium and graph-
ite is probably the most important one to consider in the LMFR. Mallett,
Gerds, and Nelson [11] reported that uranium forms three stable carbides:
UC, UCs, and U2C3. Further work on this subject [7,12] indicates that
when less than 19, U is present in bismuth, it does not react with graphite
to form carbides at temperatures below 1200°C.

However, the nitride of uranium, UN, has been identified on graphite
contacted with 0.059, U in Bi at 850°C for 28 hr. This nitrogen was un-
" doubtedly adsorbed on the surface of the graphite and had not been dis-

lodged by outgassing at high temperatures and vacuum. |

When zirconium and magnesium are present with uranium in the bis-
muth, zirconium reacts preferentially with the graphite to form ZrC. This
[cHAP. 21

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778 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

is predictable from the chemical thermodynamic data. An experiment in
which graphite was contacted with 1000 ppm U, 50 ppm Zr, and 300 ppm
Mg in Bi at 1000°C for 8 hr showed only a single intense x-ray diffraction
line corresponding to the most intense line for ZrC. The x-ray analysis
was carried out after the adherent bismuth was removed from the samples
by mercury rinsing.

These experiments indicate that the reaction of uranium with graphite
is not likely to' occur under the LMFR operational conditions and can be
prevented by the addition of zirconium to the bismuth.

Zirconium and titantum reactions with graphite. Since zirconium and pos-
sibly some titanium will be present in the bismuth, reactions of these
materials with graphite have been investigated. As described above, zir-
conium reacts to form the carbide with a strong negative free energy. At
temperatures around 550°C, ZrC and solid solutions of ZrC and ZrN have
been identified on graphite surfaces contacted with bismuth solutions con-
taining 130 ppm Zr.

On the other hand, no reaction between graphite and 1600 ppm Ti in
Bi solutions has been observed up to 800°C for contact times up to 170 hr.
A strong TiC x-ray pattern and a less intense ZrC—ZrN solution pattern
were observed on graphite contacted with approximately 0.29, Ti and
0.29, Zr in Bi at 1250°C for 44 hr.

When steel samples are reacted with U-Bi fuels containing zirconium
and magnesium, the x-ray patterns of the surface are those for pure or
very nearly pure nitrides or carbides. When graphite is contacted with the
fuel, however, solid solutions of the carbide and nitride are often found.
The unit cells vary from 4.567 Kx* to 4.685 Kx for the zirconium com-
pounds and from 4.237 Kx to 4.320 Kx for the titanium compounds. These
parameters are low for complete carbon carbide structures.

Parameters for carbon-deficient carbide structures have been reported in
the literature [13]. For ZrC the reported ao varied from 4.376 Kx at 20
atomic percent C to 4.67 Kx at 50 atomic percent C. However, up to the
present time no evidence of carbon deficient structures has been observed
in studying the graphite-fuel experiments. The low parameters are instead
believed due to nitrogen replacing the carbon atoms in the carbide lattice
(NaCl-type). Parameters for such solid carbide-nitride solutions are de-
scribed by Duwez and Odell [1].

Fission product-graphite reactions. The products of uranium fission may
also react chemically with graphite to form carbides. A series of experi-
ments have shown that materials such as cerium will definitely react with
graphite. When 25 ppm Ce in bismuth was placed in contact with graph-
ite at 700°C for 110 hr, CeC2 was identified as a film on the graphite.
Graphite contacted with 140 ppm Sm in bismuth at 800°C for 140 hr, on

*Kx = 1000x units = 1.00202 4 0.00003A..
21-6] GRAPHITE . 779

the other hand, gave an x-ray diffraction pattern of the graphite surface
which could not be identified or indexed. Under similar reaction conditions,
neodymium, barium, or beryllium in bismuth solutions have no reaction
product. However, Miller [12] has shown by an autoradiograph technique
that 180 ppm irradiated Nd in bismuth reacts with graphite at 1100°C in
100 hr, concentrating the radioactive Nd at the graphite-liquid metal in-
terface. When the same experiment was repeated with the addition of
100 ppm Zr to the solution, no Nd was identified at the graphite surface.
This experiment indicates that zirconium will probably form the car-
bide and nitride preferentially to most fission products. However, further
research is required to determine whether a zirconium carbide-nitride layer
on the graphite can be depended upon to prevent the adhesion of the fission
products to the graphite surface. This point is not only important from
the chemical and graphite surface point of view, but is also important from
the neutron economy point of view. To obtain the highest breeding ratio,
the fission products, which are all fairly good neutron adsorbers, must be
removed from the graphite core soon after their formation. This is espe-
cially true of samarium, which has a very high neutron absorption cross
section. Therefore, the experiments reported above on zirconium are quite
encouraging in that there is no trace of a samarium film on the graphite.

21-6.4 Radiation effects on graphite. The graphite core, in order to
serve as moderator and container for the flowing fuel, must be stable to
radiation. Fission recoils may cause spalling and reduction in the thermal
conductivity that might increase the thermal stresses within the core struc-
ture. The graphite must not adsorb large quantities of U or fission products.

A capsule test has been developed in which samples are irradiated in a
highly enriched U-Bi solution containing Mg and Zr inhibitors for study
of radiation effects on materials. The test has the advantage of attaining
high temperatures (700°C) and a high fission recoil density.

Graphite samples have been exposed in these capsules under conditions
given in Table 21-7. Metallographic examination of these graphite sam-
ples indicates that there is no excessive spalling or corrosion.

However, some samples were treated with Bi containing Zr at 1300°C
to obtain 10- to 30-micron ZrN-ZrC layers on the graphite prior to irradia-
tion, and postirradiation examination indicated some change of this layer.
The cause of these effects is still being determined.

The effect of neutrons on the growth and thermal conductivity proper-
ties of special low-permeability grades of graphite has been measured.
Three sets of samples have been irradiated to exposures as great as 5 X 1020
thermal neutrons/cm? in the temperature range 400 to 475°C. Results of
these tests are listed in Table 21-8. The change in physical length of the
graphite is primarily contraction and should not present a major engineer-
ing problem.
[cHAP. 21

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782 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

On the other hand, the neutrons reduced the thermal conductivity by
25 to 359, of the preirradiation value. Moreover, some of these graphites
had a preirradiation conductivity only 659, that of the usual reactor-grade
graphite. The combination of low permeability and neutron irradiation
therefore reduces the thermal conductivity to 509, that of the usual reactor-
grade graphites.

Bismuth penetration into graphite permits the diffusion of some uranium
into the graphite. The resulting fission recoil particles may further reduce
the thermal conductivity of graphite. Tests are now under way to deter-
mine the degree of damage produced.

21-6.5 Bismuth permeation and diffusion into graphite. Early in the
work on the Liquid Metal Fuel Reactor it was recognized that a special
type of graphite was required if it were to be both a moderator and a con-
tainer for the reactor fluid. Such an impermeable graphite would have not
only the usual advantages of being both structural material and moderator,
but would also not hold up quantities of coolant or fuel, with resulting
decreased neutron efficiency and control. Besides these characteristics,
because the design fluxes are of the order of 10!5 n/(cm?)(sec) it is necessary
that the graphite have a high degree of resistance to radiation damage,
specifically to physical growth and reduction in thermal conductivity.

In considering impermeable graphite, two characteristics of the graphite
are concerned: liquid pickup and permeability. Liquid pickup refers to
the amount of fluid which is held in the interior pore volume of the graphite
in the manner of a sponge. Permeability is the rate at which a fluid can be
made to flow through the graphite. Both these properties depend primarily
on the accessible void volume and the pore spectrum.

In research work on graphite, it is customary to divide the size of pores
into two categories: macropores larger than 1 micron and averaging 2.5
microns in radius, and micropores with radii less than 1 micron and pre-
dominantly below 0.5 micron.

The development of new types of “impermeable’” graphite has necessi-
tated an examination and improvement of manufacturing processes con-
current with experiments on bismuth uptake [14].

The classic process of graphite manufacture is based on cokes and pitch
binders, baking and pitch reimpregnations, and finally graphitization.
Around this scheme has evolved a complex technology involving careful
particle and flour sizing to obtain optimum compaction, elaborate baking
and graphitizing schedules, and extremes of pressure vacuum treatments.
The production of the new relatively impermeable grades was made pos-
sible by three significant advances over the older technology: (1) the ability
to use raw materials of more advanced form, including graphites and blacks
of various types, (2) impregnation techniques now span a variety of resins
21-6] GRAPHITE 783

with various viscosity and wetting properties, and (3) new forming tech-
niques permit much more uniform and finer-grained artificial graphites.
This last includes, in particular, the development of pressure baking, where
heat is applied by passing electric current through the carbon in the mold
and under pressure.

The conventional pitch-type impregnation has the effect of increasing
density without markedly reducing permeability. This is apparently due
to the tendency of the pitch to coke out only in the voids of large effective
pore radius. Conversely, the newer impregnating materials, with their
lower viscosities and increased wetting, tend to block the pores themselves
as well as the larger open volumes. Consequently, there is no general rela-
tionship between density and permeability. Several conclusions may be
drawn: (1) The newer impregnates primarily attack the macropore distri-
bution and shift it from the 1- to 5-micron range into the submicron range.
(2) Optimum particle packing in the original base material is, in general,
not an advantage upon reimpregnation. (3) It is essential to use base ma-
terials in which the long tail at high pore radii is missing. The relatively
minor effect of the present impregnates upon the micropore distribution
demonstrates that the present materials, markedly improved as they are,
do not as yet represent the achievable ultimate.

The amount of bismuth uptake in graphite is probably the most im-
portant property concerned in evaluating the graphite, and was one of the
first investigated. For this purpose, a simple pot arrangement is used to
hold samples of graphite in molten bismuth at pressures from vacuum to
550 psi and at temperatures from 550°C. The graphite samples (0.5-in.
OD and 1.75-in. long) are outgassed in a vacuum at 550°C and then sub-
merged in the bismuth. Helium pressure is then applied to the molten bis-
muth. The amount of bismuth uptake into the graphite is determined by
the difference in the density of the sample before and after submersion.
The accuracy of this measurement is within 0.01 g bismuth per cc graphite.

The degree of bismuth uptake by graphite as a function of time, de-
termined at a pressure of 250 psi at 550°C, is shown in Fig. 21-13. As was
mentioned above, there is no correlation between uptake and graphite
density. The densities of these graphites range from 1.73 to 1.92 g/ce. The
total percent of void volumes in the impregnated grades varies from 16.0
to 19.0% of the bulk volume. Of these totals, the inaccessible volumes
range from 6 to 10%. Although samples EY-9 and ATL-82 have nearly
the same density, the bismuth absorption differs by as much as a factor of
3 because of the difference in pore spectrum.

As can be seen from the figure, the amount of bismuth uptake varies as
a function of time. This behavior was obtained using separate samples for
each point on a curve and also by measuring the same sample at various
time intervals. In making determinations, considerable oscillation about
784 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

 

l l I | |

    
 

   

Right Hand
Scale

    

Bismuth Uptake, grams/cc graphite

 

 

 

Immersion, Hundreds of Hours

Fic. 21-13. BIi penetration; successive immersion of same specimen. Tempera-
ture = 550°C, pressure = 250 psi, outgassed 550°C for 20 hr.

a mean value is found for investigations extended to as much as 3500 hr.
Evidently these variances are caused by outgassing by the graphite over
the time interval of the experiment. Outgassing the graphite at 900°C in-
stead of at 550°C reduced the amplitude of the excursions, but the mean
value remained from 0.425 to 0.525 g bismuth per cc graphite for most of
the types investigated. When the outgassing temperature was increased
to 900°C, the saturation or maximum value of uptake was reached in some
cases within 2.5 hr.

The rate of bismuth penetration into graphite was determined in order
to estimate the effect of an unexpected pressure excursion in the reactor.
Samples were subjected to 250 psi for times varying from 5 sec to 5 min,
as shown in Fig. 21-14. This time span far exceeds that expected for a
reactor pressure surge. The test conditions were 250 psi at 550°C after an
outgassing period of 20 hr at 550°C. The data indicate that the practical
maximum uptake is reached in about 10 sec for all graphites except types
A and G. These graphites, which are essentially coatings instead of bulk
impregnations, have their uptake increased continuously with time. In
a long-term test their equilibrium values were not reached until after some
800 hr of submersion.

Since the core of the reactor will be subjected to various pressures, a
study was made of the effect of pressure on the absorption of bismuth by
graphite. Long-term tests covering hundreds of hours were conducted at

Bismuth Uptake, grams /cc graphite
21-6] GRAPHITE 785

 

{ | T TTT] T lllllllll Tevel T 7

 

e

 

 

 

Bismuth Penetration, in Grams Bi/cc Graphite

 

 

 

0 | I c1at | oo laaar [y
1 : 10 : 100 500
Time of Immersion, in Seconds

 

Fic. 21-14. Short-time Bi penetration. Temperature = 550°C, He pressure = 250
psi, outgassed 550°C for 20 hr.

125 psi to duplicate the test discussed above. It was found that the bismuth
uptake is approximately the same as for the 250-psi pressure and the rela-
tive absorption remain the same between the different grades of graphite.

In another series of tests, samples were immersed for 20 hr at 550°C
at varying pressures, as shown in Fig. 21-15. The samples for each curve
were first evacuated for 20 hr at 550°C before being immersed in bismuth.
With each type of graphite, the bismuth uptake at 450 psi corresponds ap-
proximately to the values attained at 250 psi for longer periods of sub-
mersion. Type R, CCN, HLM, and ATL-82 are insensitive to pressure
increases beyond 200 psi. The remaining three types, A, G, EY-9, do in-
crease continuously in bismuth uptake and furthermore show a threshold
pressure below which no bismuth penetrates the graphite for the 20-hr
duration of the test.

However, results of long term tests at 125 psi showed that graphites
having a threshold pressure at 20 hr do absorb bismuth after several
hundred hours.

After a pressure surge in the reactor core, the operating pressure will
return to approximately 120 psi, and the amount of bismuth in the graphite
might decrease. To investigate this, samples impregnated at 450 psi were
resubmerged in bismuth at 25 and at 100 psi to determine what quantity
of bismuth might leave the graphite. The dotted lines in Fig. 21-15 con-
nect these points. It can be seen that there is no significant reduction of the
bismuth contained in each type of sample.
786 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

 

  
 
  

   

 

 
 
    

 

 

 

 

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0 100 200 300 400 500

Helium Pressure in psi

Fic. 21-15. Effect of pressure on Bi penetration; successive immersions of 20
hr. Temperature = 550°C, outgassed 550°C for 20 hr. Dashed lines for reduced
pressure after 450-psi impregnation.

Calculations of the percent of voids filled with bismuth were made for
the maximum bismuth uptake obtained in the experiments shown .in
Fig. 21-15. Table 21-9 gives these calculated values for the approximate
saturation level reached. In this table, the last graphite, AGOT, is the
conventional reactor graphite. The 1009, filling of the voids is obviously a
good check of the assumption that it is quite permeable. All the other
graphites are impermeable grades under development by various comi-
panies. The percent voids filled for these graphites do not represent total
saturation of the accessible voids. Rather, these values show that about
1/3 to 1/2 of the accessible void volumes have been filled in these experi-
ments.

In studying threshold penetration effect, surface tensions of bismuth on
various surfaces of graphite were measured (Table 21-10). Although dif-
fering from the accepted values, these determinations probably represent
more closely the actual circumstances in a reactor core. In none of the four
cases was wetting of the graphite obtained by the bismuth or bismuth
solution.

Uranium diffusion into bismuth in graphite pores. Since a certain amount
of fuel absorption will have to be tolerated with the graphites now avail-
able, 1t is essential to measure the diffusion of uranium into graphite by
21-6] GRAPHITE 787

TABLE 21-9

Voips FiLLED AT 450 psi AFTER 20 HR.

 

 

 

 

 

 

 

 

 

 

 

Graphite type Voids filled with Bi, 9

100 37

EY-9 44

A 17

G 37

HLM 32

M 23

R 32

AGOT 100

TABLE 21-10
SURFACE TENSIONS
Graphite surface Constituents Time afte.r Wettm.g Surface tension,
contact, min | properties dynes/cm

Smooth Bi 1 None 276
29 257
78 241
Rough and loose | Bi 5 None 153
particles 60 142
Polished Bi + 350 ppm Mg 15 None 66
90 66
Polished Bi+ 350 ppm Zr 15 None 285
120 275
240 282
360 283

 

 

 

 

 

 
788 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

means of the bismuth solution. Experiments to measure this effect have
been made. This was done by first impregnating graphite with bismuth so-
lution containing magnesium and/or zirconium. After bismuth impregna-
tion at a given pressure, uranium was added to the solution and the graph-
ite allowed to soak in the bismuth solution for a period of time. The amount
of uranium which diffused into the graphite was measured by sectioning
the graphite and analyzing for uranium concentration as a function of
distance from the surface of the sample. These experiments were run at
550°C with a pressure of 200 psi. The graphite was first allowed to soak
in the bismuth solution for 90 hr; then the uranium was added and the
conditions were maintained for the duration of the experiment. Results of
two experiments are given in Table 21-11. In the first experiment, the
bismuth contained 390 ppm Mg and 1000 ppm U. The uranium concen-
tration in the graphite specimen was found to be less than in the melt
solution and decreased from the sample face inwards.

The second experiment was performed exactly like the first except that
no magnesium was present. The graphite specimen (Great Lakes Type
HLM) absorbs less bismuth than the EY-9 graphite used in the first ex-
periment. However, the uranium concentration near the surface of the
specimen built up to an amount considerably greater than that initially in
the solution, and the concentration gradient is much steeper than was
found when magnesium was present in the solution. This high value for
the uranium-to-bismuth ratio near the interface may be explained by
assuming that uranium reacted with impurities present on the graphite
surfaces. Apparently when magnesium is present in the solution it reacts
preferentially with these impurities.

These experiments definitely show that uranium and other solutes
present in the bismuth can be expected to diffuse into the graphite as far
as the bismuth has penetrated. For the graphites now at hand, this means
diffusion through the entire thickness of the graphite for the long-term
exposures contemplated in a reactor core. Of course, since the diffusion of
uranium itself takes considerable time, fission will convert it to other prod-
ucts before it has an opportunity to diffuse many inches into the graphite.
The effect of diffusion of the various solutes and fuel into graphite on
neutron economy and reactor operational characteristics is recognized, and
studies have to be made in large-scale experiments.

In general, it is believed that the graphites at hand will meet the require-
ments for the first experiment of an LMFR reactor. It is already possible
to produce some of these in sizes as large as 40 to 60 in. in diameter. As
this development progresses, graphites of greater impermeability will be
produced. Improvements in graphite have taken place steadily, and
markedly improved materials are anticipated in the future.
21-6]

GRAPHITE

TaBLE 21-11

UraNiuM DIFrusioN INTO GRAPHITE

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Specimen no. Distance, in. Bi, 9, Mg, ppm U, ppm
A. EY-9 Graphite
1 0.0312 32.2 1150 900
2 0.0937 30.5 1050 840
3 0.1562 29.6 750 810
4 0.250 28.0 800 760
5 0.500 26.0 820 740
B. HLM Graphite
1 0.0312 16.35 5600
2 0.0937 16.87 2630
3 0.1562 16.38 460
4 0.250 16.78 70
5 0.500 17.78 30

 

 

789
790 MATERIALS OF CONSTRUCTION—METALLURGY [cHAP. 21

REFERENCES

1. P. Duwez and F. OpELL, Phase Relationships in the Binary Systems of
Nitrides and Carbides of Zirconium, Columbium, Titanium, and Vanadium,
J. Electrochem. Soc. 97, 299-304 (1950).

9. D. T. Keating and O. F. KamMeRER, Film Thickness Determination from
Substrate X-ray Reflections, Rev. Sci. Instr. 29, 34 (1958).

3. L. S. DARkEN et al., Solubility of Nitrogen in Gamma Iron and the Effect
of Alloying Constituents—Aluminum Nitride Precipitation, J. Metals 3, 1174~
1179 (1951).

4. J. R. Weeks and D. H. Gurinsky, Solid Metal-Liquid Metal Reactions in
Bismuth and Sodium, in ASM Symposium on Liquid Metals and Solidification,
ed. by B. Chalmers. Cleveland, Ohio: The American Society for Metals, 1958.

5. O. F. KAMMERER et al., Zirconium and Titanium Inhibit Corrosion and
Mass Transfer of Steels by Liquid Heavy Metals, Trans. Met. Soc. AIME 212,
20-25 (1958).

6. G. W. HorsLey and J. T. Maskrey, The Corrosion of 2Y4% Cr—1% Mo
Steel by Liquid Bismuth, Report AERE M/R-2343, Great Britain, Atomic
Energy Research Establishment, 1957.

7. J. R. WEEKs et al., Corrosion Problems with Bismuth-Uranium Fuels, in
Proceedings of the International Conference on the Peaceful Uses of Atomic Energy,
Vol. 9. New York: United Nations, 1956 (P/118, pp. 341-355); D. H. GURINSKY
and G. J. Diexes (Eds.), Nuclear Fuels. Princeton, N. J.: D. Van Nostrand
Co., Inc., 1956. (Chap. XIII); J. R. WeEeks, Metallurgical Studies on Liquid
Bismuth and Bismuth Alloys for Reactor Fuels or Coolants, in Progress in
Nuclear Energy, Series IV, Technology and Engineering, Vol. I. New York:
Pergamon Press, 1956. (pp. 378-408)

8. W. C. Lesuie and M. G. FonTaNA, Mechanism of the Rapid Oxidation of
High Temperature, High Strength Alloys Containing Molybdenum, Trans.
Am. Soc. Metals 41, 1213 (1949).

9. L. S. Marks (Ed.), Mechanical Engineers Handbook. 4th ed. New York:
McGraw-Hill Book Company, Inc., 1941. (p. 232)

10. W. E. MARKERT, JR. personal communication to J. R. Weeks, Mar. 20,
1958.

11. M. W. MaLLeTT et al., The Uranium-Carbon System, USAEC Report
AECD-3226, Battelle Memorial Institute, 1951; The Reactor Handbook, Vol. 3,
General Properties of Materials, USAEC Report AECD-3647, 1955. (p. 316)

12. W. E. MiLLEr and J. R. WEEKs, Reacttons between LMF R Fuel and Its
Container Materials, USAEC Report BNL-2913, Brookhaven National Labora-
tory, 1956.

13. G. V. Samsonov and N. S. RoziNova, Some Physicochemical Properties
of Zirconium-Carbon Alloys, Izvest. Sektora Fiz.-Khim. Anal. Inst. Obshchet.
Neorg. Khim. Akad. Nauk. S.8.S.R. 27, 126-132 (1956).

14. W. P. EATHERLY et al., Phystcal Properties of New Graphite Materials for
Special Nuclear Applications, paper prepared for the Second International
Conference on the Peaceful Uses of Atomic Energy, Geneva, 1958.
CHAPTER 22
CHEMICAL PROCESSING*

22-1. INTRODUCTION

The Liquid Metal Fuel Reactor offers the opportunity for continuous
removal of fission products from the fluid fuel by chemical and physical
processing. By this procedure the poisoning effect of the fission'products
may be kept to a low level, and thus make possible a good breeding ratio
in this thermal reactor. In this chapter, the various chemical and physical
processes for removing the fission products are discussed.

To simplify the discussion, the fission products are classified into four
basic groups as follows:

(1) Gaseous elements or compounds that are volatile at reactor operat-
ing temperature. This group is ordinarily abbreviated FPV.

(2) Nonvolatile elements forming compounds more stable than the cor-
responding uranium compound. The abbreviation for this group is FPS.

(3) Nonvolatile elements forming compounds that are less stable than
the corresponding uranium compound and more stable than the correspond-
ing bismuth compound. The abbreviation for this group is FPN.

(4) Nonvolatile elements forming compounds less stable than the cor-
responding bismuth compounds. The abbreviation for this group is NFPN.

In the FPV group there are four elements: bromine, iodine, krypton,
and xenon. Of these, 6.7-hr I'35 and its daughter 9.13-hr Xe135 are the
important ones. Xel35 1s by far the most important because of its cross
section, 2,700,000 barns. Since this is so large, it is necessary to remove
most of the iodine and xenon as soon as formed.

The other major poisons occur in the FPS group. In calculating the
average atomic weight and cross section of these groups, it is convenient
to use the fission yield in milliatoms. Normally, it is assumed that two
atoms of fission products are produced by the splitting of one atom of
uranium. Thus, 2000 milliatoms of fission products are produced by fission
of one atom of uranium, and 1% yield is equal to 10 milliatoms. On this
basis, Table 22-1 presents the FPS nuclides with the important informa-
tion on their poisoning effect. As can be seen, Sm4? is the most important
element to be dealt with in this group.

The last group, commonly called the noble fission products, represents
a combination of groups (3) and (4) in the above classification. The im-
portant poisoning information on all these nuclides is given in Table 22-2

*Based on contributions by O. E. Dwyer, A. M. Eshaya, F. B. Hill, R. H. Wiswall,
W. S. Ginell, and J. J. Egan of the Brookhaven National Laboratory

791
TABLE 22-1

FuseEp-SALT SOoLUBLE FissioNn Propucts [1]

Precursors have half-lives less than 5 days.

 

 

 

 

 

 

 

 

 

 

 

 

 

Fission Cross sec-
Nuclide Half-life yield y, |[tion o, barns Yo Type
milliatoms*|(at0.025 ev) t poison]
Rb?85 Stable 20 0.90 18 3
Rb86 19d 36 1.0 36 3
Rb87 6.2 X 100y 46 0.14 6.4 3
Sr88 Stable 54 0.005 0.25 3
Sr89 54d 61 110 6,700 2
Sr0 20y 64 1.0 64.0 3
Y —Zr°! 61d; (stable) 66 1.52 100.0 3
Xe—(Cs!33 5.27d 66 29.0 1,920 3
(stable)
(Csl35 3 X 108y 70.5 15.0 1,060 3
Csl37 37y 71.5 2.0 143 3
| Bal38 Stable 71.1 0.6 43 3
Lal39 Stable 70.5 8.4 590 3
Ba—La—Cel40 12.8d; 40h 68.5 0.63 43 3
(stable)
Ce—Pri4 32d (stable) 61.5 11.2 688 3
Cel42 Stable 55.0 1.8 99 3
Pr—Ndi43 13.5d 45.5 290.0 13,200 2
(stable)
Ce —»Pr—Ndl44 280d; 17m 36.0 4.8 173 3
(stable)
Nd145 Stable 27.0 52.0 1,400 2
Ndl46 Stable 20.0 9.8 196 3
Nd —»Pm—Sm!47 11.6d; 2.6y 14.0 60.0 840 2
(stable)
Ndu48 Stable 10.0 3.3 33 3
Sm!149 Stable 7.0 47,000 329,000 1
Nd150 Stable 5.0 2.9 14.5 3
Sm13! 73y 2.6 7,200 18,700 1
Sm152 Stable 1.6 150 240 2
Eu?s3 Stable 0.9 420 378 2
Sin154 Stable 0.5 5.5 2.8 3
Eulss 1.7y 0.3 13,000 3,900 1
Eu—Gd15 15d (stable) 0.2 750 150 2
Gdrs7 Stable 0.1 {160,000 16,000 1
Total 30 nuclides 1052.3 394,010

 

*Percent yield multiplied by 10; total yield is 200%, or 2000 milliatoms.
foavg = 374 barns.
1o > 1000 = type 1; o 50 to 1000 = type 2; o < 50 = type 3.

 
 

 

 

22-1] INTRODUCTION 793
TABLE 22-2
Fusep-SavT INsoLuBLE Fisston Propucts [1]
Fission Cross sec- T
Nuclide Half-life yield y, |tion o, barns Yo ype
milliatoms [(at0.025ev) T poson

Se’? Stable 0.4 40 16 2
Se’8 Stable 1.1 0.4 4.4 2
Se” 6 X 10y 2.0

Se80 Stable 2.8 0.53 1.5 2
Se82 Stable 5.5 0.055 0.3 2
7192 Stable 67.5 0.25 17 2
Zr93 5 X 108y 68.0 3 204 2
Zr9%4 Stable 67.5 0.08 5.4 2
Zr—NDb? 65d; 37d 66.0 13.4 880 2
796 Stable 64.0 0.05 3.2 2
Mo? Stable 59.0 2.10 124 2
Mo98 Stable 56.0 0.13 7.2 2
T 2.1 X 105y 48 .0 100 4 800 1
Mo'00 Stable 35.0 0.2 7.0 2
Ru!! Stable 26.0 12 312 2
Ru!02 Stable 24 .0 1.2 29 2
Ru!l03 40d 8.8 150 1,320 1
Ru'%¢ N.I.* Stable 6.2 0.7 4 2
Pd1os Stable 4.6 18 83 2
Ru106 1.0y 3.3 (15)% (50) (2)
Pdi07 5 X 108y 2.2 750 1,650 1
Pdio8 Stable 1.3 11.1 14 2
Agl0® Stable 0.9 84 75 2
Pdio N.1. Stable 0.4 0.4 0.2 2
Cdit Stable 0.3 750 225 1
Cdl12 Stable 0.2 0.03 0.01 2
Cdiis Stable 0.2 25,000 5,000 1
Sb123 Stable 0.2 3.86 0.76 2
Snl24¢ N.1. Stable 0.4 0.2 0.08 2
Sn —>Sb—Te!2% |10d; 2.7y (stable) 0.6 1.5 0.90 2
Tel26 Stable 0.9 0.8 0.72 2
Tel28 Stable 5.7 0.16 0.91 2
Tel30 Stable 25.0 0.31 7.8 2

Total 33 654.0 ’ 14,843 .38

 

 

 

 

 

 

 

 

*N.I., not identified as fission product on G.E. Chart, 1952.
T0avg = 22.7 barns.
JAssumed from values for daughter, Pd108,
794 CHEMICAL PROCESSING [cHAP. 22

under the group heading FPN. As can be seen by examining the column
headed yo, none of these nuclides is a very important poison, compared
with xenon and samarium.

From the data given in these tables, it is possible to calculate the poison
level in an LMFR as a function of time of operation. Besides the charac-
teristics of the fission products themselves, the poison level is dependent
mainly on the core fuel volume, the total fuel system volume, and the
average core flux. In Fig. 22-1, the poison level is given as a function of
days of operation for a 500-Mw LMFR reference design [1] with 600 ppm
U233 in Bi. It is assumed that the volatile poisons, FPV, can be removed
in a steady-state operation and the poisoning level kept to 19%. The other
two classes, of course, steadily increase, based on the assumption of no
chemical processing of the core. After a certain poisoning level is reached,
the continuous chemical processing will serve to keep the poisoning at a
constant value. This level must be chosen by a careful economic optimiza-
tion procedure.

Figure 22-1 shows that while the FPS group is the most important, as-
suming that the volatiles can be removed as desired, the FPN group does
gradually accumulate, and after about 400 days of operation has a 1%
poisoning effect. Hence, over long-term operation, processing of all the
groups becomes desirable if a low poison level is to be maintained.

The poisoning in a U235-fueled reactor is expected to be 10 to 20%
higher than in a U233-fueled reactor [2,3] depending on the average resi-
dence time of the fission products in the fuel. This is due to a shift in the
fission product spectrum toward higher cross section nuclides. The cu-
mulative poisoning effect of the higher uranium isotopes is also slightly
higher for U23°,

In connection with this last point, the higher isotopes of uranium grad-
ually build up throughout the operation of the reactor. In the calculations
used in the reference design of Chapter 24 and in BAW-2 [1], the poison-
ing effect of the higher uranium isotopes has been assumed as 2% for a
U233 fuel. Since these higher isotopes are chemically the same as the fuel,
no provision can be made for a chemical separation from the U233. The
gradual buildup of the higher uranium isotope poisons can actually be
tolerated over a number of years before becoming important in the eco-
nomics of the reactor operation, as is shown in Chapter 24.

In all the foregoing discussions, it is assumed that corrosion products
contribute very little to the poisoning in the reactor. However, this may
not be so. As was described in Chapters 20 and 21, the corrosion rate of the
containing metals by the bismuth fuel is rather high. Corrosion products
such as iron and chromium at a concentration of 300 ppm in bismuth would
contribute a poisoning effect of about 19;,. However, the same processes
which remove the FPS and FPN will also remove all the corrosion products.
22-2] VOLATILE FISSION PRODUCT REMOVAL 795

 

15 | l [ I | | 1 | |

 

 

 

 

 

?
:‘ |—
e L
©
a.
5 l—
e /
| // FPN "
/ -‘/-- —
b / o¢/
~ /.o/ FPV (A d) )
- ssume
---—--.# -------- L X 3 -= ------ - ey
ol T | | 1 1 1 1 1 1
200 400 600 800 1000 1200 1400 1600 1800 2000

Time of Operation, Days

Fic. 22-1. Poison level after startup vs. time of operation for all fission products.
Core fuel volume, 1800 ft3.

22-2. VoraTiLE FissioN Propuct REMovaLn [20]

22-2.1 Xenon and iodine removal. For a 19, poisoning level, assuming
no Xe adsorbed on, or absorbed by, the graphite moderator, the concen-
trations of 9.13-hr Xe!35 and total Xe in the fuel are calculated to be 1.5 and
12.9 ppb, respectively. Compared with the 9.13-hr Xe!3%, the combined
poisoning effect of all the other FPV’s is negligible, so that the problem
of FPV removal is really one of Xe!3% removal. Some typical statistics on
the FPV’s are summarized in Table 22-3. These figures are based on three
assumptions: (a) that Xe buildup on the graphite is negligible, (b) that
negligible amounts of Br and I are volatilized with the FPV’s, and (c) that
Kr and Xe have the same removal characteristics.

In Article 20-3.3 it was shown that the actual solubility of xenon in
bismuth may well be in the ppb range; McMillan calculated the solubility
as 10-12. Since the amount of xenon generated is probably larger than its
solubility in bismuth, it is necessary to determine the behavior of the gas
in relation to the surfaces of the reactor core and fuel conduits, as it will
have a strong tendency to escape from solution.

Since the xenon is the decay daughter of I35, it is born not only in the
reactor core but throughout the fuel system wherever I'3° is present.
Therefore the chemical and kinetic behavior of I, its decay precursor, 1is
important. The Xe!3% removal problem might be solved by desorp-
tion of I!35: however, it is found that the I'3° decays so rapidly that
at least 759, of the I'35 would have to be removed with the FPV’s in
796 CHEMICAL PROCESSING [cHAP. 22

TABLE 22-3
STATISTICS ON FPV’s uUNDER CONDITIONS OF
19, REAacTOR PoisoNING FOR A 500-Mw REacTOR
1000 ppm U233; 150 tons of Bi

 

1. Concentrations, ppb
(a) Kr 2.8
(b) 9.13-hr Xe!35 1.46
(c) Total Xe 12.9
(d) Total FPV’s 15.7
2. Removal rates, g/day
(a) Kr 23.1
(b) 9.13-hr Xe!35 12.0
(c) Total Xe 106.0
(d) Total FPV’s 129.1
3. Per cent, by weight, total fission products 23.8
4. Average atomic weight of FPV’s 122.3
5. Rate of radiant energy release, kw/g 605

 

 

 

order to significantly reduce the amount of Xe!35 formed. This is probably
too much to be hoped for. Experimental results indicate that such a large
fraction of the I cannot be volatilized from U-Bi fuel. Thermodynamic
analysis indicates that the I, for the most part, should react with the Rb,
Sr, Cs, and Ba fission products to form monoiodides with about 709, of the
I going to Csl.

These alkali and alkaline-earth iodides would presumably have low solu-
bilities in Bi and, as a result, have a tendency to leave the U-Bi fuel and
collect on unwetted solid surfaces. These iodides also transfer heavily to
the salt in the FPS-removal process, but the rate of processing would be
too slow to extract significant quantities of I'3% and, in fact, most of the
other iodine nuclides. Thus there appear to be two predominant modes
by which I departs from the fuel: physical expulsion in the form of iodides
and radioactive decay.

22-2.2 Xenon and iodine adsorption on graphite and steel. Graphite is
not wet by the fuel; moreover, it has a void volume of almost 209, largely
composed of interconnected cells. These facts suggest the possibility of Xe
buildup in an LMFR core.

A factor in this problem is the behavior of iodine in the LMFR fuel.
The iodine may form rather insoluble iodides, then adsorb on unwetted
surfaces, and there decay to Xe. Both kinetic and thermodynamic analyses
indicate that this may be a real possibility.
22-2] VOLATILE FISSION PRODUCT REMOVAL 797

In 1956, an in-pile loop [4] was operated at Brookhaven in which fission
products were generated in U-Bi fuel, where the natural U concentration
was 800 ppm. The concentrations of fission products were therefore
several orders of magnitude below those for an LMFR. Two steel rods,
1/2 in. in diameter and 4 in. long, were suspended vertically in the gas
space of the surge tank, 2 in. above the liquid metal level. One was exposed
for a period of 60 hr and showed an I'32 concentration of 9.0 X 107 atoms/
em? at time of removal; the other, exposed for 85 hr, showed 1.6 X 107
atoms/cm2. The corresponding 133 concentration in the flowing metal
was 1.1 X 10° atoms/cm3, which means that for every 100 atoms of I133
per cc of fuel there were roughly 1 to 8 I'33 atoms/cm? of exposed surface
in the gas space. The temperatures of the rods and liquid metal were the
same, 500°C.

Several steel tabs immersed for extended periods in the flowing metal
showed I133 concentrations on their surfaces roughly 100 times those found
on the rods suspended in the gas phase. Moreover, it was estimated that
less than half the I in the system was in the Bi; about 609, was found on
the container walls contacting the Bi and about 19, on the gas walls. The
tabs were, for the most part, unwetted by the Bi.

The loop had a degassing chamber in which the metal flowed in a thin
layer over a baffled plate. Samples of gas taken from this chamber showed
I concentrations too small to measure, even radiochemically.

To get a better understanding of this general problem, a two-part ex-
perimental program is underway at BNL. In the first part, capsule scale
experiments are being carried out to determine the action of iodine and
xenon on graphite and steel capsules containing U-Bi fuel. These capsules
are irradiated in the BNL pile and then examined for iodine, xenon, and
radioactivity across the radius of the specimen. The second part of the
program is a kinetic study of the removal of iodine and xenon in degassing
equipment.

In-pile capsule experiments. In one series of experiments, capsules made of
2107 Cr-19,, Mo steel and graphite were filled with Bi containing 500 to
1000 ppm of natural U, 350 ppm Mg, and 350 ppm Zr. The capsules were
degassed under vacuum for 3 hr at 800°C before being filled. They had the
dimensions 1.27 cm ID, 1.60 cm OD, and 10 em long. The capsules were
irradiated in a flux of 2 X 10!'2/(cm?)(sec), with the U-B1 mixture frozen,
for periods up to 2 wk. After irradiation, the capsules were held at 500°C
for periods ranging from 10 min to 117 hr. They were then cooled quickly
~ to room temperature and sectioned into 10 disks for radiochemical analyss.
The concentrations of Xe!33 133 and U were measured at the center of
each disk and in a 1-mm ring on the periphery of the Bi. The results are
summarized in Table 22—4.

These experiments are exploratory. They were carried out to determine
 

 

 

 

798 CHEMICAL PROCESSING [cHAP. 22
TABLE 224
REsurts oF IN-PILE STUDIES ON THE
Benavior oF IopiINE AND XENON IN LMFR FukL
Concentration Concentration .
of 1133 of Xel33, Agitated

Sample Contaiper atoms/g Bi atoms/g Bi during
number | material equilibration

Core | Periphery| Core | Periphery time
S—010 steel 5x 108 | 5x 101 | 7x 107 | 6 X 101 No
S—020 ” 7xX 108 [ 2x 101 | 2x 10 | 7 x 10U »
G010 graphite | 2 X 100 | 6 x 101 — — ”
G-020 »” 2x10° | 1x 101 | 2x 10° | 2 x 1010 ”
G-030 » 4x10% | 2x10°9 [ 1Xx10° | 2x 10° ?
G-040 » 3x10° [ 3x 10" | 1x 10 | 4 x 101 »
G-080 ” 5x 1010 | 3 x 101! | 2Xx 10° | 7 x 10° Yes
G-150 ” 7X 100 [ 6 x 101 | 1 X 10° | 7 X 10° »

 

 

 

 

 

 

 

 

 

roughly the extent to which iodine and xenon concentrate on interfaces.
However, in spite of the limitations of the experiments, the following con-
clusions are warranted.

When the concentration of iodine generated by fission reaches a level of
about 10! to 10'2 atoms/g Bi (capsules S—-010, S-020, G-010, G-020), the
iodine concentrates at the interface between the Bi and the container wall.
The concentration at the interface is about 1000 times higher than that in
the bulk of the Bi for the steel capsules, and about 100 times higher than
that for the graphite capsules.

When the concentration of Xe reaches a level of about 10! to 102
atoms/g Bi (capsules S-010, S-020, G-010, G-020), its concentration near
the Bi-steel interface is about 10,000 times that in the Bi. This ratio for
graphite, G, is only 10 (G-020). The difference between the steel and graph-
ite capsules is believed to be due to the fact that Xe diffuses into the latter.
This penetration by fission-product gases has been found in other experi-
ments and confirmed by autoradiographs and material balances.

When the concentrations of iodine and Xe are lower, i.e., about 10°
atoms/g, the differences between interface and core concentrations are
much smaller, though still statistically significant (Xe in G-040, iodine
in G-030 and G-040). For iodine the concentration ratios vary slightly
from less than 10 for G-030 to 100 for G-040. For Xe the ratio is only about
3 for G-040, and no significant separation was observed in G-030. These
22-2] VOLATILE FISSION PRODUCT REMOVAL 799

 

F1g. 22-2. In-pile capsule experiment with molten bismuth fuel, showing xenon
and iodine diffusion into graphite. |

lower Xe ratios are again attributed to the loss of Xe from the interface to
the graphite.

Samples G-080 and G-150 were agitated (by rotating them at 15 rpm
around an axis passing at right angles through the middle of the capsule)
while being equilibrated at 500°C for 75 hr. It is seen that in the case of the
agitated samples Xe segregation was unaffected but I separation was
appreciably reduced. However, the great bulk of the I was still found on
the outer layer of the Bi.

Besides these experiments, another series was carried out in which the
bismuth, containing uranium, was molten during irradiation, so that the
xenon and iodine had a chance to escape as soon as formed. Figure 22-2
is an example of a typical experiment. In the figure, the central dark area
is the bismuth core. The bright band is that part of the graphite into
which xenon and iodine have diffused at 500°C. This band is about 1.5 mm.,
since the picture represents a magnification of 4 times. The conditions for
this particular experiment are given in Table 22-5. The irregularities
observed in the photograph are in accordance with the heterogeneity of
graphite.

It should be noted that fission products other than iodine and xenon
may be and possibly are involved in the formation of the high-intensity
800 CHEMICAL PROCESSING [cHAP. 22

TABLE 22-5
CaprsurLe Trsts witHh MoLTEN FUEL

1000 ppm U235 in Bi 4 350 ppm Mg + 350 ppm Zr. Irradiated for 15 days
at a flux 2 X 1012 n/(cm?)-(sec) at 500°C. Graphite G capsule

 

Xe!33 concentration in graphite about 1 X 10'3 atoms/g of graphite

Xel33 ” ” bulk ” 3 X 10° atoms/g of Bi
J131 » ” graphite 7’ 5 X 10'3 atoms/g of graphite
I131 ” ” bulk ? 3 X 1019 atoms/g of Bi

 

 

 

regions. The penetrations in the graphite appear to be due to radioactive
gases exclusively.

The results of all these experiments show that I and Xe concentrate
very heavily on surfaces in contact with the U-Bi fuel. There is evidence
that Xe and radioactive gases penetrate the graphite and are immobilized
therein. This may present a very serious problem in keeping the LMFR
fission-product poisoning to the low levels required for economic breeding.
The reported experiments, however, have been limited by the available
neutron flux of the BNL pile to concentration levels about 1/1000 those
anticipated in an LMFR breeder. Extrapolation of the present results to
the LMFR levels is not justified, since it is conceivable that because of
saturation effects the concentrations at the interfaces may not increase
proportionately. However, the penetration of Xe in the graphite, as con-
trasted to its accumulation at interfaces, is a potentially serious problem
because of the large surfaces available inside the graphite.

The results of these experiments clearly indicate that the removal of the
FPV’s is not a simple degassing operation. An increased research program
is under way to learn more about the release and movement of the FPV’s
in both the reactor core and in the fuel streams. While degassing equipment
designed to afford a large fluid surface for escape of the gases will probably
be the best kind of equipment, the volatiles may very well never arrive at
the degasser at all. Instead, they may adhere to the graphite walls and to
the steel walls. Operation of the LMFR Experiment No. I should give
extremely valuable information on this particular question.

22-2.3 Design of equipment for FPV removal. In the LMFR, the fuel
would flow continuously through several parallel loops to external heat
exchangers for cooling. Degassing equipment would, in all probability, be
located in each of these loops. For a 500-Mw reactor, if all heat-exchange
22-3] FUSED CHLORIDE SALT PROCESS 801

streams were processed continuously, the fraction of FPV in the fuel re-
moved per pass would only be about 0.004. Since the solubilities of Kr
and Xe in Bi increase with temperature, the degassing equipment should
preferably be located in the coldest part of the system, but since the fuel
flow through the reactor is upward, and since the degassers must be located
at the top of the system because of hydrostatic pressure, it 1s not very
practical to locate them at the coldest point.

The main objective would be to prevent excessive amounts of Xe from
being adsorbed on, or absorbed in, the graphite moderator. To achieve
this, two conditions are necessary: first, the relative amount of I settling
on the graphite must be kept low, and second, the degassers must be very
efficient. The problem is not so much one of desorbing Xe from a Bi solu-
tion as it is one of controlling the accumulation of I and Xe on unwetted
surfaces. To minimize I buildup on the graphite, the fuel velocity in the
core should be as high as practical and there should be solid surfaces
located somewhere between the core and the degassers to collect I.

On the basis of present knowledge, the degassers should be so designed
that a large interfacial area is provided and that the liquid metal surface
is as turbulent as possible. Theoretically, a degasser should work with good
efficiency. A theoretical analysis by McMillan (BNL-353) showed that
xenon has a tremendous tendency to concentrate on liquid bismuth sur-
faces. For a spherical volume, the number of xenon atoms on the surface
was estimated to be about 10® times the number dissolved in bismuth at
300°C. At 500°C this ratio came close to 10°.

A sieve-plate column, in which the fuel descends in fine streams, would
be such a degasser. It is felt that sparging of an inert gas into the fuel is
not necessary to promote gas desorption, since Xe is so insoluble. However,
depending on the gas pressure in the degasser, the use of an inert carrier gas
may be desirable. The effluent fission gases would be collected in refriger-
ated charcoal beds.

22-3. FuseEp CHLORIDE SALT PROCESS

In processing the molten bismuth for the removal of fission-product
poisons, the ideal process would be a pyrometallurgical one operating at
substantially the same temperature as the fuel. Furthermore, this process
should either leave the uranium fuel in the bismuth or treat it in such a
manner that it is relatively easy to recharge it as a metal into the bismuth
stream for reuse. The LMFR thus offers an excellent opportunity for the
application of pyrometallurgical chemical reprocessing methods. From a
procedural point of view, such methods should inherently be cheaper than
presently known aqueous processing methods. It will be necessary, however,
802 CHEMICAL PROCESSING [cHAP. 22

to await an economic comparison of the aqueous and pyrometallurgical
processes before one is finally chosen for use with an LMFR.

However, since the LMFR offers such an excellent opportunity for the
application of cheap pyrometallurgical processing, this path has been
explored quite extensively. In this section a fused chloride salt process
for the removal of fission poisons is described. In following sections a
fluoride volatility process and a noble fission product removal process are
described.

22-3.1 Equilibrium distribution. Chemistry. The FPS group consists of
the lanthanides and the elements in groups IA, ITA, and ITTA of the Periodic
Table. Within this group the lanthanides account for about 94% of the
total poisoning effect of the FPS elements. In the case of a typical 500-Mw
reactor [1] the concentration of FPS elements in the bismuth amounts to
about 17 ppm. To reduce this concentration to acceptable levels, a process
has been developed whereby the FPS elements are oxidized by and then
extracted into a fused salt.

Following the original suggestion by Winsche that fission products
might be extractable from a liquid U-Bi fuel by molten salts in a manner
similar to solvent extraction, experiments were conducted by Bareis using
the LiCI-KCl eutectic and lanthanide-bismuth alloys [6]. If the mechan-
1sm was indeed one of liquid-liquid extraction, then the lanthanide distribu-
tion should follow a simple distribution law and as such be independent of
total concentration. Experimentally, this was not the case, and it was
subsequently shown by Wiswall [7,8] and later independently by Cubic-
ciotti [9] that the results could be explained by assuming that a chemical
reaction had occurred as follows:

3LiClsarty + La(siy <== LaCl3za) + 3Lis;. (22-1)
From the free energies of formation of the halides involved (Table 22-6)
we may calculate AF* =+ 33.6 kcal for Eq. (22-1). From this and the
relationship AF®= —RT In Kq, the equilibrium constant, Ke, is found
to be 3.2 X 10710, Obviously, the equilibrium will be displaced far to the
left. However, if we assume an initial La concentration in the bismuth
equal to 17 ppm, equal volumes of eutectic (KCl considered here as inert)
and bismuth, and that activities are equal to mole fractions, then the ratio
of moles of lanthanum in the salt to moles of lanthanum in the bismuth at
equilibrium will be 146. Essentially, therefore, all the lanthanum will be
transferred to the salt phase.

On the other hand, for the analogous reaction with uranium:

3LiCl + U === UCl; 4 3Li (22-2)
22-3] FUSED CHLORIDE SALT PROCESS 803

TABLE 22-6

AF or CERTAIN Havmpes aT 773°K [10]

 

 

 

Free energy of formation F,
Compound keal/atom Cl
KCl 88.6
SmCI:g 84 .1
LiCl 82.6
NaCl 81.4
LaClg 71.4
CeCl3 69.8
NdCl; 67.4
MgClz 61.7
UCl; 57.5

 

 

 

 

the standard free energy change is +75.3 kcal, and Keq = 5.2 X 10722,
At equilibrium, assuming the initial uranium concentration in the bis-
muth = 1000 ppm, the ratio of the mole fraction of U in salt to the mole
fraction of U in Bi will be equal to 6.8 X 1074, Thus, in principle, a selec-
tive oxidation of the lanthanides may be achieved in the presence of
uranium. Of course, the assumption that activities are equal to mole
fractions is only an approximation.

Ternary salt. As a consequence of these reactions, lithium metal builds
up in the bismuth phase and, in view of its high thermal neutron cross
section, replacement of the lanthanide by lithium offers no advantage m
terms of neutron economy.

Therefore another low-melting salt, the ternary eutectic of MgClz(50
mole 9,), KCl (20%), and NaCl (30%) (MP 396°C) was investigated.
In this system, the free energy of formation of MgClz is intermediate
between those of the lanthanide chlorides on one hand and uranium
trichloride on the other and a satisfactory, although not complete, separa-
tion should be achieved.* Furthermore, the low neutron cross section of
Mg is more favorable than that of lithium, and a low concentration of Mg
in the fuel (250 ppm) appears to be necessary in order to minimize cor-
rosion and mass transfer in the steel equipment. The magnesium con-
centration in the bismuth will therefore control the extent of the reaction:

*The stability of NaCl and KCl is so much greater than that of MgCl; that their
contribution to the oxidizing potential of the salt may be neglected. However, they
do exert an influence upon the activity coefficient of MgCl..
804 CHEMICAL PROCESSING [cHAP. 22

3Mg012(3a1t) + 2La(Bi) 4__._)'_ 2La(313(sa1t) -|— 3Mg(Bi), AF0 = —9H8.2 kcal,
(22-3)

3MgC12(sa1t) + 2U(Bi) .(_—"'2 2UCI3(39,11;) + 3Mg(]3i), AFC = -|—25.2 kcal,
(22-4)

but will not influence the degree of separation which may be achieved.

Thermodynamacs of FPS transfer and distribution data. The equilibrium
constant for reaction (22-3), in which lanthanum is taken as being repre-
sentative of lanthanides in the 43 oxidation state, is given by

2 3
ALaCl; AMg

Keg = . 22—
7 afa afrecn, (22-5)
Expressed in terms of mole fractions, Eq. (22-5) becomes
Keq — X%aCls X13VIg . (ffi:.Cls)z (f;\.jfg>3. (22_6)

X%,a. Xi;VIgClz (ff’a)2 (fMgClz)3

In the above, a = thermodynamic activity, X = mole fraction, and f and
f® are activity coeflicients. f* is the limiting activity coefficient at infinite
dilution, which is assumed to be independent of concentration at the
concentrations encountered in this investigation. It is equivalent to the
Henry’s law constant [11].

Solved for the experimentally detesminable quantity Xrp.ci/X1a, Eq.
(22-6) becomes

 

XLa.Cl3 _ Keq X%’IgC]z 1/2 _
Xt ‘( K, Xi, ) ! (22-7)
“where
_ (ffa013)2 (f;\o’lg)?’.

T (20)? (f mecn)?

In logarithmic form, (22-7) may be written

X als e 2 3
logTLIS—:—glog XMg—I—%log Ig—q—(—j%%gl)—, (22-8)

whereupon, a plot of log Xraci,/X1a versus log X, should result in a
straight line of slope = —3/2. Figure 22-3 is a plot for most of the FPS,
~ uranium, and zirconium based on the best experimental data. In the
case of La, the best line has a slope of —3/2. From the position of the line,
the constant term of Eq. (22-8) may be calculated by
22-3] FUSED CHLORIDE SALT PROCESS 805

emf, Volts vs Zn/ZnCly (2) //

 

—0.40 —0.50 =0.001
]000 .1 T 1 l vV T l T 01 ' T T 3 1 -
- Sm La -
500 |— —
200 | -
100 —

50

  
 

! lllllll

f 111111

T
c
wn
3

20

| llllll]

] | lnul

k, (wt. % insalt) / (wt. % in bismuth)
o

 

 

 

2 \Zr
_ N\ 4
N\
1 N\ -
= D -
- \ -
0.5 \ \ —
L \ \ —
o \\ Zr  _
0.2} -
\\U
0.1} N\, = L
= l G Fia. 22-3. Distribution of solutes
00501 liie I Lol 17 - - 1-
s Sl 0\ = between MgCle-NaCl-KCl and Bi-Mg.

Mg in Bi, ppm

K eq (X MgClz) 3 .

B (constant) = % log K
s

(22-9)

Experimental values of these constants are given in Table 22-7.
Comparison of theory and experiment. In order to compare theory with
experiment, Ke,, Xmgcl,, and the activity coefficients of the pertinent
substances in each phase must be known. K., is easily calculated from
the AF>® for the appropriate reaction by means of the relation AF®
= —RT In Keq; Xmec1, may be considered essentially constant and equal
to 0.5, since MgCls is present in the salt phase in large excess over the other
reactants and its concentration changes only very slightly during the
reaction. An exact calculation of K, is not possible at this time, owing to
the paucity of information regarding activity coefficients in fused salts
and in liquid bismuth. However, in one case, that of cerium, it is possible
to estimate K, from measured activity coefficients if one assumption is
allowed. Recently Egan [12,15] has measured the partial molar free energy
of mixing, AF, of magnesium in bismuth and cerium in bismuth by galvanic
cell methods. From AF, and AFc., it was possible to calculate fy, and
f&, the activity coefficients at infinite dilution, in bismuth at 500°C.
These values are estimated to be f¥=2X 1073 and f& =3 X 1071

 

 
[cHAP. 22

CHEMICAL PROCESSING

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(LNVISNO))) g Jd0 SHATVA

-GG 14V T,

 
22-3] FUSED CHLORIDE SALT PROCESS 807

(Table 22-8). Neil [13] by similar galvanic cell techniques, has meas-
ured the activity coefficient of MgClz in the ternary salt eutectic
MgClo-KCI-NaCl at 500°. The best value to date is fmgci, = 0.34. If it
is assumed that fGeci, = 0.1 in the ternary salt (and this value appears
reasonable), then K, for cerium is given by

_ (J&cw)?(fRe)® _ (1071H)*(2 X 1073)°

T (f&)2(fmeor,)® (B3 X 10714)2(0.34)3 = 2.3 X 108,

K;

 

 

 

The experimental value of K;, 5.6 X 1020, leads to a value of 2 X 1071°
for f&ci,. The agreement is considered satisfactory, in view of the ex-
ponential character of the equations and the uncertainties in the available
data.

For example, the entire difference between the experimental value and
calculated value of K; may be reconciled if one assumes an error of
1.4 kcal/atom Cl in the AF of formation of CeCls. Such an error is well
within the limits with which the standard free energies of formation are
known at these temperatures. The estimated activity coefficients of
metals in bismuth may also be in error by as ymuch as a factor of 2 to 3.

The experimental values of the constant Bi. and Bng (Table 22-7)
may be used to calculate the activity coefficients of lanthanum and
neodymium in the bismuth if it is assumed, as in the case of cerium, that
ffac, = fR4ac, = 0.1. The values so obtained, ff.=4 X 107 and fRq
=2 X 10713, are quite low, and are in general agreement with the
measured f2..

In the case of samarium, SmCly is thermodynamically more stable
than SmCl; by 14.6 kcal/atom of Cl at 500°C, and hence the equilibrium
reaction is

Sm iy + MgClaeay === SmClasarty + Mgsi)- (22-10)

In a manner analogous to the treatment of the trivalent lanthanides, we
obtain

log 3_(;&(:_1_2_ = — log Xmg + B/, (22-11)
Sm
where
B =1 Keq XMgClz and .K} — féomClz ffig.

f OSOm f MgCl2
808 CHEMICAL PROCESSING [cHAP. 22

TABLE 22-8

Activity COEFFICIENTS AT
INFINITE DIiLuTiON

 

 

 

System Temperature, °C far

Ce-Bi 500 3 X 10714
Mg-Bi 500 2 X 10-3
U-Bi 500 1x10°°
Li-Bi 450 1X 1075
Na—-Bi 500 8.5x 1073
Zr-Bi 700 7% 104

 

 

 

 

 

Equation (22-11) predicts that a plot of log Xsmcl,/Xsm versus log Xmg
should yield a straight line of slope —1. The curve is shown in Fig. 22-3
and the line is drawn with a slope of —1. This line yields the value of B’
given in Table 22-7. With the assumption that f&.c1, = 0.1, the estimated
activity coefficient at infinite dilution of samarium in bismuth is
f&m=3X 10718,

The validity of Eq. (22-6) is dependent upon the assumption that side
reactions, such as the oxidation of bismuth by the salt, are negligible.
Since AF? for these reactions are large positive numbers, it is reasonable
to consider bismuth as inert in this respect. Bismuth, of course, interacts
with the lanthanides and magnesium very strongly, but this is taken into
account by the use of activity coefficients.

It is also assumed that the reactions

3NaCl 4 La === LaCls + 3Na, AF9 = +4-30.0 keal,

3KCl+ La === LaCl;s 4 3K, AF0 = 451.6 keal,

do not contribute significantly to the transfer of lanthanides to the salt
phase, in view of the large positive free energy change. This approximation
was checked experimentally by determining the concentrations of Na and
K in the bismuth phase after an equilibration experiment. No detectable
amounts of alkali metals were found in the bismuth. This result also
indicates that salt solubility in bismuth is negligibly low. Analysis of the
salt phase for bismuth yielded low, erratic results, possibly due to the slight
solubility of bismuth in 1 N HCl which occurred during the aqueous
separation of salt and metallic phases. It is highly improbable that bismuth
would be soluble in a salt of this type.
22-3] FUSED CHLORIDE SALT PROCESS 809

Another assumption made in this analysis involves the reversibility of
the oxidation of the lanthanides by MgCls. This point was checked by
equilibrating a series of Mg-Bi alloys with a salt eutectic containing
Cel43 Cls. The distribution data of cerium as a function of Mg concentration
in the bismuth derived from the reduction of CeCl; by Mg shows the
reaction is reversible within an experimental error of 109 [14].

Included in Fig. 22-3 are Nd distribution data obtained in the presence
of 0.1 w/o uranium and 0.03 w/o zirconium in the bismuth. (Zirconium
will normally be present in the LMFR fuel as a corrosion inhibitor.)
Within this concentration range, zirconium and uranium do not affect
the Nd distribution.

Data for process design. It will be noted from the relative positions of
the lines of Fig. 22-3 that it is not possible to assume, a priors, that the
order of the lanthanide distributions will be directly predictable from free
energy of formation data. For example, from Table 22-6 the order of
decreasing stability of the chlorides is Sm, La, Ce, and Nd, whereas at
constant Xy, the experimental order is La, Sm, Nd, and Ce. The differ-
ence in order is apparently due to the large variation of the activity coeffi-
cients of the lanthanides in bismuth.

The results of the lanthanide distribution experiments have provided a
basis for the design of a countercurrent, salt-metal extraction process [2].
Results of uranium distribution studies indicate that in small-scale exper-
iments, a satisfactory separation of lanthanides from uranium may be
achieved in a single equilibrium contacting stage. The experimental value
of the distribution coefficient, K, was found to be of the order of 20 to 50,
where K is defined as

X LaClj / X Lapg;
Xuct,/ Xug;

Multistage extraction should ensure efficient removal of the fission-product
poisons from the bismuth fuel stream.

22-3.2 Pilot plant equilibrium experiments. A pilot plant equilibrium
program is under way at BNL to investigate the salt bismuth-fuel equilib-
ria on a larger scale under conditions more closely simulating those in an
actual plant. The contacting vessels, made of 347 stainless steel, have a
capacity of about 2 liters. They can be fitted with liners of other metals
in order to study the effect of surfaces and corrosion. Each contacter 1s
equipped with connections through which materials can be added and re-
moved without admitting air, a sightport, gas and vacuum connections,
heaters, and thermocouples. Liquid salt and metal phases are equilibrated
in quantities large enough to allow multiple analyses, so that the effect of
810 CHEMICAL PROCESSING [cHAP. 22

changes in conditions can be directly determined by before-and-after ang]-
yses on a single system.

The most significant results of this pilot plant program are those from
experiments for which an apparatus of large capacity alone could serve.
These are studies of the stability of the solutions for long periods and of
the changes in equilibrium distribution resulting from addition of various
reagents. In general, the distribution coefficients obtained in this equip-
ment confirm those found in the small-scale work. However, the precision
of the results is less.

In carrying out experiments in these equilibrium vessels, a stability suf-
ficient for most practical purposes can be achieved, given the right operat-
ing conditions, but there are still unsolved problems. A solution of Bj,
U, Mg, rare earth, and Zr can be kept at 500°C under helium in a stainless-
steel vessel indefinitely without change of composition. If then a quantity
of pretreated salt* is added to the system, a significant drop occurs in the
U concentration in the metal, e.g., from 1000 to 900 ppm. Some U appears
in the salt phase, but not in an amount equivalent to the loss from the
metal. Thereafter, the U concentration remains constant but the Mg in
the metal suffers a slow decline, losing between 1 and 10 ppm per day.
As its concentration decreases, the distribution of elements such as the
rare earths changes in about the way which would be predicted from the
results of the gram-scale experiments. The U remains nearly constant
unless the Mg is allowed to drop below about 20 ppm, in which case U
begins to transfer to the salt.

From some of these systems a solid material has been recovered which
gives the x-ray pattern of uranium nitride, and it is possible that nitrogen
from the container walls is somehow involved in the mysterious behavior
of U and Mg. Since very small quantities of the various materials are in-
volved in these reactions, it is quite possible that solid surface adsorption
effects are also playing a part in the instability of composition.

In a second series of experiments, the change in equilibrium distribution
from the addition of reagents is being studied. In the FPS extraction proc-
ess, a sequence of columns operated at different oxidation potentials is
proposed. Most of these changes in equilibrium are controlled by the
addition of BiCls or Mg to the system at appropriate points. Experiments
have been done in which these reagents have been added to a salt-metal
system at equilibrium. The results of two such experiments will illustrate
the behavior of these systems. In the first, an initial equilibrium was es-
tablished in which the metal phase contained a fairly high concentration

*Molten salt which has been equilibrated for many days with molten Bi con-
taining high concentrations of Mg and U. When each phase has reached constant
composition and there is no U in the salt, it is ready for use.
22-3] FUSED CHLORIDE SALT PROCESS 811

of Mg. Its approximate value, together with those of the other constitu-
ents, are given in the first row of Table 22-9, Run 1. In view of difliculties
in sampling and analysis, these figures may be in error by 10 to 20%.
A quantity of BiCls which was more than equivalent to all the Mg was
then added. When a new equilibrium had been reached, it was found that
all the Mg and much of the Zr and U had been removed from the metal
phase; the second row gives the analytical figures. Metallic Mg was then
added to reverse the reaction; and the final equilibrium situation is given
by the figures in the last row. It can be seen that the U was restored to its
original concentration in the metal. It is of interest that this occurred
even though the final Mg concentration was much less than the original;
that is, the U distribution coefficient was insensitive to the Mg concentra-
tion when the latter’s value was 100 ppm or more. This agrees with the
laboratory experiments discussed previously. Additional confirmation was
obtained in Run 2, in which an amount of BiCls was added which was less
than equivalent to the Mg. The results are given in Table 22-9, Run 2.
Here, when the Mg concentration was lowered from 320 to 140 ppm, the
Ce distribution coefficient increased, as one would expect, but the U re-
mained unchanged.

22-3.3 Reaction rates. Previously, the equilibrium for the salt-metal
reactions was discussed. It was shown that most probably more than one
equilibrium contact will be required to remove the FPS. This means that
some kind of contacting between two flowing streams will be required in

 

 

 

 

 

TABLE 22-9
Concen-
tration, Concentration, ppm
mole 9,
Zr U Ce
Mg | Mg
Salt | Metal
1" salt | Metal | Salt | Metal | Salt | Metal
Run No. 1
Initial equilibrium 50 440 | 20 | 240 10| 800 | 15 | 11
After BiCl; addition 50 10 | 20 | 160 |1070{ 330 | 73 0.1
After Mg addition 50 110 | 20 | 210 17| 810 | 56 4
Run No. 2
Initial equilibrium 50 320 | — | 200 20 790 | 28 9
After BiCls addition 50 140 | — | — 30 790 | 58 5

 

 

 

 

 

 

 

 

 

 

 
812 CHEMICAL PROCESSING [cHAP. 22

the over-all chemical processing, using the fused chloride salt method.
Therefore, an examination of the reaction rates is important. When sev-
eral equilibrium contacting stages are required, and it is desired to do this
in a flowing countercurrent system, it is necessary for the mass transfer
rates to be fast.

In this reaction there are at least three stages: transport of the reactants
to the salt-metal interface, the reaction proper, involving exchange of elec-
trons, and transport of the products away from the interface. Situations
are conceivable in which any of these could be rate-limiting. In investi-
gating so complex a situation experimentally, it is often possible to order
things so that one or more stages are fast relative to the others, thus per-
mitting the kinetics of the latter to be studied alone. If, for example, the
reaction is made to go under conditions which are far from equilibrium,
i.e., the reverse reaction proceeding to only a negligible extent, the trans-
port kinetics of certain species can be excluded from consideration.

A series of experiments of this type was carried out in which 2.2-mm drops
of Bi containing 200 ppm Sm!%3 fell through 31 em of molten ternary salt
eutectic. At the bottom, the drops were drawn off and analyzed. The
salt phase was initially free of rare earths, and its volume was 500 times
that of the total Bi which fell through, so transport of species in the salt
phase should not be rate-limiting. The contact time for each drop was
about 0.6 sec. Analysis showed that 759 of the Sm was extracted into the
salt. If we calculate the amount that would have been extracted had the
rate been limited by diffusion of solute to the surface of a spherical drop,
assuming rapid reaction at the interface, a smaller figure results. It may
be concluded that some turbulence exists within the drop, assisting the
diffusion process, and that the interface reaction is indeed fast.

Although further rate studies are required, the results at hand show that
considerable latitude is available to the process engineer in designing the
over-all process using these equilibrium and rate data. These possible de-
signs may range from straight batch type contacting to completely auto.
mated countercurrent contacting.

22-3.4 FPS removal process. In the process design described [20], the
oxidant is BiCls and the carrier salt is the ternary eutectic NaCl-KCl-
MgCls,  which melts at 396°C. Sufficient oxidant is added to the salt to
remove the FPS, leaving the U for the most part behind. The FPS form
chlorides, which are considerably more stable than UCl3, the most stable
chloride of U.

Equilibrium partition coeflicients for Ce, Zr, and U, as functions of Mg
concentration, are shown in Fig. 22-3. For a particular Mg concentration,
the ratio of the Ce coefficient to that of U is a direct measure of how diffi-
cult it is to achieve a given degree of separation between the two solutes.
22-3] FUSED CHLORIDE SALT PROCESS 813

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Exit Salt
2436 lbs/Day
FPS 220.5 FPS 215.0 (238 gm/Duy)
Zr 908 |/ Zr  537(595 " )
FPS 18.4 U 122.0 @ U 1.8( 20 " )
Fuel St Mg 300.0 > /
vel Stream 7, 2500 LT
From Reactor {y 1500.0 r_l__ FPS 16.7
— \ Mg 88.0 Mg 185
40600 |bs/day Zr 120.0
| @\‘ ~—— ‘ U 390.0
o 'i;::; ';Zeacl;olr 533 2 %7 ‘ L-—Mg
. U 655.0
Mg 12 |Ibs/Day Mg 159 Zr 247 © 2 .60% ® 63 gm/Day
Bi 2.6 Bi
A S
.81b/D
500 lbs/Day |  Bi10.4% 0.8lb/Day
BiCl3
I
93 Ibs/Day 750 |lbs/Day
U 966 /
inlet | B 1357 Ibs/Day Biors | )
Salt Processing
BiCl3 Ucl3 Bi Stream
2 lbs/Day 870 gms/Day Bi 7970 FPS O
Mg 0.7 Ib/Day MZg g
500 lbs/D r
C S ay o 1 [Ql = ‘ uo
[ +] i
<4 Ibs/Day

Note: Concentrations in Parts Per
106 Unless Otherwise Indicated

Fig. 22-4. Flowsheet for the removal of FPS and Zr fission products from an
LMFR fuel.

Ce is one of the least stable of the FPS chlorides, but in the treatment
which follows, FPS salt-metal equilibrium coefficients are taken to be the
same as that of Ce, that is, they are conservative. The slope of the Ce and
U lines in Fig. 22-3 is —1.5, signifying trivalency in the salt phase, while
that of the Zr is shown as —1, signifying divalency. The Zr line is drawn
dashed because experimental results are still preliminary, and the slope of
—1 was assumed rather than being firmly established by experiment. Un-
fortunately, the data available at this writing were obtained over a rather
short range of Mg concentration. Of the FPS, Rb and Cs are univalent
and Ba, Sr, and Sm are divalent, but all of these lie well above the Ce line
and would, therefore, be more easily extracted.

The total energy release per fission in the LMFR is estimated to be
194 Mev. For a reactor having a heat rate of 500 Mw, this means that
542 g of U235 would be fissioned per day. Since the FPS represent about
449, of the total fission products by weight, 238 g of FPS’s must be re-
moved per day to maintain a steady concentration in the fuel. (See Table
22-10.) The Zr concentration is kept at about 250 ppm for purposes of cor-
rosion inhibition, and the steady-state removal rate of this fission product
will be approximately 59 g/day. It is interesting to note that about 11%
of the fission products end up as Zr. For a reactor with a heat rate of
500 Mw and a total fuel inventory of 150 tons, a fission-product Zr con-
814

CHEMICAL PROCESSING

TABLE 22-10

[cHAP. 22

STATISTICS ON VARIOUS FissioNn-Propucr GROUPS

For a 500-Mw reactor having a 150-ton Bi inventory containing 1000 ppm U233,

 

 

 

 

: Approximate | Removal | Weight fractién of
Typical con- :
Group : reactor rate, total fission prod-
centrations .
poisoning, 9, | g/day ucts produced
FPS 18 ppm 0.8 238 0.44
Zr 250 ppm 0.1 59 0.11
FPN 2 ppm 0.8 0.6 0.0011
N FPN (less Mo) 174 ppm 59 0.110
Mo 1 ppm 0.0 54 0.10
FPV 16 ppb 1.0 129 0.24

 

 

 

 

 

 

centration of 250 ppm corresponds to a 590-day average residence time in
the fuel and gives a reactor poisoning effect of slightly less than 0.1%.
 Figure 224 is a simplified flowsheet showing how the FPS may be re-
moved from an LMFR fuel of a 500-Mw reactor. The high concentration
of Mg makes it difficult to extract the FPS, but the high concentration of
Zr makes it easier to extract that particular element. The high Mg con-
centration rules out the possibility of using a buffer method and necessi-
tates the use of a stoichiometric method in the FPS removal step. Suffi-
cient oxidizing agent (in this case BiCls) is added to the salt to remove
the required fractions of FPS and Zr. At the same time, most of the Mg
in the fuel is unavoiadably oxidized.

After a suitable holdup period, the fuel flows at the rate of 0.34 gpm
through column 1, the removal column. This column, as- shown, has a
separative capacity equivalent to two equilibrium stages, The separative
capacity of the column is illustrated in Fig. 22-5, where concentrations,
relative flow rates, and equilibrium partition coefficients are shown. The
bottom stage operates under oxidizing conditions, while the top one oper-
ates under reducing conditions. This brings about relatively high concen-
trations in the middle of the column. The increase in the U concentration
in the fuel, in passing through the column, is to provide the necessary U
makeup for the reactor. The principal effects of increasing the number
of stages to three would be to lower slightly the Mg concentration in the
exit fuel stream, to increase considerably the FPS/U ratio in the exit salt,
and to decrease appreciably the Zr/U ratio in the exit salt. The first of
these by itself would be of little consequence, the second would be very
22-3] FUSED CHLORIDE SALT PROCESS 815

 

 

 

 

 

 

Fuel Inlet FPS = 220.5 ppm
FPS = 18.4 ppm | 40578 2436 Zr — 908 “
Mg =300 * lb/Day Ib/Day U=122

Zr — 250 *
U= 1500 *
Stage 1
(Reduction)

kFps — 3.33

kZr = .295

kU =.0666

A

— FPS — 1012 ppm

l;:: ~ 593 "o | 40503 2451 | 7, — 1048
7r — 308 Ib/Day Ib/Day U= 5630 *“
U—1834 “
Stage 2
(Oxidation)
kFPS — 185
kzr = 4.27
ku =37

 

 

 

   
 

Inlet Salt Stream
FPS = 5.5 ppm
Zr =367 "
U=6.55 '
Bi = 2.69 %

FPS = 5.5 ppm

Mg =159 * | 40600
Zr — 247 * | lb/Day
U=1533 *“

 

Fic. 22-5. Typical concentrations in an FPS removal column with two equilib-
rium stages.

desirable, and the third would be undesirable. The last effect is actu-
ally controlling, which means that a three-stage separation is not as
good as a two-stage separation. Going in the other direction, a one-stage
separation gives very much lower FPS/U and Zr/U ratios in the exit salt,
thereby increasing the difficulty of subsequent U recovery. However,
certain advantages result from a single-stage operation—higher Mg con-
centration in the exit fuel allows easier control of the process, and higher
Zr concentration in the exit salt makes it easier to remove the Zr. The op-
timum number of equilibrium stages probably lies between one and two.
In column 2, the U in the salt stream from column 1 is recovered by ex-
tracting it into a second Bi stream. This column operates under the buffer
system, even though the Mg concentration in the metal stream drops 52%.
The separative capacity of this column is equivalent to four equilibrium
stages, and the variations of solute concentrations throughout the column
are shown in Fig. 22-6. The U losses in the exit salt stream were set ar-
bitrarily at 2 g/day, for purposes of illustration. Obviously, in actual prac-
tice this quantity would be determined by economic considerations, i.e.,
it would be at such a value that the cost per gram of recovering any ad-
ditional U would be more than it is worth. The process design of column 2
is controlled by the fact that the concentration of Zr in the exit salt stream
has to be 54 ppm for a salt flow rate of 2436 1b/day and a reactor with a
500-Mw heat rate, i.e., 59 g of fission-product Zr must be removed per
816 CHEMICAL PROCESSING [cHAP. 22

 

I I I | I I I I

400 - —
Equilibrium Curve

Operating Lline

 

 

300 —

Uranium-Recovery Column
with Four Equilibrium Stages

Uranium Concentration in Metal, ppm

 

 

 

 

 

 

 

 

 

. Concentration in Bismuth Stream, ppm
100 Solute Inlet Outlet —

FPS 0 16.7
Mg 185 88

Zr 0 120

U 0 390

1 | | | | ] | |
0 20 40 60 80 100 120 140 160 180

Uranium Concentration in Salt, ppm

Fig. 22-6. Uranium recovery column with four equilibrium stages.

day. When the Zr concentration in the exit salt stream is fixed, the con-
centration of the FPS is also fixed, for the ratio FPS/Zr in the exit salt
stream must be the same ratio in which these materials are generated in
the fuel. The concentration of the FPS in the inlet fuel stream to column 1
was varied while the Zr concentration was held constant, until this condi-
tion was achieved. The concentration of 18.4 ppm for the FPS’s corre-
sponds to a total FPS poisoning effect of about 0.8%.

The processing Bi stream, column 2, contains 185 ppm Mg but no other
solutes. In passing through the column, the Mg concentration in the Bi
drops to 88 ppm, which means that the oxidation reduction potential be-
tween the salt and Bi phases changes appreciably throughout the column.

The fission products, Mg and U in the Bi stream from column 2, are all
oxidized completely into incoming salt stream B in vessel 3. The stripped
Bi, after addition of 185 ppm Mg, is then returned to column 2 to repeat
its cycle. The Mg—Bi stream is so small that a few days’ supply could be
prepared on a batch basis if continuous addition of Mg to the recirculating
Bi stream proved difficult to control.

Vessel 3, conditions in which are highly oxidative, could be a short
column; its only function is to provide good single-stage contact between
the Bi and salt streams. The U makeup for the reactor, shown added as
22-3] FUSED CHLORIDE SALT PROCESS 817

UCIl; to the salt stream entering this vessel, is transferred to the fuel in
column 1. Alternatively, the bred U from the blanket could be transferred
from a Bi solution to the incoming salt. This Bi stream would be joined
to that from column 2 and later separated from it after leaving vessel 3,
or it could be contacted with the incoming salt in a separate vessel.

The exit salt from column 2 can be treated with a Ba—Bi or Ca—Bi solu-
tion to remove the FPS and U, thus making it possible to recirculate the
salt. The FPS’s and U could then be slagged out of the Bi into a low-cost
salt mixture for storage.

The flowsheet in Fig. 22-4, for the sake of simplicity, does not show
holdup and storage tanks, instrumentation, pumps, or heat exchangers.
There are several possible variations of this flowsheet but, for the most
part, they include the three types of operations described above.

Owing to the fact that the oxidation-reduction potential varies con-
siderably throughout the FPS removal column, it may be preferable to
operate it with concurrent flow within each stage and countercurrent flow
between stages. Alternatively, two separate concurrent columns could be
used. The U recovery column, on the other hand, would clearly be operated
with countercurrent flow because, chemically, conditions are reductive
throughout the column.

Design of extraction columns. The mechanical design of a proposed ex-
traction column is shown in Fig. 22-7. Fuel enters at the top of the column
and is dispersed by the slots in each tray as it falls through the column.
The flow paths are indicated by arrows. Coalescence of the fuel drops
occurs on each tray. Salt, as the continuous phase, may flow either con-
currently with or countercurrently to the fuel. Fuel coalescence promotes
thorough local mixing in the fuel and at the same time tends to minimize
axial dispersion in each phase.

Columns of the type shown in Fig. 22-7, about 3 to 6 ft long and 3 to 4 in.
in diameter, are expected to have satisfactory performance characteristics.
Such columns have not yet been tested under conditions simulating actual
practice, although their fluid dynamical behavior has been studied with
H,0 and Hg as substitutes for salt and Bi1.

22-3.5 Process control of fused chloride process. The object of the
process described above is to remove 59 g of fission-product Zr and 238 g
of FPS from the fuel per day, at the same time losing only 2 g of U. For
this, careful control of the process is required. Continuous measurement
of the U concentrations in the salt streams from columns 1 and 2 will be
required. The U concentrations in these streams are good indicators of
column operation, i.e., if the U concentrations are correct, those of the
7r and FPS should also be correct. Assuming constancy of fuel composition
and all flow rates, the two operating variables affecting the process are,
818 CHEMICAL PROCESSING [cHAP. 22

Gas & Vacuum System

 

 

 

 

 

Salt

     
  

Location
of Fuel-Salt
Interface

 

 

Fuel

F1c. 22-7. Extraction column.

first, the BiCls concentration in the inlet salt stream to column 1 and the
Mg concentration in the inlet Bi stream to column 2. Each of these must
be controlled to give the proper concentrations of U in the salt leaving
columns 1 and 2. The operation of column 1 is the more difficult to control.
There are three inlet salt streams which eventually merge into the single
stream entering column 1. Stream A contains about 929, of the total
BiCls requirements, B contains about 29, and C contains the remainder.
Streams A and B are separated because of difference in corrosiveness, and
stream C provides fine control of the total BiCls addition. At least one
day’s supply of each stream would be prepared in advance.

The Mg concentration in the exit fuel is a sensitive indication of the
rate of BiCls addition to the column and, consequently, of the U concen-
tration in the exit salt. Thus controlling the rate of addition of BiCls to
column 1 by this Mg concentration would be more satisfactory than con-
trolling it by the U concentration in the exit salt, because of the much
quicker response of the Mg concentration to changes in the rate of BiCls
addition. The damping effect of the column should then result in a fairly
uniform U concentration in the exit salt.
22-3] FUSED CHLORIDE SALT PROCESS 819

 

 

 

 

 

 

400
[ [ I 2000
U in Exit Fuel
300 1 > —3 1500
£ - Zr in Exit Fuel £
8 g
o o
g‘ c
2 200 —{ 1000 -2
S S
5 £ g
§ 3 in Ex; §
S 1t SO/f S
100 |- U in Exit 5qp 500
FPS X 100 Zr in Exit Saly
in Exit Fuel
8 10 20 30 40

Mg Concentration in Exit Fuel, ppm

F1ac. 22-8. Effect of Mg concentration in exit fuel on the compositions of the exit,
streams from the FPS removal column.

Figure 22-8 shows the effect of variation in the Mg concentration in the
exit fuel stream of column 1 on the steady-state concentrations of FPS,
Zr, and U in the exit fuel and salt streams. It is seen that changes in Mg
concentration have less effect the higher the Mg concentration; e.g., in the
case shown, the column would be much easier to control at an exit Mg con-
centration of 25 than at one of 15.

The results of studies at Brookhaven indicate that it should be possible
to measure continuously the Mg concentration in the exit fuel by means of
a galvanic cell. For this, Marsland [17] has used the following type of cell:

Zn /ZmCl2(19, solution in NaCl-KCIl-MgCl2 eut)pyrex//
NaCl-KCIl-MgClz eut/Mg (Bi).

The emf from such a cell would control the voltage to another, large elec-
trochemical cell. This second cell, shown in the flowsheet, would add
BiCl;s to inlet salt stream C, the rate depending on the demand from the
controlling cell.

The control of column 2 should be much less of a problem. The Mg con-
centration in the inlet Bi stream must be kept within certain limits to
maintain the desired concentration of U in the exit salt. Actually, column
2 can, to a considerable degree, correct for malfunctioning of column 1.

In the event that control of the process described in Fig. 22—4 should
turn out to be difficult, several steps which can be taken to correct the
difficulty: (a) decreasing the separative capacity of column 1, (b) increasing
820 | CHEMICAL PROCESSING [cHAP. 22

the salt flow rate, and (c¢) inserting a holdup tank between columns 1 and 2
to assure uniform composition of the salt entering column 2. As an ex-
treme measure, the first column could be replaced by equilibration vessels,
but this would appear to be an unlikely eventuality. The magnitude of
the problem is defined by the continuous processing requirements, namely,
maintaining close control of very low uranium and fission product concen-
trations in streams of three interdependent contacting towers.

22-3.6 Processing to reduce radiation hazard. The continuous process
described above is based on an FPS concentration of approximately 18 ppm,
and calls for a processing rate of 0.45 gpm. These conditions were chosen
on the basis of poisoning considerations. If, however, safety considerations
were the determining factor, the processing rate would be greatly increased.
If the whole fuel stream were processed daily for removal of FPS, the con-
centrations of Sr% and Cs!37, the two worst fission nuclides from the stand-
point of biological hazard, could each be kept down to about 0.1 ppm.
This might well be a very desirable objective, and the processing rate would
still be only about 3 gpm.

22-3.7 Pilot plant program for fused chloride process. Plans for an ex-
tensive pilot plant program for the fused chloride process are currently
being made. Some work on mechanical component development and ma-
terials of construction has already started. Several small loops are in opera-
tion at BNL, circulating fused salt. In these loops, mechanical components
such as pumps, valves, and control instruments are under development and
test. Concurrently, a corrosion test program is under way, as was discussed
in Section 21-5. A full-sized prototype pilot plant for the testing and
operation of a single extraction column is now being fabricated and con-
structed (Loop N). This pilot plant has been designed to circulate quantities
of bismuth fuel and fused salt comparable to those for a full-sized reactor,
as discussed previously in this chapter. In this pilot plant it is planned to
obtain operational data which will lead to a full-scale processing plant.

22-3.8 Heat generation by fission products. The problem of heat re-
moval is an important consideration in the design of processes and equip-
ment for handling radioactive fission products. This is particularly true in
the present case, because of the relatively short age of the fission products
at the time of their extraction from the fuel. However, heat removal from
fused salts does not present a difficult problem.

Figure 22-9 shows a family of curves giving the specific heat rates for
the FPS as a function of average residence time in the reactor and time
after removal therefrom [2,16]. The curves were calculated from fission-
22-4] FISSION PRODUCT FLUORIDE VOLATILITY PROCESS 821

 

 

 

 

105 l T T T I 3

= -

I~ —

LSS Day ~

_— —

] — a5 Days 3

~ = -~

0 - -

‘5 — -

3 — ]

v ~ .

a | ™S~ 20 Days -
M.
£

‘g ]03 E— -

2 e =

3 ~ T «< . 100 Days ~

E - —

o = -
S

o - -
X
o

21— |

Z‘; 10 E “

L N

10 | ] ] | |
1 10 102 103 104 105 106

Time After Removal From Reactor, sec

F1a. 22-9. Energy release from FPS.

product heat-release data obtained from the Argonne National Laboratory.
Extrapolations to short decay times were made with the aid of the Way-
Wigner expression for fission product decay heat.

The energy will be divided about equally between beta and gamma
radiation. For this fused chloride process, the heat release will, of course,
depend on the poison concentrations, but will probably range from 100,000
to 500,000 Btu/(hr)(ft3 of salt).

22-4. FLUORIDE VOLATILITY PRrROCESS FOR FissioN PRropucets

As an alternative to the fused chloride process, a pyrometallurgical
process based on the volatility of UF¢ has been suggested by the Argonne
National Laboratory. A schematic flowsheet is given in Chapter 24 as
Fig. 24-18. This process would be operated batchwise.

In this process a batch of molten salt made up of 509, ZrF4 and 509, NaF
is poured into the graphite hydrofluorinator and heated to 600°C. The
fused salt is then sparged with HF gas, dissolving approximately 3 w/o.
After the HF gas is cut off, bismuth-uranium core fluid containing FPS is
introduced at the top of the vessel. Salt-metal contacting time is long
enough to permit hydrofluorination of the FPS, FPN, and U in the core
fluid and subsequent extraction of the resulting fluorides into the fused
822 CHEMICAL PROCESSING [cHAP. 22

salt melt. Excess HF is sent to a scrub tower not shown in the figure. The
stripped bismuth is continuously withdrawn from the bottom of the
column into. a storage tank, leaving enough Bi for a liquid seal. The
fluoride salt containing all the poisons and UF4 is then routed to a nickel
fluorination vessel in which UF4 is fluorinated to UF¢ by direct contacting
with fluorine gas. Other salts, such as MoFs, TeFs, RuFg, AsF3, IFs
and MbF5 [18], are also formed in this step, since they are present as
fission poisons. Since all of these materials are volatile, they will leave
the fluoride melt with the excess fluorine, and will then be condensed in a
cold trap maintdined at approximately —40°C. An NalF trap removes
traces of Tel's, AsF5, and Rufs from the fluorine before it is recycled.
These volatile fluorides are then sublimed from the cold trap by heating
to about 100°C, and distilled in order to complete the separation and puri-
fication of UF¢ from the other volatile fluorides.

The gaseous UF¢ is reduced by bubbling it with an excess of hydrogen
through fresh molten fluoride salt. The resultant UF4 is trapped in this
clean salt melt. As shown in Fig. 24-18, the salt containing UI'4 is next
contacted with the stripped bismuth stream in an electrochemical reduction
step. In this step, the UF4 is reduced to metal at a flowing bismuth cathode
while fluorine gas is released at the anode. The resultant bismuth-uranium
alloy, to which Mg and Zr have previously been added, is ready for re-entry
to the core.

As an alternative to this last electrochemical step, the UF4 can be reduced
in the salt by direct contact with Mg-Bi.

Although the development work on this process is not as far advanced
as on the fused chloride process, enough work has been done so that the
process appears feasible. Small-scale laboratory work has indicated that
the hydrofluorination step can be.carried out successfully. Previous work
in other areas of the atomic energy program has supplied considerable
information on the direct fluorination step and the volatile fluoride dis-
tillation step. In the other areas of this process, less information is cur-
rently available.

The chief advantages of the fluoride Volatlhty process is that it will be
operated batchwise and will give a complete, clean separation between the
uranium and all the fission products. This allows comparatively easy
control of the cleanup of the fuel and preparation of new fuel for the reactor.
Since each step of this process is batch, the instrumentation would be
comparatively simple and the operators would have complete independent
control of each step.

On the other hand, there are many difficult problems being encountered
in developing this process further. One of them is the severe corrosion
encountered in the various steps. The chemistry of the hydrofluorination
in the first step has to be proven out conclusively. It is believed that by
22-5] NOBLE FISSION PRODUCT REMOVAL 823

close temperature control the oxidation of bismuth can be prevented.
However, the free energy of formation of bismuth fluoride is rather close
to those of some of the fission products. From an economic point of view,
some means will probably have to be found for cutting down the cost of
fluoride salts sent to waste. Zirconium fluoride is quite expensive and
could be an important item in the total expense of the program.

29-5. NosBLE FissioN Propuct REMOVAL

22-5.1 Characteristics of FPN poisoning. Owing to the fact that they
include no nuclides which are particularly high neutron absorbers, the FPN
can be allowed to build up to relatively high concentrations in the fuel.
The two worst poisons are Tc? with a 19-barn thermal cross section, and
stable Rh1%3 with 150. For the reactor conditions described earlier, the
poisoning effect of the FPN (less Mo) is essentially proportional to their
concentration or average residence time in the fuel. The relationship is

Percent poisoning = 0.0020 (average residence time in days).

Table 22-11 shows the concentrations of the FPN and NIFPN elements
after a 400-day operating period. It is seen that the FPN group represents
only about 19, of the total soluble FPN.

 

 

 

 

 

 

 

 

TABLE 22-11
FPN CoNCENTRATIONS AFTER 400 Days oF OPERATION
FPN group NFPN group
Element ppm Element ppm
Ag 0.21 Se 0.75
Cd 0.44 Nb 5.0
In 0.07 Tec 39.0
Sn 0.58 Mo* 1
Sb 0.42 Te 23.0
Ru 80.0
Total 1.72 Rh 17.0
Pd 9.2
Total 175.0

 

 

*Solubility of Mo is less than 1 ppm; if solubil-
ity had not been exceeded, its concentration would
be 146 ppm.
824 CHEMICAL PROCESSING [cHAP. 22
The FPN group, minus Mo, represents 11 a/o of the total fission products.
With practically all the Mo out of solution, a 400-day residence time gives
an FPN concentration of 177 ppm with a reactor poisoning effect of about
0.8% for a 500-Mw reactor. To maintain that concentration, the fuel
~would have to be processed at the rate of only 9.2 gal/day, assuming com-
plete removal of the FPN’s. The size of the batches, and therefore the
frequency of processing, would be determined by economic factors. Pro-
cessing would begin probably after 400 days of full-power operation.

22-5.2 Chemistry of NFPN removal by zinc drossing. The process
adopted for the NFPN fission products is basically the same as the Parkes
[19] process for desilvering Bi. Experiments conducted by the American
Smelting and Refining Co., under a research subcontract with the Brook-
haven National Laboratory, and by Argonne National Laboratory have
given very encouraging results. A few results are given in Tables 22-12
and 22-13 to illustrate the high efficiency of the process. In a series of
experiments, Ru, Pd, and Te were dissolved in Bi at 500°C. Zn was added
in three concentrations, 1, 2, and 3%. In each case, the mixture was
agitated and then cooled to 400°C. The concentrations of the original
solutes, both in the filtered Bi solution and in the skimmed-off dross, are
shown in Table 22-12.

TABLE 22-12

REMovaL or NFPN MEeTALs FrROM B1 witH ZN

 

 

 

 

 

 

 

 

 

 

 

 

Concentrations at 400°C, ppm
Metal Dross
Amount of Zn
dded,
added, 7 Ru Pd Te Ru Pd Te
0 15 62 25 18 64 357
1 3 22 1.5 324 1280 610
2 0.6 5.3 1.0 216 1038 320
3 <0.5 1.9 <0.1 187 847 213

 

 

As shown by the results in Table 22-13, the amount of Zn required de-
creases as the precipitation temperature is lowered. The less Zn added,
the less to be extracted later. |
22-5] NOBLE FISSION PRODUCT REMOVAL 825

TABLE 22-13

RemovaL oF NFPN MgeraLs FroM Br witH 0.59, ZN
AT VARIOUS TEMPERATURES

 

 

 

 

Concentrations in Bi, ppm
Temperature, °C
Ru Pd Rh Te
Original solution, 500 44 26 12 100
Zn added, 450 31 31 9.5 8
400 12 11 1.2 0.6
350 2.4 4 0.5 0.6
300 1.5 1.6 0.5 0.6
Freezing point 1 0.9 0.5 0.6

 

 

 

 

 

 

 

22-5.3 FPN removal for the fused chloride process. The zinc drossing
process has been modified for use with either the fused chloride or the
fluoride volatility process. In this article, the modification for the fused
chloride process is discussed, and in the ‘next, that for fluoride volatility
will be described. In both cases, the NFPN removal is essentially the same.

Flowsheet. The proposed process is represented in Fig. 22-10. From the
FPS-removal plant, the fuel is charged to vessel 1, which is an equilibration
tank having both agitation and heat-removal facilities. Here it is contacted
with ternary chloride salt and just enough BiCls to oxidize the FPS, Mg,
Zr, and U into the salt. The fuel stream is then fed into vessel 3, where
practically all the FPN fission products are removed from the Bi with
7Zn. The more noble fission products form intermetallic compounds with
Zn, which are skimmed off the top of the Bi after cooling it close to its
freezing point. Thus far, it is known that Se, Nb, Te, Ru, Rh, and Pd of the
NFPN group and Ag of the FPN group can be removed from Bi by Zn
treatment. It is a general observation that all elements more noble than
Bi are removable by Zn. The extents to which the FPN elements Cd, In,
Sn, and Sb and the NFPN element Tc are removed by Zn have not yet
been determined. It is probable that both In and Sn will not be appreciably
removed. :

The concentration of Zn required is less than 0.5%, which is well below

~ its solubility limit at 500°C. The Bi from vessel 3 1s charged to vessel 4,

where the residual Zn and FPN are removed by oxidizing them to chlorides
with ternary chloride salt containing BiCl3(Cl sparging could also be used).
The stripped Bi is then fed to vessel 2, where it is contacted with the salt
from vessel 1. Sufficient Mg is added to the Bi in the vessel to transfer all
826 | CHEMICAL PROCESSING [cHAP. 22

Salt

     
  

 

Fuel From FPS—Removal Plant

FPS

Mg
Zr
U
FPN

NFPN

Zr

Sludge to Waste

 
 
 
 
 
  

Fresh Ternary Salt

To Salt Recovery

 
 

FPN
In

Bi to Fuel System

F1a. 22-10. Flowsheet for the removal of FPN fission products from an LMFR
fuel.

the FPS, Zr, and U in the salt back into the metal and still leave about 300
ppm Mg in the Bi as it is returned to the reactor. In vessel 2, the Fe and
Cr concentrations in the Bi should be brought back up to those in the
incoming fuel.

A portion of the stripped Bi from vessel 4 may be sent directly to a
holdup tank and used for fuel-adjustment purposes. Similarly, a concen-
trated solution of U in Bi may be made in vessel 2, also for fuel adjustment
purposes.

Vessel 2 can be eliminated and its function taken over by vessel 1. The
two vessels were included in Fig. 22-10 for convenience in explaining the
process. All vessels are similar in design and equal in size. To handle 275
gal (one month’s accumulation) of metal, they should have a total volume of
about 350 gal. The operations in vessels 1 and 2 should be conducted in
O2-free atmospheres, but this condition is not necessary for the operations
carried out in vessels 3 and 4.

M olybdenum removal. Mo is really a special member of the NFPN group.
Its solubility at 375°C, probably the coldest fuel temperature, is estimated
to be less than 1 ppm. Moreover, it is produced at a rate equivalent to
0.38 ppm/day. Thus, probably the most feasible way to remove Mo would
be to precipitate it out of solution onto cold fingers immersed in the circu-
lating fuel. The precipitation rate for a 500-Mw reactor would be 54 g/day,
most of which would be stable Mo. Even with cold traps, some Mo will
22-5] NOBLE FISSION PRODUCT REMOVAL 827

very likely precipitate throughout the fuel system where it is generated.
Information on this will be obtained in the LMFR Experiment No. 1.

Polonium removal. The behavior of Po in the FPS and FPN removal
processes described above is not clear. Chemically, it is more noble than
Bi and should not form an intermetallic compound with Zn, indicating that
it should always remain with the Bi. In preliminary equilibration experi-
ments with chloride salt mixtures, it was found that about 1% of the Po
transferred to the salt, but whether this was due to chemical oxidation or
volatilization is not presently known.

Heat generation rates. The maximum rate of heat removal from the charge
in vessel 1 is estimated to be about 290 kw (250 from the fission products
and 40 from the Po) and from vessel 3 about 240 kw (200 from the fission
products and 40 from the Po). These values can be greatly reduced by
allowing the 275 gal of fuel to ‘‘cool off”” for several days before processing.
The rate of heat removal can be kept sufficiently low so that cooling the
vessels does not present too much of a problem.

The worst heat-removal problem arises when the NFPN’s are con-
centrated in the Zn, Bi-NFPN sludge; but the generated heat can be re-
moved satisfactorily by leaving the intermetallic sludge in contact with
some molten Bi in the “‘extraction’’ vessel for a short while until it can be
skimmed off and sent to waste without danger of excessive heating.

22-5.4 FPN removal process for the fluoride volatility process. The
flowsheet for this proposed process is given in Fig. 24-20. The feed stream
for the FPN fission product removal plant is taken from the fluoride vola-
tility plant after the bismuth is free of all the uranium and FPS. This bis-
muth stream now contains only FPN. In a batch vessel, it is brought in
contact with a small amount of zinc (approximately 0.5 w/0). As the con-
tents are cooled, the zinc forms intermetallics with the FPN and NFPN
elements, as described previously. This zinc dross floats to the top, is
skimmed off and sent to the zinc waste. From the bottom of the vessel,
the bismuth stream containing some zinc is sent to a zinc crystallizer, where
the temperature is further decreased. Some of the zinc crystallizes and 1s
removed from the top for recycle back to the first vessel. It is proposed
to remove the remaining zinc by a distillation operation shown on the
flowsheet as Still. The bismuth from the Still is ready for return to the
volatility plant for the addition of uranium, magnesium, and zirconium.

The entire FPN plant would be operated batchwise. The quantity of
material handled would probably be about the same as for the previous
process, about 275 gal. The heating problem for this process also would be
about the same as for the process described previously.
828 CHEMICAL PROCESSING [cHAP. 22

 

Blanket
400 Tons
10% Th—90% Bi
=1 187.5 ppm U
61 ppm Pa Th 890g/Day
550°C
5.52 Tons/Day

 

 

U233 500g/Day

   

 

 
  
 

A 5.52 Tons/Day]
Bismuth S .
178 ppm Th torage tor

+ PP Decay of Pa To U

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 
 

Removal of
5.52 Tons/Day |111.04 Tons/Day Thorium and [ Wastes

10% Th—90% Bi|} 5% Th—95% B; Fission
104 ppm U 93.8 ppm U Products
55.7 ppm Pa 30.5 ppm Pa 4
550°C -

A Thon.um

Uranium

Redistribution Protactinium

   
   
  
 

of Bred and Fission
Products and Product
Reconstitution Chlorides

of Slurry
Particles 5.52 Tons/Day

Bismuth
60 ppm Th A
+ | 83.5 ppm U

' 5.3 ppm Pa

 

 

 

 

 

 

 

 

   

eparation

 

 

 

 

¢
\
1

 

 

 

 

o | =T

 

 

Stripped Bi

-

KCl, Na CI, Mg Cl2, Bi CI3

 

Fia. 22-11. Flow diagram for processing a 10 w/o Th-Bi breeder-blanket slurry
to remove Pa233 and U233,

22-6. BLANKET CHEMICAL PROCESSING

As is pointed out in Chapter 20, the easiest blanket to handle in the
LMFR would be a 10 w/o thorium-bismuthide slurry in bismuth. Chem-
ical processing of this blanket would be very similar to the core processes
already described. The major problem consists in transferring the bred
uranium and protactinium from the solid thorium bismuthide to the liquid
bismuth phase, so that they can then be chemically processed. Two ex-
amples of proposed processes are shown in Fig. 22-11, which shows a proc-
ess that can be used with the fused chloride salt FPS removal process, and
in Fig. 24-19, which shows a flowsheet for a process to be used with the
fluoride volatility process. |

In the process of Fig. 22-11, a typical two-fluid 500-Mw LMFR would
have a blanket of about 400 tons of material containing approximately
109% Th and 909, Bi. The material balance shows that 5.52 tons/day
would be withdrawn and processed. In the first step, an additional 5.52
tons/day of bismuth containing fresh thorium is added to the stream,
22-7] ECONOMICS OF CHEMICAL PROCESSING 829

primarily to dilute the thorium bismuthide to half its first value. This slurry
is then raised in temperature until a complete solution is obtained. When
this is done, all the uranium and protactinium as well as fission products
dissolve in bismuth. In the next step, the thorium bismuthide is reconsti-
tuted by cooling methods such as described in Chapter 20. The U, Pa,
and fission products will remain in solution in the bismuth. Then a phase
separation at 350°C can be carried out. This gives a recycle stream of
5.52 tons/day containing 109, Th going back to the blanket and 5.52
tons/day of Bi with about 959, of the original U and Pa dissolved in it.

This stream then goes to column 1 of the fused chloride salt FPS re-
moval system, where all the Th, U, Pa, and FPS are transferred to the
ternary chloride salt. Meanwhile, the stripped Bi is returned to dilute
more blanket thorium bismuthide.

In the last step, the Pa is allowed to decay to U for about 130 days. At
the end of this time, practically all (99.59,) of the Pa would be converted
to U, and the U would be separated from the Th and FPS by the methods
previously described. As shown on the flowsheet, this would result in the
production of approximately 500 g/day of U233 for charging into the core
fluid. |

The flowsheet for the bismuthide slurry head-processing shown In
Fig. 24-19 shows a similar technique for handling the transfer of U and
Pa from the intermetallic solid to the liquid bismuth. |

As yet neither of these processes has been tried in the laboratory. As
work progresses on the bismuthide blanket system, further work on the
chemical processing will be carried out.

29-7. EconoMmics oF CHEMICAL PROCESSING

Basically, in evaluating the economics of chemical processing, the cost
of neutrons in the form of uranium fuel wasted to fission-product poison-
ing must be balanced against the cost of operating the chemical processing
units for removal of the fission-product poisons. In the over-all operation
of an LMFR reactor and its auxiliary chemical processing plant, the at-
tainment of the highest breeding ratio will not necessarily give the lowest
cost power. When the price of fuel is fixed as it is now by the U.S. Govern-
ment, or by general market conditions, then the cost of the chemical proc-
essing becomes the variable which must be operated upon in order to justify
a high breeding ratio.

As is shown in Chapter 24, the chemical processes now available and
under development are so expensive relative to the cost of fuel that op-
timization of the reactor conditions for a two-region reactor indicates a
most economic breeding ratio of about 0.86, and for a single region reactor
about 0.75. These figures can be increased toward one or better by de-
830 CHEMICAL PROCESSING [cHAP. 22

creasing the cost of chemical processing. However, it must be kept in mind
that the fission products are not the only neutron poisons present in the
LMFR. Such other neutron poisons as the bismuth, the structural ma-
terials, and the higher uranium isotopes will be present even though the
fission and corrosion products levels are kept to zero percent.

REFERENCES

1. Bascock & Wircox Co., Liquid Metal Fuel Reactor; Techmcal Feasibility
Report, USAEC Report BAW-2(Del ), June 30, 1955.

2. O. E. DwyER, A.I.Ch.E. Journal 2, 163-168 (June 1956).

3. J. B. Sampson et al., Poisoning tn Thermal Reactors Due to Stable Fission
Products, USAEC Report KAPL-1226, Knolls Atomic Power Laboratory, Oct. 4.
1954.

4. C. J. RasEMAN et al., Continuous Removal of Fission Products from Liquid
Metal Fuel, Chem. Eng. Progr. 53(2), 86-F (1957).

5. R. W. REBHOLZ et al., Chemical Reprocessing Studies for the Liquid Metal
Fuel Reactor Ezxperiment, USAEC Report UCN-42, Umon Carbide Nuclear
Co., June 28, 1957.

6 D. W. Bareis, A Conlinuous Fisston Product Separation Process. I. Re-
moval of the Rare Earths (Lanthanum, Cerium, Praseodymium, and Neodymium)
Jrom a Typical Liquid Bismuth—Uranium Reactor Fuel by Contact with Fused
LiCl-KCl M7xture, USAEC Report BNL-125, Brookhaven National Laboratory,
1951.

7. R. H. WiswaLL, Jr., The Distribution of Elements in Salt-Metal Systems
with Special Reference to the Data of D. W. Bareis, USAEC Report BNL-201,
Brookhaven National Laboratory, Sept. 30, 1952.

8. D. W. Bareis et al., Fused Salts for Removing Fission Products from U-Bi
Fuels, Nucleonics 12(7), 16-19 (1954).

9. D. Cusiccrorri, An Ezxplanation of the Effect of Added Metals on the Dis-
tribution of Rare Earths between Liquid Bismuth and KCl-LiCl, USAEC Report
NAA-SR-202, North American Aviation, Inc., 1953.

10. A. GrAssNER, The Thermochemical Properties of the Oxides, Fluorides, and
Chlorides to 2500°K, USAEC Report ANL-5750, Argonne National Laboratory,
1957.

11. O. KusascaEwskI and E. Evans, Metallurgical Thermochemastry. New
York: Pergamon Press, 1956. (pp. 42-43)

12. James J. Ecan and R. H. WiswavLy, Jr., Applying Thermodynamics to
Liquid-Metal-Fuel Reactor Technology, Nucleonics 15(7), 104-106 (1957).

13. D. NEiL, personal communication.

14. W. 8. GINELL, paper presented at San Francisco Meeting, American
Chemical Society, April 1958.
REFERENCES 831

15. R. H. WiswaLL, Jr., et al.,, Recent Advances in the Chemustry of Liquid
Metal Fuel Reactors, paper prepared for the Second International Conference
on the Peaceful Uses of Atomic Energy, Geneva, 1958.

16. O. E. Dwyer et al., High Temperature Processing Systems for Liquid
Metal Fuels and Breeder Blankets, in Proceedings of the International Conference
on the Peaceful Uses of Atomic Energy, Vol. 9. New York: United Nations, 1956.
(P/550, pp. 604-612)

17. D. B. MaRrsLaND, A Reference Electrode for Fused-Salt Studies, PH.D.
thesis, Cornell University, Ithaca, N. Y., 1958.

18. S. LaAwroski, Survey of Separations Processes, in Proceedings of the
Interndtional Conference on the Peaceful Uses of Atomic Energy, Vol. 9. New
York: United Nations, 1956. (P/823, pp. 575-582)

19. D. M. LipperL (Ed.), Handbook of Non-ferrous Metallurgy, Vol. II. New
York: McGraw-Hill Book Company, Inc., 1945. (p. 201)

20. O. E. Dwykgr, A. M. Esuava, and F. B. HiLL, Removal of Fission Products
from Urantum-Bismuth Fuel, paper prepared for the Second International Confer-
ence on the Peaceful Uses of Atomic Energy, Geneva, 1958.
CHAPTER 23

ENGINEERING DESIGN

23-1. ReEacTor DEsigN*

The LMFR readily lends itself to a wide variety of designs and arrange-
ments. The concepts proposed to date may be classified according to type
as being internally or externally cooled and either compact or open ar-
rangement of cycle. Such classification has been brought about in an at-
tempt to present designs which minimize bismuth and uranium inventories.

If we assume the cost of U235 to be $20/g and that of bismuth to be
$2.25/1b, a U-Bi solution of 700 ppm uranium by weight would cost ap-
proximately $6000/ft3. This high volume cost makes it very important to
design the LMFR with the minimum possible holdup.

In addition to the variety of cycle arrangements, several different
coolants are possible. The U-Bi may be used directly to produce steam, or
a secondary fluid such as NaK or sodium may be used. The LMFR has
also been proposed as the heat source for a closed-cycle, gas-turbine power
plant [2].

23-1.1 Externally cooled LMFR. In an externally cooled LMFR the
fuel is circulated through the core to an external heat exchanger, where the
heat is removed by the secondary fluid. This type provides the simplest
core design, requiring simply an assembly of graphite pierced with holes
for circulation of liquid-metal fuel. The major problems of heat transfer
are essentially removed from the core design. |

23-1.2 Internally cooled LMFR. The internally cooled LMFR is de-
signed so that the liquid fuel remains in the reactor core. The core thus
acts as a heat exchanger in which the heat is transferred to a secondary
fluid flowing through it to an external heat exchanger or steam generator.

The internally cooled design offers a means of substantially reducing
the U-Bi inventory of the system, but considerably complicates the design
of the core. The core must be designed to accommodate two fluids and suf-
ficient surface for transferring heat from one to the other. The introduction -
of a secondary fluid in the core requires a greater uranium concentration
than in the externally cooled system, which has only U-Bi and graphite in
the core. The required concentration cannot be achieved with U-Bi so-

*Based on material by T. V. Sheehan, Brookhaven National Laboratory, Upton,
L. I.,, New York.

832
23-1] | REACTOR DESIGN 833

From

Recuperator
To Turbine
A

     
  
 

Volatile
Fission Prod.

 

 

 

 

U-Bi

< Core
\@J N~—— ]mm{mmlflflflmuflm

Graphite Reflector

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Heat Exchanger

Dump

Fig. 23-1. Externally cooled compact arrangement LMFR for closed-cycle gas
turbine.

lutions, since these concentrations approach the solubility limits for the
temperatures presently being considered (400 to 550°C).

23-1.3 Compact arrangements. The compact arrangement may best be
described as an integral or “pot type’’ design and may be internally or ex-
ternally cooled. In such a design [1] the fluid fuel remains in the core ex-
cept for small amounts which are withdrawn for reprocessing. The breeding
fluid acts as a coolant by circulating through blanket and core and thence
through heat exchangers which are also contained within the primary
reactor vessel.

Figure 23-1 shows a concept of an externally cooled compact reactor
834 ENGINEERING DESIGN [cHAP. 23

arrangement for a closed-cycle, gas-turbine power plant [2]. In this ar-
rangement the fuel is circulated through the core and heat exchanger,
which are contained inside the same vessel. The compact arrangement
offers a means of reducing the U-Bi inventory over a particular reactor
designed with an open-cycle arrangement. It does, however, increase the
problems associated with design of the core, blanket, and reactor vessel.
The compacting of all the equipment into a single vessel reduces the flexi-
bility of mechanical design which the open arrangement allows, as well as
intensifying the problems of thermal expansion. The reactor vessel not
only becomes larger, but the number of openings is also increased, both of
which complicate the vessel design. Nevertheless, as operating experience
with materials and equipment becomes available, the compact arrange-
ment may provide a means of improving the economics of the LMFR
system.

23-1.4 Open arrangements. The open arrangement is the type receiving
the most consideration at present because of the flexibility and simplicity
of design it affords the system components. Figure 23-2 shows one con-
cept of an externally cooled LMFR using the open-cycle arrangement [3].
In this design both blanket and core fluids are circulated to heat ex-
changers located outside the reactor vessel. This type of arrangement also
allows greater freedom of design for maintenance of equipment. Means
must be provided for removal and/or maintenance of system components
under radioactive conditions. The open arrangement makes it easier to
provide such facilities. The major disadvantage of this arrangement is the
high U-Bi inventory.

The open-cycle arrangement may also be employed in an internally
cooled LMFR to reduce fuel inventory, but it introduces those problems
peculiar to internally cooled systems.

23-1.5 Containment and safety requirements. The high negative tem-
perature coefficient and low amount of excess reactivity available make the
LMFR inherently stable and safe. However, any rupture of the primary
system, whether by reactor excursion or otherwise, would release fission
products and polonium to the surrounding atmosphere. The primary
system must therefore be surrounded with a secondary vessel for contain-
ment of radioactivity in case of such a failure. Since all materials in the
reactor core have very low vapor pressures, the containment vessel need
not be designed to withstand any appreciable pressure. The containment
problem in the LMFR is one of containing the high-temperature liquid
metal together with fission products, and such containment can be ac-
complished by lining the reactor and primary circuit cells with a gastight
steel membrane. This containment vessel also acts as a catch basin for
recovery of U-Bi in case of leaks.
23-1] REACTOR DESIGN 835

-——§ 1= Na To Superheater

/ = Vet
Saturated 600 / And Reheater

Steam To Super-

'1 heater Reheat And Superheat

Sodium Inlet | Liquid Metal Pumps

And Outlet Box:

'.

 

 

 

       

 

Water Inlet
and
Outlet Box ™

 

 

 

 

 

 

il
i

 

pay L L L ! !
Gas Takeoff ’ 3 (iQTE
Jh
Processed Th-Bi

Boiler Bundle 41— % { , Slurry Return

 

 

 

 

 

 

 

 

 

 

 

 

Slurry Suction

Blanket c:! Header

(Th-Bi, Graphite)

‘ /. L :._:4%\
l i = 1
Liquid;;fl/ l g

 

Superheat + Reheat
Bundle

 

 

- — Slurry Drawoff
Graphite Core

 

   

 

 

 

 

From Slurry
Coolers

 

To Slurry

 

 

 

To Fuel

 

 

 

 

 

 

 

 

 

 

- —=4— 5000 Coolers
Processing Slurryl
System Discharge
— Header ‘
Liquid Fuel Pumps l To Fuel Th-Bi Slurry Pump
Dump

F1ag. 23-2. Externally cooled open-cycle arrangement LMFR.

The arrangement of the containment vessel also depends on the heat-
removal design. If an intermediate heat-transfer fluid such as sodium or
NaK is used, the containment may be handled as above. If a direct U-B1 to
steam cycle is used, a double-wall heat exchanger must be used to maintain
double containment, unless the entire building is constructed to act as the
second containment barrier.

In the event of a leak in the system, the U-Bi would be drained to a
dump tank. This tank would be provided with adequate cooling to remove
the decay heat from fission products.

23-1.6 Design methods. The vessels in an LMFR are designed in ac-
cordance with the Code for Unfired Pressure Vessels [4]. Vessels would be
of welded construction with all seams radiographed and stress-relieved.

The design temperature used can be as high as 1100°F. For 2% Cr-1%
Mo steel, this gives allowable stresses of 4200 psi for normal operating condi-
tions and 9200 psi for emergency, short-duration conditions. These figures
correspond to 1% creep strength for 100,000 hr and 10,000 hr, respectively.
836 ENGINEERING DESIGN [cHAP. 23

23-1.7 Maintenance and repair provisions. Provisions for maintenance
and repair of the LMFR raise several problems. It is anticipated that a
substantial level of activity will be induced in the system by the circulating
fuel. This means that the system should be designed so that it can be main-
tained despite the high radiation level. Several approaches, not mutually
exclusive, to this problem are being considered:

(1) If maintenance or repair to a component is required, the entire com-
ponent will be removed from the system and a new one inserted.

(2) All connections between components will be made in one area, fully
biologically shielded from the components themselves. When a com-
ponent is to be removed, its connections are shielded from adjacent connec-
tions by portable shielding if the work is to be done directly rather than re-
motely. The connections are broken and the shielding is removed above the
pipes leading to the component in question. The component is removed
with the overhead crane and a new one set in place. The shielding is re-
placed, and the connections are remade. The connections are accessible
and pipes do not overlay each other so as to prevent removal of any
disconnected component. Unfortunately, placing all connections in one
channel increases the fuel inventory since the piping for this arrangement is
somewhat longer than that required for a more conventional arrangement.

(3) Both mechanical and welded connections are being studied, with a
view toward the ease with which connections can be made and broken both
directly and remotely.

(4) Remote methods of performing maintenance tasks (welding and cut-
ting pipe, making and breaking flanged joints and closures) are being
studied, since direct maintenance will not be possible in some areas.

(5) Fluidized powders, shot, and liquids are being studied as possible
portable shielding media.

23-2. HEAT TRANSFER*

In the open-cycle externally cooled, two-fluid LMFR, the bismuth-
uranium solution serves as the primary coolant as well as the fuel. In the
reactor itself, there is no actual heat transfer. Instead, the solution acts as
a transporter of heat to an external heat exchanger. In evaluating bismuth
as a primary coolant, it is helpful to make a comparison between it and
three other coolants: sodium, a typical alkali metal coolant; LiCl-KCl
eutectic, a typical alkali halide salt mixture; and water. (The salt eutectic
used here would not be a suitable primary coolant for a thermal reactor.
Its heat transfer properties, however, are typical of salt coolants.)

*Based on material by O. E. Dwyer, Brookhaven National Laboratory.
23-2] HEAT TRANSFER 837

The ideal primary coolant for a nuclear power reactor should have the
following characteristics:

(1) High heat-transfer rates.
(2) Good gamma, absorptivity.
(3) Low pumping power requirements.
(4) Low melting point.
(5) Low vapor pressure.
(6) Low corrosion rate.
(7) Low chemical reactivity.
(8) Low neutron absorption.
(9) Low induced radioactivity.
(10) Low cost.

In order to have the above characteristics, the coolant should have the
following physical properties in either a high or low amount:

(1) Density (high): affects pumping power requirements, heat-transter
characteristics, and gamma shielding requirements.

(2) Thermal conductivity (high): affects heat-transfer characteristics.

(3) Specific heat (high): affects heat-transfer characteristics and coolant
flow rate. :

(4) Viscosity (low): affects pumping power requirements and heat-
transfer characteristics.

(5) Melting point (low): affects auxiliary heating requirements.

(6) Vapor pressure (low): affects mechanical design of reactor and system
components.

(7) Volume change on fusion (low): affects startup and shutdown pro-
cedures.

(8) Coefficient of volumetric expansion (high): affects thermal pumping
capacity and, where primary coolant is also the fuel, reactor reactivity.

(9) Electrical resistivity (low): affects applicability of electromagnetic
pumps.

Table 23-1 summarizes the physical properties of bismuth which are
relevant to nuclear reactor design and in the temperature range of practical
interest from the standpoint of electrical power generation [5,6].

23-2.1 Nuclear aspects of coolants. From the nuclear standpoint, five
important characteristics of primary reactor coolants are their capacities
for (1) absorbing thermal neutrons, (2) slowing down neutrons to the
thermal energy level, (3) absorbing gamma radiation, (4) developing in-
duced radioactivity, and (5) resisting radiation damage.

In Table 232 the thermal neutron absorption cross section and neutron-
slowing-down power of Bi are compared with those of Na and H20. Bis-
838 [cHAP. 23

ENGINEERING DESIGN

TABLE 23-1

PaYSICAL PROPERTIES oF BisMuTH

 

 

 

 

 

 

Atomic weight 209

Melting point 271.0°C (520°F)
Boiling point 1477°C (2691°F)
Volume change on

fusion —3.329,

Temperature, °C 300 400 500 600
Temperature, °F 572 752 932 1112
Vapor pressure,

mm Hg 109w 3.5X1075*%12.3x1074|6.3x 1074
Density, g/cm3 10.03 9.91 9.79 9.66
Specific heat,

cal/(gm)(°C) 0.0343 0.0354 0.0365 0.0376
Viscosity, centipoises 1.66 1.37 1.16 1.00
Thermal conductivity, |

Btu/ (hr) (ft) CF) 9.9 9.0 9.0 9.0
Electrical resistivity,

ohms 128.9 134.2 139.8 145.2
U solubility, ppm 480 1850 5100 13000

 

 

*Extrapolated results.

muth with a macroscopic cross section of 9.0 X 104 ecm™1! at 450°C has
the lowest neutron absorption characteristic of any practical coolant, with
the exception of D20 and certain gases. Its “‘reactor poisoning’’ effect is
at least an order of magnitude below those of sodium and water. The
slowing-down power of Bi is very low, however, which means that when it is
used as the primary coolant in a thermal reactor it contributes very little
moderating capacity. The term £No, in Table 232 represents the decrease
in the natural logarithm of the neutron energy per centimeter of travel
through coolant.
The gamma absorption coefficient, u, is defined by the equation
dl = —uldx (23-1)
and has the units of em™!. Values of u for 450°C Bi at several gamma
energies are shown in Table 23-3, along with those for Na and H-O.
Bismuth, because of its high density, is an excellent absorber of gamma
radiation, which means that it provides considerable internal shielding.
The values presented in Table 23—-3 are estimates based on the theoretical
calculations of Davission and Evans [8].
23-2]

 

 

 

HEAT TRANSFER 839
TABLE 23-2
SoME NUCLEAR PROPERTIES OF VARIOUS REACTOR COOLANTS
Thermal Macro- Thermal )
: : Slowing-
neutron scopic scattering £ * down
Coolant Temp., Cross CTross Cross dimeilsion- ovr;r
°C |section [7],| section [7],| Density| section, 1 P ‘
ess ¢No,
O, N a9, P, O, — 1’
barns cm™! g/cm3 | barns om
Bi 450 0.032 0.00090 | 9.82 9 0.0095 | 0.0024
Na 450 0.505 0.011 0.841 4.0 0.084 0.0074
H-0 250 — 0.018 0.802 — — 1.23

 

 

 

 

 

 

 

 

 

 

*Average decrease in the natural logarithm of the neutron energy per collision.

TABLE 23-3

VALUES oF u, THE GAMMA ABSORPTION COEFFICIENT,
FOR VARIOUS REACTOR COOLANTS AS A FuNcTioN oF ENERGY

 

 

 

 

 

 

 

 

 

 

Energy, Mev
Coolant Te;%p "
0.5 1.0 1.5 2.0 3.0
Bi - 450 1.57 0.70 0.52 0.44 0.41
Na 450 0.070 0.051 0.042 0.036 0.029
H20 250 0.078 0.057 0.046 0.039 0.032

 

 

 

Regarding the tendency for developing induced radioactivity, Bi has a
serious disadvantage, owing to the formation of Po?!0, a very energetic

alpha emitter with a 138.3-day half-life. Its formation and decay can be
represented as follows‘:

Bjr(19mb) s B;210 _—__B) Po210 — % Pp206

5d

138.3d

Po0210 is one of the most poisonous materials known, the maximum allow-
able concentration in air being 7 X 10~ uc/ml or 3.75 X 10~8 ppm.
Another troublesome feature of Po210 is its tendency to scatter throughout
840 ENGINEERING DESIGN [cHAP. 23
any accessible volume, due to recoil from its high-energy alpha emission.
Thus, spillage of solutions containing Po?!0 constitutes a most serious phys-
iological hazard. In the LMFR, however, it is not believed that the presence
of Po?!0 in the fuel stream creates a more serious radioactivity problem
than already exists as a result of the fission products.

Sodium is not free of the radioactivity problem either, but as a primary
coolant it is not as bad in this respect as Bi. Water is comparatively free
of induced radioactivity after short holdup times. For the same flux
conditions, Na will give over 20,000 times as much radioactivity, on a
roentgen basis, as H20.

Liquid metals, because of their simple atomic structure, suffer no radia-

tion damage.

23-2.2 Pumping-power requirements. An important criterion for as-
sessing the relative merits of different coolants is the amount of pumping
power required for a fixed rate of heat removal in a given application. The
three main pressure drops in the primary coolant circuit are those in the
reactor, the external heat exchanger, and the interconnecting piping. A
comparison of the four different types of coolants will now be made on the
basis of their relative pumping-power requirements, with respect to the
interconnecting piping and the heat exchangers. The physical properties
of the coolants are listed in Table 23—4. The properties of the first three
are evaluated at 450°C, as a typical average primary coolant temperature
for such coolants, and those for water at 250°C.

TABLE 234

PaYsicAL ProPERTIES oF SOME TyricaL REAcTOR COOLANTS

 

 

 

 

 

 

 

 

Property Bi Na KCI-LiCl | H20
450°C 450°C 450°C 250°C
Melting point, °F 520 208 664 32
Boiling point, °F 2691 1621 — 212
Density, 1b/ft3 615 52.5 103 50.0
Specific heat, Btu/(1b)(°F) 0.036 0.304 0.31 1.16
Heat capacity, Btu/(ft3)(°F) 22.1 15.95 31.9 57.8
Thermal conductivity,
Btu/ (hr)(ft) (°F) 8.95 39.5 1.47 0.357
Viscosity, cp 1.28 0.245 3.4 0.110
Prandtl number, C, u/k 0.0125 0.00454 1.7 0.863

 

 
23-2] HEAT TRANSFER 841

The pumping power required to circulate the coolant through the piping
system per unit rate of heat transport for a fixed temperature rise in the
coolant has been shown [9] to be

_ pumping power uo2 ~
bp Tont load (a constant) 520,28 (23-2)

 

The quantity u2/p2C,28, represented here by the symbol X, is an index
of the pumping power required to circulate a coolant through a fixed
piping system, for a given heat load. Table 23-5 gives relative values of X
for the four typical coolants mentioned above. The units and values of the
physical properties used in evaluating X are the same as those given in
Table 23—4.

TABLE 23-5

RELATIVE VALUES OF X FOR VARIOUS COOLANTS
Frowing THROUGH A FIxXED PIPING SYSTEM

 

 

 

Coolant Temp., °C X x 104
B1 450 308
Na 450 77
LiCl-KCl 450 32
eutectic
H20 250 1.7

 

 

 

 

 

The very large spread in pumping-power requirements 1s striking. Bis-
muth has about four times the pumping-power requirements of sodium and
both have manifold greater requirements than that of water, which has
the least of any known liquid. The tremendous superiority of water as a
heat-transport medium is due to its low viscosity and very high volumetric
heat capacity.

23-2.3 Heat transfer for LMFR. So far as is known, no heat-transfer
data have been obtained for liquid bismuth. However, several investigators
[10-14] have published experimental heat-transfer results on the bismuth
lead eutectic and on mercury. For these results the Lubarski and Kofiman
equation [15] expresses the results most closely:

’-‘kg = 0.625(DV,Cp/k)*4. (23-3)
842 ENGINEERING DESIGN [cHAP. 23

This equation may be used for turbulent flow in round tubes or for turbulent
flow outside round tubes.

In obtaining the heat-transfer coefficients for comparison with bismuth,
the sodium coefficients were calculated from the Martinelli-Lyon relation-
ship. The coefficients for molten salt and water were calculated from the
conventional Dittus-Boelter equation.

Using the above relationships and assuming (1) total fixed heat load,
(2) fixed diameter of tubes, (3) fixed inlet and outlet temperatures, (4) av-
erage bulk temperature of coolants same as in Table 23—4, and (5) combined
heat-transfer resistance of tube wall and second fluid equals 0.001, a typical
value for 1-in. ID alloy steel tubes with 0.1 in. wall the values in Tables
23-6 and 23-7 were calculated. Although the heat-transfer characteristics
of bismuth are slightly inferior to those for sodium, it is clear from these
two sets of calculations that all four coolants behave similarly.

The heat-transport capability of bismuth are simply related to its volu-
metric heat capacity. The values of this property are given in Table 23—4.
Bismuth is definitely superior to sodium but inferior to the fused salt and
water.

To achieve good thermal contact between bismuth and a solid metal
surface, the surface must be cleaned to a high polish, the bismuth must
be free of oxide and dissolved gases, and the system must be filled under a
high vacuum. Gases or oxides on the heat-transfer surface can greatly
reduce the heat-transfer coefficient for bismuth. Bismuth has a less stable
oxide than the oxides of iron, chromium, and nickel which may be present
on the tube surfaces. Hence the bismuth would have a tendency to non-
wet the walls.

Good wetting of alloyed steels by bismuth may be achieved by adding
small amounts of alkali or alkaline earth metals, by heating to high tem-

TABLE 236

CoMPARISON OF CoOLANTS IN HEAT-EXCHANGER DESIGN
WHEN NUMBER OF TUBES IN ParaLLEL 1s FIXED

 

 

 

Coolant Xflggiy h, U, Relative size of
£t /sec’ Btu/(hr)(ft)2(°F) | Btu/(hr)(ft)2(°F) | heat exchangers
Bi 15 2700 730 1.00
Na 20.8 10230 910 0.87
LiCl-KCl 10.4 2400 706 1.12
eutectic
H20 5.73 2360 . 703 1.12

 

 

 

 

 

 

 
23-3] COMPONENT DESIGN 843

peratures (above 1200°F), or by both. For good heat transfer with bismuth
extreme care must be taken to ensure oxide- and gas-free systems.

23-2.4 Heat-exchanger design. In a commercial liquid-metal fuel sys-
tem, the primary bismuth coolant would probably exchange heat with a
secondary metal coolant before generating steam. Typical conditions for
a 5-Mw countercurrent bismuth-sodium heat exchanger are given in
Table 23-8.

23-3. CoMPONENT DEsiGN*

This section discusses the design and development experience obtained
on components required in LMFR systems. Besides the requirements for
these systems, considerable component development is needed in the re-
search and development program for experimental apparatus. Both kinds
of components are treated here in detail and by case histories.

23-3.1 Pumps. In the case of liquid-metal pumps, which can be classified
as mechanical or electromagnetic, a good deal of preliminary development
work has been done by the Fairchild Engine and Airplane Corporation
Nuclear Energy for Propulsion of Aircraft Division (NEPA), the Allis-
Chalmers Co., the Babcock & Wilcox Co., and the Government Labora-
tories, KAPL, ORNL and ANL [19].

TABLE 23-7

CoMPARISON OF Coorants IN HEaT-ExcHANGER DESIGN
AT Fixep LiNEAR VELocCITY OF 15 FT/SEC

 

 

 

e
Coolant | of tubes Temp., h, v, size of
‘o °C | Btu/(hr)(ft)2(°F) | Btu/(hr)(ft)2(°F) heat

parallel exchangers
Bi n 450 2770 730 1.00
Na 1.38n 450 8810 897 0.88
LiCI-KCl | 0.69n 450 3200 762 1.03

eutectic

H-0 0.42n 250 5150 837 0.94

 

 

 

 

 

 

 

 

*Based on a contribution by C. Raseman, H. Susskind, and C. Waide, Brook-
haven National Laboratory.
844 ENGINEERING DESIGN [cHAP. 23

TABLE 23-8

TypricAL CoNDITIONS IN A COUNTERCURRENT,
Bi—-NA HeaT EXCHANGER

 

Tube material Low Cr-Steel
Thermal conductivity of tube, Btu/(hr)(ft) (°F) 15.8
Tube inside diameter, in. 0.70
Tube thickness, in. 0.100
Tube spacing (triangular), in. 0.250
Bi temperature (bulk), °F 850

‘Bi velocity (outside tubes), ft/sec 15.0
Bi heat transfer coefficient, Btu/(hr)(ft)2(°F) 3,390

Na temperature (bulk), °F 750

Na velocity (inside tubes), ft/sec 25.5
Na heat transfer coefficient, Btu/(hr) (ft)2(°F) 12,300
Over-all heat transfer coefficient, Btu/(hr)(ft)2(°F) 1,015
Fraction of resistance offered by tube wall 0.60
Heat flux (outside tube surface), Btu/(hr)(ft)2 101,500
Power density, Btu/ (hr) (ft)3 510,000

Bi, ft3/mw heat 0.56
Na inventory, ft3/mw heat 0.45

 

 

 

 

Electromagnetic pumps. In the early days of the LMFR project, a mag-
netic pump for Bi was described by B. Feld and L. Szilard [20,21]. The
Fuel Processing Group of Brookhaven National Laboratory required pilot-
plant pumps that would circulate uranium-bismuth fuel with absolutely
no leakage. The U-Bi1 fuel was eventually to be circulated through an
experimental hole in the Brookhaven reactor where fission products and
polonium would be generated. Since a small flow rate of approximately
1 gpm was desired and efficiency was of little concern, it was decided to
use an electromagnetic pump. |

An experimental loop [22] was set up to circulate nonradioactive U-Bi
by means of a General Electric Model G-3 ac (Faraday) electromagnetic
pump. This loop ran continuously for 2400 hr. During the first 160 hr the
rig was operated isothermally at a temperature of 645°F; during the
remainder of the time, the loop was run isothermally at 840°F. The U-Bi
solution was circulated for most of this period at a rate of 1 gpm. There
was no sign of plugging or flow restriction.

The General Electric G-3 Ac pump was calibrated (Figs. 23-3 and 23—4)
in another AISI type-347 stainless steel liquid bismuth loop at 930°F
[22]. It was operated continuously for over 13,000 hr.
23-3] COMPONENT DESIGN 845

0.16

 

0.14 |— —

0.12 — ’ —

0.10 — —

0.08 — —

Efficiency, %

0.06 - —

0.04 — —

0.02 —

 

 

l l l I 1 l l |
0 0.2 04 06 08 1.0 1.2 1.4 1.6 1.8
Flow, GPM

F1e.23-3. AC electromagnetic pump efficiency. Molten bismuth in AISI type-347
stainless steel cell. (Manufactured by General Electric Co.)

 

 

10 l I | | | I | |

200 Volts
8 I~ —

 

150 Volts

 

Head, PSI

2 100 Volts —

;
: = 50 Volts 1 | | 1 |

0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8
Flow, GPM

 

 

 

 

Fig. 23-4. AC electromagnetic pump characteristics. Molten bismuth in AISI
type—347 stainless steel cell. (Manufactured by General Electric Co.)

The same pump was used to circulate bismuth at 930°F in a 239, Cr-19,
Mo steel loop. - The efficiency and characteristic curves were somewhat
lower than those obtained in a stainless steel loop. This is probably due
to short-circuiting of magnetic flux in the ferritic steel walls.

A theoretical study [23] was prepared by the Atomic Energy Research
Establishment at Harwell, England, for linear-induction pumping of
bismuth. The report indicates the feasibility of using this type of pump.
Linear-induction pumps have been built and successfully used at Ames
846 ENGINEERING DESIGN [cHAP. 23

Laboratory to circulate Mg-Th eutectic (37 w/o Th) and Bi-U alloy
(65w/oU) in an Inconel-enclosed tantalum loop [24,25]. The pump
operated successtully in the Mg-Th system for 2000 hr at 1470°F with a
temperature differential of 250°F, and in Bi-U for 5250 hours at 1740°F
with a temperature differential of 210°F. For calibration, about 1 gpm
of Bi-U was pumped at 750°F against a head of 0.5 in., with an efficiency
of 0.169.

Mechanical pumps. Most pump development work has been aimed at
pumping sodium or sodium-potassium alloys. The most serious problem
relative to the design of a mechanical liquid-metal pump appears to be
that of suitable bearings and seals.

Bismuth was pumped by NEPA in 1950 [26]. The system was operated
for 37 hr, the maximum flow rate measured was 2 gpm, the maximum head
developed was 66 psi, and the maximum bismuth temperature reached was
1765°F. The pump was a modified Browne and Sharpe No. 206, machine-
tool-coolant pump.

In another experiment [27] NEPA circulated bismuth with a 50-gpm
centrifugal pump for 100 hr at a mean temperature of 1500°F with a
temperature differential of 500°F. An accumulation in the sump of a
residue high in oxide content and dissolved elements reduced the flow and
forced suspension of operation. This residue probably resulted from an
impure inert atmosphere above the liquid metal. The container material
selected was AISI type—347 stainless steel which had shown some promise
in bismuth solubility tests at temperatures up to 1800°F.

The California Research and Development Corporation made a survey
of the various types of pumps that might be used for liquid bismuth and
came to the conclusion that a centrifugal pump would best fit the need.
A test unit was built that operated for 1037 hr, and a report [28] stated
that the centrifugal pump proved to be a very satisfactory means for cir-
culating bismuth in an isothermal system at 700 to 750°F. This pump and
its driver are on a common shaft, the shaft being top-suspended with all
bearings 1n the motor chamber. Space was provided for a labyrinth to
separate the pump chamber from the motor chamber, although no seal was
used during operation. This pump has also been used to circulate mercury
in a test loop at BNL. It has been run successfully for an accumulated time
of over 4000 hr. |

Brookhaven has developed a totally canned overhung-impeller centrif-
ugal pump. Figure 23-5 shows the major design features of this pump.
These units pump 5 to 25 gpm Bi against heads up to 30 ft while operating
at 525°C. These sump-type pumps run with no bearings in the liquid metal
and have proved reliable so long as sufficient internal baffling is included
to stop surface splashing.

There are several centrifugal pumps that have been used to circulate
23-3] COMPONENT DESIGN 847

   
  
   
  
   

Induction Motor

Gas & Vac.

Liquid
Level

Cooling
Probes

Water

Flange Cooling
Tubes

Shaft Cooling
Chamber

Pump Baffles &
Splash Can

Overhung
Impeller Shaft

Liquid Level
Side Arm Discharge Inlet

Fig. 23-5. Canned-motor centrifugal pump developed at Brookhaven.

lead-bismuth eutectic [29,30]. They are all vertically mounted sump
pumps with overhung shafts and impellers. All would require a can around
the motor and shaft for a hermetic seal.

The University of California has used a double-volute pump which is
rated at 30 gpm and a 40-ft head at 1000°F.* The lower bearing is 2 ft
above the liquid metal. The pump utilizes a packing gland (Johns-Manville
Super Seal No. 6) adjusted to allow helium at 2 psig to leak out of the
system at a rate of about 10 ft3/hr.

The pump used at North American Aviation, Inc. [30] is made of cast
steel. The lower bearing is cooled with a water jacket and a graphite seal

*The vendor is Berkeley Pump Co., Berkeley, Cal. The 73-in.-diameter impeller
and the pump casing are made of AISI type—410 steel. The pump is V-belt driven
by a 30-hp motor.
848 ENGINEERING DESIGN [cHAP. 23

minimizes gas leakage from the casing. A flow of 0.82 gpm at 400-rpm
shaft speed and a temperature of 700°F was maintained until oxide dross
forced shutdown of the pump after 496 hr.

A completely canned, modified Series T-3¢ MD Duval stainless steel
pump was used at the University of California to circulate mercury [31].
The packing gland was replaced with a bushing and any metal leakage
was drained to a reservoir. The pump was driven by a 5-hp, 3-phase in-
duction motor at a shaft speed of 1200 rpm.

North American Aviation, Inc. has circulated tin with a graphite pump
at 2 gpm against a head of 22.5 psi at 1830°F [32]. The pump has a 4-in.-
diameter impeller and is driven by a variable speed (20 to 2000 rpm) pc
compound-wound motor mounted outside the gastight enclosure to avoid
the high temperatures. A rotating Graphitar bushing on hardened steel
provides the gas seal. The spindle bearings are in a cooled housing. The
pump was operated for 500 hr in one run; this was followed by additional
runs. To overcome differential thermal expansion, a molybdenum adapter
joins the graphite shaft to the stainless steel spindle.

A miniature canned centrifugal pump to circulate bismuth, ideally
suited for in-pile work, has been developed by the Atomic Energy Research
Establishment at Harwell, England. The over-all pump dimensions are
32-in.-diameter by 24% in. long, with a 2-in.-diameter impeller. The bis-
muth flow is 1.5 gpm with a head of 9 ft. The motor rating is 0.75 hp and
2800 rpm. Two gas-lubricated bearings are utilized. The material of con-
struction is 21 Cr-1 Mo steel.

The Allis-Chalmers Manufacturing Co. has built a canned rotor centrif-
ugal pump with fluid piston-type bearings to pump bismuth at 1050°F.
The pump is rated at 10 gpm and a head of 25 ft, with an efficiency of 109,.
Those parts of the pump in contact with the bismuth are made from AISI
type—410 steel. The pump was used in loop G at BNL to pump bismuth
at 1020°F with a temperature differential of 300°F. After 15.5 hr the pump
failed, due to scoring of the bearings and seizure of the can by the rotor.

23-3.2 Valves. The standard-stem packed gate valves used in early
NEPA bismuth tests [26] proved that special valves would be required
for successful liquid-bismuth operation. High leakage rates through the
packing caused maintenance difficulties throughout the tests.

A 1i-in. Fulton-Sylphon bellows-type stainless steel valve was cycled
1000 times at the rate of 77 times/min against bismuth at a temperature
of 1000°F and a pressure of 25 psig. No failure of the bellows or other
valve parts occurred. NEPA also checked valves for metal-to-metal
self-welding effects [33]. Tests of valve operation reached 1500°F with
liquid bismuth on Standard Stellite-faced poppets and seats without indi-
cation of self-welding effects.
23-3] | COMPONENT DESIGN 849

The two types of valves which have seen extensive service up to 1050°F
in liquid-metal fuel systems are standard Y pattern globe valves and
needle valves. Due to the stringent requirements of zero gas leakage
(into or out of the metal systems), the only acceptable stem seal has been
a steel bellows. Packings are unacceptable.

Brookhaven National Laboratory has used both types of valves exten-
sively [22,34]. The 1/2-in. IPS 150-lb Y pattern globe valves constructed
from AISI type-347 stainless steel for all parts in contact with bismuth
(including bellows, stem, and disk) have been used continuously for over
8000 hr at 930°F without mishap. Similar valves with mild carbon steel
disks (instead of type—347 stainless steel) have been used at 930°F for over
13,000 hr without failure or extensive corrosion.

A high-velocity loop operating with bismuth at 1020°F at BNL uses
1-in. IPS 150-1b Y pattern globe valves made from 239, Cr-19, Mo steel,
AISI type—430 steel bellows and disk, and AISI type—416 steel stem.

Needle valves (1/8-in. IPS AISI type—347 stainless steel construction,
including the bellows) have been in use for intermittent service (i.e., drain
valves).

As an additional safety measure, 1/2-in. IPS globe valves used in an
in-pile loop at Brookhaven National Laboratory have utilized two sets of
bellows [34]. The space between the two bellows was pressurized with inert
gas which was continually monitored to detect pressure changes (thus
indicating a valve leak). None was detected.

The valve drives have been modified to facilitate remote operation.
The globe valve handwheels are replaced by gears and these are, in turn,
connected to extension rods projecting through the enclosures. Extension
rods are welded directly to the needle valve bellows. Universal joints and
right-angle gear drives are used for changes in direction between valve and
operator. When relatively gastight enclosures are desired, as in in-pile
loops, the extension rods project through rubber-gasketed compression
seals.

Oak Ridge National Laboratory has reported on the use of special high-
temperature packing [35] for valve stems. This packing consists of suc-
cessive layers of Inconel braid, graphite, nickel powder, and another layer
of Inconel braid. |

It has been shown at practically all AEC installations that two sections
of a circulation system can be isolated from each other by freezing a short
section of connecting pipe. This plug can be remelted and flow resumed
after a short wait. This type of seal is undesirable for uranium-bismuth
solutions, however, since the uranium will deposit at the cold surface.

23-3.3 Piping. Layout features. The most important considerations in
designing piping for a liquid-metal fuel system are the considerable thermal
850 ENGINEERING DESIGN [cHAP. 23
expansion of the pipe when heated from room temperature to operating
temperature, and the expansion of bismuth upon freezing (3%). The former
condition prescribes the type of supports required, while the latter de-
termines the methods and techniques for freezing the metal.

In general, it is desirable to hang pipe from overhead supports, prefer-
ably spring-loaded hangers with straps around the pipe insulation. Heavy
vessels may be anchored to hangers by brackets welded to the wall. Care
should be taken to see that these brackets do not act as a large heat sink.
If the system is supported from below, heavy vessels should “float” by
locating them on freely moving bearing raceways.

Freezing the liquid metal in the system, especially in components with
bellows, should be avoided. However, in case of emergency, the metal
should be frozen towards the free surface. For this reason, a system should
always contain a surge (or expansion) tank, located at the highest ele-
vation.

The use of an integral fill tank, located at the lowest point in the system
to permit charging the loop with metal through a pipe “dip leg”’ com-
pletely immersed in the metal, is recommended. The application of gas
pressure on the fill tank will transfer the metal slowly into the loop. By
charging the metal from the bottom, into a previously evacuated system,
gas entrainment will be minimized. A sintered metallic filter should be
used to remove oxide and other scum from the metal while filling the loop.
This filter should always be located outside the fill tank, since this will fa-
cilitate removal of the filter when it becomes clogged and will prevent crack-
ing of the pores if the contents of the fill tank freeze.

The loop may be drained into a vessel which can be either the fill tank
or a separate drain tank. Piping lines should be sloped to facilitate drain-
age; undrainable pockets should be provided with separate drain lines or,
if possible, eliminated. A typical liquid-bismuth loop layout is shown 1n
Fig. 23-6.

Bellows. Several types of metal bellows have been used at Brookhaven
National Laboratory in bismuth systems at 930°F. AISI type—347 stainless
steel welded bellows have been used continuously in 1/2-in. IPS globe valves
for periods as long as 13,000 hr. The bellows have not, however, been ex-
tensively cycled in bismuth. Their dimensions are 2 in. OD by 1-in. 1D
by 0.018-in. thick and contain 32 convolutions. Two AISI type-410 steel
welded bellows have been bench cycled 32,000 and 120,000 times, respec-
tively, in bismuth at 1020°F and should, therefore, be satistactory. They are
used in pressure transmitters, and are 13 in. OD by 3/8-in. ID by 0.009-1n.
thick and contain 22 convolutions. At this time, one AISI type—430 steel hy-
draulically formed bellows, used in 1/2-in. IPS globe valves, has been
bench-cycled with helium over 200,000 times at 1020°F without failure. Its
dimensions are 13-in. OD by 7/8-in. ID by 0.008-in. thick (two-ply).
23-3] COMPONENT DESIGN 851

Bellows tests at Argonne National Laboratory [36] have yielded the
following data:

(1) Failures have generally occurred at a weld; therefore bellows with
the least number of welds are favored. However, mechanically formed
bellows should be examined for cracks and other flaws that may be intro-
duced in the forming.

(2) There was no evidence that corrosion played a part in the failure of
any bellows.

(3) One predominant factor determining bellows life is the relative
amount of travel. |

(4) Other factors affecting bellows life are temperature and the relative
distribution between compression and extension. It was found that the
outer bellows failed before the inner bellows which operated at a higher
temperature.

(5) Some bellows designs had not failed up to 10° cyecles, at which point
the test was stopped.

Joints. Melal systems. In general, in these metal systems, all joints
should be welded for tightness and structural soundness. All weld joints
are made by standard inert-arc procedures. Complete procedure specifica-
tions have been prepared by BNL for inert-arc welding of AISI type—347
stainless steel pipe, fittings, and vessels for use with liquid metals. This
procedure was developed through the cooperation of the Metallurgy
Division of Oak Ridge National Laboratory. Specifications have, likewise,
been prepared at BNL for welding 239% Cr-19% Mo steel. AISI type—-502
steel welding rods are used in welding 23% Cr-19% Mo steel pipe. A pro-
cedure for welding 0.030-in.-thick tantalum tubing, as well as AISI type-316
stainless steel to tantalum, has been prepared at Ames Laboratory [24,25].

Experimental and operating procedures, however, often make it ad-
vantageous to have removable joints. These have been successfully used
at a number of installations. An oval cross-sectional ring for a flanged
joint was used by NEPA [37] in a bismuth system between 520 and 660°F
at 300 psig, and by the California Research and Development Corporation
[28] on 13-in. piping containing bismuth at 700 to 750°F.

Standard metallic ring-joint flanged connections have also been satis-
factorily used at the University of California and Brookhaven National
Laboratory [22,29]. The rings were of soft iron (in lead-bismuth systems)
and AISI type—347 stainless steel (in bismuth systems). At a temperature
of 930°F, the AISI type—-347 stainless steel joint has been found to be
helium leaktight to a mass spectrometer.

The ability of liquid metals and liquid salts to leak through extremely
small openings has made the use of helium mass-spectrometer leak testers
a specified test step. Halogen leak testers should never be used because of
the absorbed halogen which remains in the surfaces after the tests.
852 ENGINEERING DESIGN [cHAP. 23

Graphite system. Several graphite loops have been operated with bis-
muth at a maximum temperature of 2550°F [38] and with tin at tempera-
tures up to 2730°F [32]. Spherical joints held together with steel flanges
and bolts, or tapered joints threaded for assembly under tension, have been
the best. In addition, the joints may be fused to reduce leakage by coating
furfural on one face and hydrochloric acid on the other. However, even
with all these precautions, the systems were not absolutely tight to bis-
muth or tin.

Sight ports. Sight ports have been used to facilitate viewing the liquid
metal inside a closed system at the University of California [29] and at
Brookhaven National Laboratory. A satisfactory port consists of a glass
plate at the end of a steel bellows welded to the pipe. A normally closed
butterfly valve isolates the glass from lead or bismuth vapors. The valve
is moved by an externally mounted magnet or a handle projecting through
a Teflon-packed gland.

23-3.4 Heating equipment. Flexible Nichrome heater wire consisting of
a Nichrome inner wire, asbestos and glass insulation, and a flexible stainless
steel protective braid, is extremely useful for maintaining systems at tem-
peratures up to 1100°F for periods of time in excess of 10,000 hr [22].
Figure 23-6 shows the application of this type of heater in loop work.
Strip and tubular heaters have been in in-pile service for over 8000 hr [34].
A resistance heater has also been used as an internal heater submerged in a
lead-bismuth eutectic system [29].

Induction heating has been used on bismuth with good results [27].
A heating transformer in which the metal stream is the secondary circuit
has been used at Ames Laboratory in magnesium-thorium and uranium-
bismuth systems at temperatures up to 1740°F for periods of up to
5000 hr [24].

The use of graphite as a resistance heater in graphite loops has been suc-
cessful at temperatures up to 2700°F for short times (about 500 hr) [32,38].

23-3.5 Insulation. Samples of 26 insulating materials were tested for
possible reaction with molten bismuth [51]. In general, results indicated
that little or no reaction occurred when molten bismuth at 1832°F came
into contact with the unheated materials, but that none of the materials
would withstand contact with the bismuth for more than a few hours when
both were at 1832°F.

At BNL, Johns-Manville Co. Superex preformed pipe insulation and
Carborundum Co. Fiberfrax bulk insulation have been used extensively.

23-3.6 System preparation. Cleaning of equipment. Owing to the cor-
rosive nature of most bismuth compounds and the necessity for maintain-
23-3] COMPONENT DESIGN 853

 

F1g. 23-6. Typical liquid bismuth loop.
854 ENGINEERING DESIGN - [cHAP. 23

ing definite concentrations of additives in the fuel systems, the type of
container employed and the condition of the container-liquid interface is
of great importance. The presence of oxygen and other impurities in
soluble or insoluble form can accelerate the attack upon the container
material. As a result, it is desirable to remove all foreign material from
liquid-metal fuel systems before charging. Cleaning techniques for the
more important liquid-metal container materials are summarized as
follows.

Stainless steels. The committee of stainless steel producers of the Ameri-
can Iron and Steel Institute [40] recommend several techniques, depending
upon the type of impurity to be removed. In addition to these methods,
BNL has found electropolishing to be useful in removing surface oxides
[34]. In all cases, after the use of cleaning solutions the material is rinsed
thoroughly with water and dried by allowing a final alcohol or acetone
rinse to evaporate. |

Low-chrome steels. Several methods have been used for cleaning metals
of this type. One method is described [30] for PbBi systems in which
boiling detergent solution is used to remove dirt and scale, followed by a
distilled water rinse and drying under conditions of heat and vacuum.
The same reference describes the following cleaning procedure:

(1) Inhibitor: 109, HCI for 12 hr.

(2) Neutralization of HCI with Na2COs.
(3) Water rinse.

(4) 109, phosphoric acid wash.

(5) Drying with heat and vacuum.

The following technique has been developed at BNL for use with large
vessels:

(1) Degrease with trichlorethylene.

(2) 39, HNO3-179%, HCI solution for 30 min at room temperature.
(3) Flush with water.

(4) Repeat steps (2) and (3).

(5) Add 209, HCI solution for 5 min at room temperature.

(6) Rinse with water.

(7) Rinse with alcohol.

(8) Dry with inert gas blast.

Leak testing. Liquid-metal fuel systems which involve solutions con-
taining uranium and oxygen-sensitive additives (such as the magnesium
used in LMFR systems) require that precautions be taken to prevent air
leakage into equipment. In general, a sequence of leak detection is followed
in which gross leakage and structural faults are first eliminated by pressure
testing.  Suspected leaks can be verified by application of soap solution.
23-3] COMPONENT DESIGN 855

The helium mass-spectrometer leak detector has been found to be the most
useful as a final test. Systems found leaktight to helium are acceptable for
use in uranium-bismuth systems.

Preheating. A procedure for preheating equipment has been used at
BNL and elsewhere [34,41] in which the equipment is first evacuated to
less than 100 microns pressure and then heated, at a rate slow enough to
prevent pressure surges above 100 microns, to operating temperature.
This procedure has the advantage of removing condensables from the con-
tainer walls before they can react with the wall at elevated temperatures.

With the equipment at or above operating temperature, purified hydro-
gen may be introduced to reduce any surface oxide that might be present.
This step is frequently done with the liquid metal present in the charging
vessel in order to reduce oxides present in the charge.

Charging procedures. The procedures described here are specific for the
preparation of LMFR fuel solutions, but they are at the same time some-
what typical of the handling techniques necessary for other liquid-metal
fuel systems that have been suggested. Basically, the procedures result from
the need for maintaining system cleanliness, stability of additives, minimum
oxygen contamination, and uniformity of solutions.

Bismuth preparation. Bismuth ingots are cut to a size suitable for loading
and surface oxide deposits are mechanically removed. The metal is then
charged to a melt tank and heated to the charging temperature under
vacuum. Zirconium and magnesium, in the appropriate amounts, are
suspended in the melt to establish the proper concentrations of additives.
Samples are taken to verify this. When the concentrations of the additives
are stable, the bismuth is considered satisfactory for charging to the test
equipment.

Equipment charging. The bismuth from the charging vessel is forced, by
inert-gas pressure, through a porous metal filter to remove oxides, and into
a sump tank in the test equipment. From this tank the metal can be raised
by gas pressure into the operating sections of the equipment.

Addition to flowing bismuth. The addition of uranium, magnesium, and
zirconium to flowing streams is accomplished by inserting a steel basket
containing the additive into the bismuth stream through a sampling port.
Initial uranium additions to a system are not made until sampling has
shown that the concentrations of magnesium and zirconium are stable.

23-3.7 Operation and handling. Blanket gas. The blanketing of bismuth
with inert gases is necessary to provide protection against oxidation. In
many cases it has been found necessary to purify commercial grades of gas
to meet system requirements. A survey of active metals for use in the puri-
fication of rare gases has been made at Ames [42].

Several methods are in use for the determination of oxygen in gases in the
856 ENGINEERING DESIGN [cHAP. 23

 

 

 

 

 

 

 

 

 

 

Helium
[ ——
Vacuum
Compression Hollow
Seal\ L.— Sampling Rod
V
.@FI%.P.# acuum
o Atm
-« Ball g
Valve
Bubbler
1/16 In. fl
Opening

Fic. 23-7. Thief-type sampler.

ppm range of concentration. KAPL [43] and Oak Ridge [44] have developed
techniques for this analysis and commercial units have also been developed
for use in this range. At BNL, the purity of gas is checked by passing it
over a polished uranium chip at 550 to 600°C. If the chip is not tarnished,
the gas is considered suitable for use.

Conditioning operation. In addition to the system preparation steps de-
scribed in previous sections, it has been found desirable to provide a period
of system operation in which a corrosion-inhibiting layer of zirconium ni-
tride can be formed on the container walls. In general, this is done by
charging the system with bismuth to which zirconium and magnesium
have already been added and then operating the system isothermally until
analyses have shown the additive concentrations to be stable.

Sampling. Thief-type samplers have been used almost exclusively for
liquid-metal fuel systems. Sampling in this manner is accomplished by
inserting a sample tube into the metal through an airlock mounted above
the vessel. The airlock is separated from the vessel chamber by a full-
opening ball valve. By bubbling helium through a hole near the bottom of
the sample tube, it is possible to control the depth at which the sample is
taken. At the time of sampling the pressure inside and outside the tube is
equalized and the liquid enters the tube, which is then withdrawn [22].
23-3] COMPONENT DESIGN 857

  
 

Sample Cup Holder

Sample Cup vGas &
acuum

Connection

 

Section A A
Gas & Vacuum Connection

Rack Guide \fl Pinion
\ |

Rack§ 5_
QS ]  poeeree

— e g
] L 0 I Tqgooe 11 _
] \ » ‘r Sample Cup

    

 
   

  
   
  

//‘//// A

\\\‘r///

 

 

 

   

 

 

\\\‘\\\i /////// -

oY)

\ i A ~Outlet Line
Guide Support Ball Valve Water Cooling Passages

Ball Valve

Air Lock End This End Heated With Resistance Wire

Fig. 23-8. In-line bismuth sampler.

This method is shown in Fig. 23-7. A variation of this technique has been
adapted for taking filtered samples; an inverted sample cup, which has
been closed at one end by a filter, is lowered into the metal stream and filled
by increasing the system pressure. Another variation involves the use of a
sliding valve on the sample tube. This valve is opened and closed by a
rotary bellows-sealed drive that controls the time at which the sample is
taken. Radioactive samples have been taken using thief-sampler tech-
niques. The activity levels encountered were not high enough to require
remote manipulation, but drybox techniques were necessary to protect
against alpha contamination.

Corrosion study samples are used extensively in developmental systems
and consist of carefully prepared and examined metal or graphite pieces
which are included in the system piping during fabrication and removed
after each experimental run. Samples have also been inserted into flowing
streams through thief-sampler airlocks to study corrosion effects and inter-
actions between the sample and fuel stream components.

A line-type sampler, in which the liquid-metal stream is drawn through
a sample line to a sample container, is shown in Fig. 23-8. In this device,
small cups may be filled in succession and then withdrawn through the
airlock. The sampler is manipulated externally by the pinion gear.

High-temperature radiography. Techniques for radiographing operating
bismuth systems at elevated temperatures have been developed to study
plug formation, gross corrosion effects, and operating characteristics such
as liquid levels and gas entrainment. Gamma-ray sources are used in this
work [45].
858 ENGINEERING DESIGN [cHAP. 23

Repair techniques. In making repairs on systems which have contained
liquid-metal fuels it is essential to observe certain precautions:

(1) Whenever possible, the system should be thermally cold.

(2) Blanket gas should always be maintained on the inside of the system.
When the system is opened, a flow of gas from the system should be
maintained.

(3) In making welds, any surface deposit of bismuth must be removed
before a successful weld can be assured. Removal of a part of the inner
pipe wall by reaming has been found necessary. Cooling coils placed on
the pipe at the end of the reamed section will keep bismuth from melting
and flowing into the weld.

(4) In cases where bismuth fuels have undergone neutron irradiation,
proper protection against polonium contamination must be provided. It
has been found that polonium and nonvolatile fission products contained
in solid bismuth can be handled without little difficulty, since they are
largely immobilized by the bismuth. Repairs of contaminated equipment,
including welding operations, have been made without hazard [34].

23-3.8 Instrumentation. Liquid level measurement. Determination of
liquid levels in a closed metallic system, such as that generally en-
countered in liquid-metal work, can be approached either as a single-
level problem or as a continuously indicating level problem. The require-
ments for the former are:

(1) A metallic probe, preferably of the same material as the metallic
container.

(2) High-temperature insulation between the probe and the vessel in
which the liquid level is to be determined.

(3) A gastight seal between insulation and both adjacent metallic parts.

(4) An appropriate external circuit to note the attainment of the par-
ticular level.

Experience at Brookhaven National Laboratory [22] has shown that the
most successful method for providing both good insulation and a satisfactory
high-temperature seal in a single-level probe is by the use of automotive
spark plugs. It is suggested that the seal be removed from direct contact
with the heat source by means of an appropriate pipe extension. A probe
can be welded to the spark plug after removal of the bent side electrode.
The probes may be made from AISI types—347 and 502 steel for bismuth
systems or of tungsten in a tin system [32]. The external circuit consists
of a transformer, relay, and indicating lights. By the use of two probes and
interlocked relays, it is possible to indicate a level beneath the lower probe,
between probes, or above the upper probe.
23-3) COMPONENT DESIGN 859

There are two general types of continuous level indicators: a manually
adjustable resistance probe, and a variable inductance probe.

The movable probe, consisting of the proper metal rod or tube, is adjusted
through a suitable compression fitting. Modified Parker fittings [29] and
Wilson fittings with Teflon packing glands are recommended. The liquid
level is determined by comparing the probe height with a previously cali-
brated scale. |

The variable inductance probe consists of a doubly wound coil in a ce-
ramic form [22]. The coil is inserted into a pipe well inside the tank and,
as the liquid-metal level rises, the inductance of the coil changes. The
change of inductance is detected in a bridge circuit, with the degree of
unbalance being a measure of the level. This method has the advantage,
especially important in handling radioactive fluids, that the system is
hermetically sealed at all times.

If it is not possible to utilize the fluid itself for level indication, the liquid
level may be obtained in a roundabout manner by means of a stainless-steel
float. A stainless-steel tube long enough to protrude from the tank is
attached to the float. A short length of cold-rolled steel rod is contained
in the uppermost section, which is completely enclosed so that no liquid can
come in contact with it. The liquid level is obtained by locating the position
of the cold-rolled steel rod with a search coil wound about a tube concentric
with the one protruding from the tank.

Pressure measurement. Several methods are available for measuring
the pressure exerted by liquid-metal fuels. These include seal pots, gas- or
spring-balanced nullmatic transmitters, and bourdon-type gauges.

The seal pot measuring devices are simple to construct and have been
used most extensively [22,29,30] in this work. The pressure is transmitted
from the metal to a trapped inert gas that is monitoned by a conventional
gas-pressure gauge. This inert gas maintains a constant metal level in the
seal pots, as determined by means of a float [29] or spark plug probes
[22,30]. The float (with an extension rod) or the High-Low spark plug
probes actuate solenoid valves connected to gas supply and vent lines.
The probe separation is 1/4-in., thereby regulating the liquid level to 41/8
in. Since there is no barrier between metal and gas, metal may splash into
the gas space and freeze the gas lines. This may be partly alleviated by
providing long vertical gas lines, a means of heating these lines, and
baffles.

A variation consists of measuring the relative height of a column of
bismuth, backed up by gas pressure in a steel pipe [25]. The level is deter-
mined by radiography with an Ir'9? source. This method finds special
application in measuring differential pressure heads (i.e., orifice).

The gas-balanced [46,47] or spring-balanced [48] nullmatic pressure
transmitters provide a metal bellows or diaphragm seal between the liquid
860 ENGINEERING DESIGN [cHAP. 23

Inlet Gas Pressure
.Zero Adjustment Screw

   

Transformer Core

  
  
  
  

Return To . -
Pilot Relay §

 
     
     

Transformer .
\ Balancing

Nozzle

   

Capillary
Tube

 
 

 

 

     
  
    
  

O
T \
SBF

=~ R/ /1

  
 

 

 

 

 

 

 

Calibrated | Sensing
Balancing N\ ‘N Bellows NaK Fill
Spring N T
Bismuth : Tube
Secondary — Process E
Sealing Stream
Bellows Nullmatic Gas
Ballasted Type
Cooling Finsy| Sensing &
Diaphragm
Bismuth
. Bismuth Process Stream
Process Stream
Differential Transformer Type NaK Filled Bourdon Tube Type

Fic. 23-9. Pressure transmitters.

metal and a gas or mechanical pressure balance; this balancing pressure is
then measured. Figure 23-9 illustrates the basic design of three types of
these transmitters.

The nullmatic pilot-operated pressure transmitter can be made to be
very sensitive, with rapid response. A thin metallic bellows seals the unit
and is the sensing element. The full-range bellows movement is only a few
thousands of an inch. The backing gas is nitrogen and the sensing system
is adjusted to maintain a maximum differential of 10 psig across the bellows.
One of the difficulties with this type of element is the incomplete drainage
of Bi from the convolutions of the bellows. This trapped Bi may rupture
the bellows when 1t freezes. Another disadvantage is its large consumption
of instrument gas.

Another type of pressure transmitter utilizes a bellows-sealed differen-
tial transformer. The sensing element of this transmitter is similar to that
of the previous unit and consists of a metallic bellows. The very slight
movement of the bellows during a pressure change is transmitted to a
differential transformer by a rod with a secondary bellows seal. A matching
transformer installed in a bridge circuit allows a calibrated instrument to
indicate or record the actual pressure in the system.

The diaphragm-sealed, NaK-filled bourdon-tube type of pressure trans-
mitter has been used with two different styles of diaphragms. A thin,
0.010- to 0.015-in.-thick metallic diaphragm is used to separate the Bi system
23-3] COMPONENT DESIGN 861

from a NaK capillary system that extends from the diaphragm chamber
to a bourdon tube in a conventional pressure transmitter. Capillary lengths
up to 20 ft allow the transmitter to be placed remotely with respect to the
system. The other diaphragm style consists of two thin sheets of metal
welded together to form an envelope. The inside of the envelope contains
NaK and is connected to a bourdon element by a length of capillary tubing.
The envelope diaphragm is suspended in the Bi in an all-welded container.
This type of transmitter has proved to be reliable in the pressure range
between 10 and 175 psig. |

Flow measurement. Orifice. Flow of liquid-metal fuels, much like flows
of water or other liquids, is most commonly measured with standard
orifices [22,25,30]. Work done at the Engineering Research Center,
University of California [39,50] has demonstrated that an orifice may be
calibrated with water, and the calibration may then be used directly for
heavy metal (Bi or Pb-Bi) flow metering. The error introduced in this
manner is only between 3 and 59.

Orifice assemblies have generally been installed in the piping systems
with ring-joint and flange connections; one-piece orifice plate and metallic
O-rings are used. Either flange or vena contracta taps are used and the
pressure is measured as indicated in the previous section. Mild steel orifice
plates with sharp-edged holes are satisfactory for use in lead-bismuth
systems [29,30]. After 500 hr at 350°F, and a throat velocity of 1.5 fps,
there was no detectable erosion in one such orifice.

A rounded-edge orifice (with flange taps) made from AISI type-347
stainless steel gave very satisfactory service at Brookhaven National
Laboratory in a 1/2-in. IPS bismuth loop for 13,500 hr at 930°F [22]. The
flow was 5.5 fps through the throat. Upon examination, the hole diameter
had increased by 3%, (from 0.2662 in.) during loop operation.

A submerged orifice made from 219, Cr-1 Mo steel has been successfully
used at Brookhaven National Laboratory in over 4000 hr of operation with
bismuth at 1020°F. Its special appeal lies in the fact that liquid levels
(heads) instead of pressures are measured. Ordinary liquid level probes are
used.

Electromagnetic flowmeter. An electromagnetic flowmeter has been de-
signed and analyzed theoretically by General Electric Company and by
Babcock & Wilcox. A permanent magnet is mounted around the pipe
through which molten metal is flowing, with the faces of the magnet creating
a field perpendicular to the pipe. Two leads are welded to the pipe wall,
mutually perpendicular to both the pipe and magnetic flux. The emf gen-
erated by the molten metal when cutting the lines of flux is picked up by
these leads and can be transmitted to any potential-sensitive instrument.
The theoretical analysis of this type of flowmeter agrees within 69, with
experimental results.
862 ENGINEERING DESIGN [cHAP. 23

The electromagnetic flowmeter has been successfully used to meter
bismuth flows in AISI type-347 stainless steel at Brookhaven National
Laboratory [22]. The measured flow agreed within 109, with the theoreti-
cally determined value.

Preliminary results have shown that these flowmeters may also be used
in a 239, Cr-19, Mo steel system. However, corrections must be made for
the short-circuiting of magnetic flux in the ferritic steel pipe walls. One way
of minimizing this correction might be to use a bimetallic cell, that is, a
thin (0.010 in.) liner of 239, Cr-19%, Mo steel surrounded by an AISI
type—347 stainless-steel pipe to provide structural strength.

Temperature measurement. The temperature of liquid metal fuels is
usually measured with thermocouples of duplex Chromel-Alumel, No. 20
BWG gauge. Each wire is individually insulated with fiberglass and
asbestos and each pair is covered again with insulation.

The best and most accurate service in low-chrome or stainless-steel
systems is obtained by welding the thermocouple junction directly to the
outside of the pipe wall. The difference between the temperature on the
pipe wall and the bulk bismuth at 930°F is no greater than 10°F. If re-
quired, thermocouples located in wells have also been used in bismuth
systems.

In graphite systems the thermocouples are inserted in drilled holes, and
then cemented in place with alumina cement [32].

Temperature control for isothermal loops is obtained as follows [22].
The various parts of the loop are heated by means of individual heater
circuits. Since the current demand varies, depending on the position of the
heater in the loop, the current to the heaters is adjusted by means of in-
dividual autotransformers on each circuit. The entire heater group is
supplied from a single line whose voltage varies according to the signal
supplied to a controller by a single, centrally located thermocouple. The
voltage is varied by means of a transformer whose primary is in the feed
line. While the loop temperature remains within the neutral band around
the set point of the controller, the secondary coil circuit is closed. If the
temperature drops below the neutral band, the relay opens the secondary
coll circuit, thus decreasing the inductance of the primary, and increases
the voltage to the heaters. If the temperature rises above the neutral band,
the controller relay opens the main circuit breaker and cuts off current
to the heaters.

By proper adjustment of the individual Variacs it is possible to maintain
the temperatures around the loop within 20° of the desired value and to
operate so that the main circuit breakers are rarely opened.
863

REFERENCES

1. R. J. TeErreL, An Internally Cooled Liquid Metal Fuel Reactor Design, in
Proceedings of the First Nuclear Engineering and Science Congress, Vol. 1, Problems
in Nuclear Engineering. New York: Pergamon Press, 1957. (pp. 292-301)

2. T. V. SueenaN and L. D. StouguToN, The Liquid Metal Fuel Reactor
Closed-Cycle Gas Turbine Power Plant, Mech. Eng. 78, 699-702 (1956).

3. C. WirLiams and F. T. Mires, Liquid-Metal-Fuel Reactor Systems for
Power, in Chemical Engineering Progress Sympostum Series, Vol. 50, No. 11.
New York: American Institute of Chemical Engineers, 1954. (p. 245)

4. FrRaNnk W. Davis, Feastbility Study of Pressure Vessels for Nuclear Power
Generating Reactors, USAEC Report AECU-3062, Division of Reactor Develop-
ment, AEC, December 1955. (pp. 5-6)

5. C. WirLiams and F. T. MiLes, Liquid-Metal-Fuel Reactor Systems for
Power, in Chemical Engineering Progress Symposium Series, Vol. 50, No. 11.
New York: American Institute of Chemical Engineers, 1954. (pp. 245-252)

6. R. N. Lvox et al., Liqgutd Metals Handbook, U. S. Atomic Energy Commis-
sion and U. S. Navy. 2nd ed. Washington, D. C.: U. S. Government Printing
Office, 1952.

7. D. J. HuguEs and J. A. Harvey, Neutron Cross Sections, USAEC Report
BN1-325, Brookhaven National Laboratory, May 1955.

8. C. M. Davisson and R. D. Evans, Gamma-Ray Absorption Coefficients,
Rev. Modern Phys. 24(2), 79-107 (1952).

9. O. E. DwYER et al., Liquid Bismuth As a Fuel Solvent and Heat Transport
Medium for Nuclear Reactors, USAEC Report BNL-2432, Brookhaven National
Laboratory, 1955.

10. L. M. TreFETHAN, Heat Transfer Properties of Liquid Metals, Cambridge
University, England, Christ’s College, July 1, 1950.

11. S. E. Isakorr and T. B. Drew, Heat and Momentum Transfer in Tur-
bulent Flow of Mercury, in Proceedings of the General Discussion on Heat Trans-
fer, Institution of Mechanical Engineers (London) and American Society of
Mechanical Engineers, 1951. (pp. 405-409)

12. W. K. Stromquist, Effect of Wetting on Heat Transfer Characteristics of
Liquid Metals (thesis), USAEC Report ORO-93, University of Tennessee,
March 1953.

13. H. A. Jounson et al., Heat Transfer to Mercury tn Turbulent Pipe Flow,
USAEC Report AECU-2627, University of California, Berkeley, Institute of
Engineering Research, July 1953.

14. H. A. Jounson et al., Heat Transfer to Molten Lead-Bismuth Eutectic in
Turbulent Pipe Flow, Trans. Am. Soc. Mech. Engrs. 75(6), 1191-1198 (1953).

15. B. Lusarsky and S. J. KaAurMaN, Review of Experimental Investigations
of Liquid-Metal Heat Transfer, Report NACA-TN-336, Lewis Flight Propulsion
Laboratory, March 1955.

16. R. N. Lyon, Liquid-Metal Heat Transfer Coefficients, Chem. Eng. Progr.
47(2), 75-79 (1951).
864 ENGINEERING DESIGN [cHAP. 23

17. R. C. MARTINELLI, Heat Transfer to Molten Metals, Trans. Am. Soc.
Mech. Engrs. 69(8), 947-959 (1947).

18. O. E. DwyEer, Heat Exchanger in LMF Power Reactor Systems, Nucle-
onics 12(7), 30-39 (1954). |

19. R. L. MorgaN, Technical Information Service, AEC, 1952. Unpublished.

20. B. FELp and L. SziLarp, A Magnetic Pump for Liquid Bismuth, USAEC
Report CE-279, Argonne National Laboratory, 1942.

21. B. FELp, More Calculations in the Bismuth Pump, USAEC Report CP-326,
Argonne National Laboratory, Oct. 17, 1942.

22. C. J. RasEMaN and J. WEismAN, Liquid Metal Fuel Reactor (LMF R)
Processing Loops. Part I. Design, Construction, and Corroston Data, USAEC
Report BNL-322, Brookhaven National Laboratory, June 1954.

23. D. A. Wart, A Study in Design of Traveling Field Electromagnetic Pumps
for Liquid Metals, Report AERE-ED/R-1696, Great Britain Atomic Energy
Research Establishment, June 12, 1955.

24. G. R. WinDpERs and R. W. FisHER, An Electro-magnetic Pump and Heating
Transformer for High Temperature Liquid Metals, USAEC Report ISC-547,
Iowa State College, Dec. 6, 1954.

25. R. W. FisHer and G. R. WinbpERs, High Temperature Loop for Circulat-
ing Liquid Metals, in Chemical Engineering Progress Sympostum Sertes, Vol. 53,
No. 20. New York: American Institute of Chemical Engineers, 1957. (pp. 1-6)

26. R. 5. WiNGaRpD, Jr., Fairchild Engine & Airplane Corp., NEPA Division,
1950. Unpublished.

27. J. F. CorLins, Fairchild Engine & Airplane Corp., NEPA Division, 1950.
Unpublished.

28. J. E. WaALKEY, California Research Corporation, 1951. Unpublished.

29. H. A. JounsoN et al., The Design and Operation of a 30 Gpm 40 Kw Pb-Bi
Eutectic Heat Transfer System, USAEC Report AECU-2848, University of Cali-
fornia, Berkeley, Institute of Engineering Research, February 1954.

30. R. CyagaN, Circulation of Lead-Bismuth Eutectic at Intermediate Tempera-
tures, USAEC Report NAA-SR-253, North American Aviation, Inc., Oct. 1, 1953.

31. H. A. JounsoN et al., Heat Transfer to Mercury in Turbulent Pipe Flow,
USAEC Report AECU-2627, University of California, Berkeley, Institute of
Engineering Research, July 1953.

32. R. D. KEkN, High Temperature Liquid Metal Circulating System, USAEC
Report NAA-SR-985, North American Aviation, Inc., Aug. 1, 1954.

33. T. A. Simms, Fairchild Engine & Airplane Corp., NEPA Division, 1950.
Unpublished.

34. C. J. RasEMAN et al., Liquid Metal Fuel Reactor In-pile Fuel Processing
Loop (Loop B); Construction, Operation, Experimental Results, USAEC Report
BNL-403, Brookhaven National Laboratory, January 1957.

35. W. B. CorTrELL, Oak Ridge National Laboratory, 1952. Unpublished.

36. W. P. BiGLER, Reactor Engineering Quarterly Report for March 1, 1950,
Through May 31, 1950, USAEC Report ANL-4481, Argonne National Labora-
tory, July 1, 1950.

37. R. PorTER et al., Fairchild Engine & Airplane Corp., NEPA Division,
1950. Unpublished.
REFERENCES 865

38. W. J. HaLLETT et al., Dynamic Corrosion of Graphite by Liquid Bismuth,
USAEC Report NAA-SR-188, North American Aviation, Inc., Sept. 22, 1952.

39. R. A. SEBAN et al., Flow Metering of Molten Lead-Bismuth Euteclic, at
University of California, Berkeley, California. University of California, Berkeley,
Institute of Engineering Research, April 25, 1949.

40. Am. Machinist, Nov. 12, 1951.

41. O.J. ELGERT et al., Dynamic Corrosion of Steel by Liquid Bismuth, USAEC
Report LWS-24891, California Research and Development Co., Aug. 29, 1952.

42. D. S. Gisss et al., Purification of Rare Gases. I. A Comparison of Active
Metals in the Purification of Rare Gases, USAEC Report ISC-560, Iowa State
College, Dec. 30, 1954.

43. L. P. Pepkowirtz and E. L. SHIRLEY, Quantitative Determination of
Oxygen in Gases, Anal. Chem. 25, 1718-1720 (November 1953).

44. LELanDp A. ManN, Oak Ridge National Laboratory, personal communica-~
tion.

45. J. C. AustiN and P. Ricuarps, Radiography As a Hot Lab Service,
Nucleonics 12(11), 78 (1954). |

46. P. W. TayLor, Moore Pressure Transmitter Test Summary, USAEC Repor
CF-53-1-260, Oak Ridge National Laboratory, Jan. 22, 1953.

47. M. T. MorcaN, Hermetically Sealed High-Temperature Pressure Trans-
mitter and 'Hermelically Sealed High-Temperature Liquid Level Probe, USAEC
Report ORNL-1939, Oak Ridge National Laboratory, Sept. 15, 1955.

48. E. C. King and V. K. HeckeL, High Temperature Pressure Gauge, Tech-
nical Report No. 45, Mine Safety Appliances Co., Jan. 5, 1956.

49. E. A. Lueskg, Knolls Atomic Power Laboratory, 1952. Unpublished.

50. H. A. JounsoN et al., Orificc Metering Coefficients for Lead-Bismuth
Eutectic, USAEC Report AECU-2798, University of California. Berkeley, Insti-
tute of Engineering Research, December 1953.

51. W. S. FLesaMaN and C. G. Coruins, The Effect of Molten Bismuth on In-
sulating Materials, Report NEPA-1306, Fairchild Engine & Airplane Corp.,
NEPA Division, Feb. 9, 1950.

52. R. CygaN, Lead-Bismuth Eutectic Thermal Convection Loop. USAEC
Report NAA-SR-1060, North American Aviation, Inc., Oct. 15, 1954.
CHAPTER 24

LIQUID METAL FUEL REACTOR DESIGN STUDY*

24-1. CompraRISON oF Two-FLuip aAND SiNGLE-FLUID
LMFR DEgsigNS

In Chapter 18, the two-fluid and the single-fluid externally cooled LMFR
concepts were discussed in a general way. It was pointed out that the two-
fluid design has the better breeding possibilities but is somewhat more
complex than the single-fluid reactor. In this chapter a complete design
study of a two-fluid full-sized LMFR reactor is described and discussed,
and a shorter discussion of a single-fluid design study follows. This does
not mean that one design is necessarily favored over the other. In fact
both of these designs are being studied very extensively.

24-2. Two-Fruip REacTtor DEsiGN

24-2.1 General description. The two-fluid externally cooled LMFR
concept consists of a relatively small core surrounded, for the most part, by
a blanket containing fertile material. The core is composed of high-density,
impervious graphite through which vertical channels are drilled to allow
circulation of the fuel coolant. The fuel in the core is dissolved U233 or
U233 dissolved and suspended in liquid bismuth. The fluid fuel also acts as
coolant for the core system. The required coolant to moderator ratio is
obtained by proper size and spacing of the fuel coolant channels.

The blanket is constructed of high-density graphite through which flows
a liquid bismuth slurry containing the bred U233 fuel and thorium, the
fertile material. In this study, thorium is assumed to be suspended in bis-
muth as thorium bismuthide, although thorium oxide particles could be
used. The blanket is wrapped around the core as completely as possible
for good neutron economy. An important economic consideration is the
degree of end blanketing which can be achieved while keeping coolant
velocities below the allowable limit. Several blanket designs were in-
vestigated, but a complete study for obtaining the best end blanket design
has not yet been carried out.

*This chapter is based on studies made by Babcock & Wilcox Company for the
USAEC, BAW-1046, March 1958, and on a 17 company report BAW-2, June 30,
1955, for which Brookhaven National Laboratory contributed information and sup-
plementary design studies.

866
24-2] TWO-FLUID REACTOR DESIGN 867

24-2.2 General specifications. Unless otherwise noted, the specifications
listed below are common to all calculations performed in this design.

Total power 825 mw (thermal)
315,000 kw (electrical)

Coolant to moderator ratio in core, Vgi/Vc 1.22

Coolant to moderator ratio in blanket, Vaury/Ve  0.50

Core-blanket barrier material graphite

Blanket thickness 3.0 ft

Blanket slurry composition:
Bismuth 90 w/o
Thorium, as Th3Bis 10 w/o

Coolant inlet temperature 750°F

Coolant outlet temperature 1050°F

Nuclear calculations utilizing latest cross sections and multigroup diffu-
sion theory indicate that the values 1.22 and 0.50 listed above are close
to the optimum.

The several factors which dictated the choice of a bismuth-to-carbon
volume ratio merit some attention. There are some losses of neutrons due
to capture in graphite. Hence, one would wish to use only enough graphite
to sufficiently thermalize the reactor. If too little graphite is used, the
critical mass will be large. It is suspected that the n value for U233 may
be lower in the epithermal than in the thermal energy range. This would
make it desirable to keep the reactor thermal. It was found that bismuth-
to-carbon volume ratios in the range of 0.5 to 2.0 satisfy these various re-
quirements quite well. It may be further observed by referring to Fig. 24-1
that breeding improves with an increase in the bismuth-to-carbon volume
ratio. However, the maximum bismuth-to-carbon volume ratio acceptable
on the basis of structural limitations was 1.22, and consequently this core
diameter is 155.7 cm (61 in.) at a bismuth-to-carbon volume ratio of 1.22,
assuming a cylinder with its height equal to diameter.

Blanket slurry-to-graphite volume ratio and blanket thickness. A series of
calculations were made to estimate the most economical parameter values
for the blanket. Blanket slurry-to-graphite volume ratio and blanket
thickness were varied to give the best breeding ratio consistent with reason-
able bismuth holdup. Figures 24-2 and 24-3 demonstrate the effects of
varying blanket composition and thickness on breeding ratio. The slurry-
to-graphite volume ratio was set at 0.5 and the blanket thickness was set
at 3.0 ft.

Study of design parameters. The parameters investigated in the following
analysis are (1) end blanket design, (2) power fraction in the blanket, and
(3) fission product poison level in the core.
868 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

 

 

0.14 ! I P

   

 

0.10—

o
o
o
l

4

Breeaing Gain

  

o
o
o
I

N23/NBi = 6x10

0.02

] ] | | | ] | ]
0 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00 2.25
Bismuth to Carbon Volume Ratio

 

 

 

F1a. 24-1. Breeding gain vs. bismuth-to-carbon volume ratio in core.

 

Breeding Ratio
o

 

 

 

0.9 1 | |
2.0 25 3.0 3.5 40
Blanket Thickness in Ft.
F1a. 24-2. Breeding vs. blanket thickness for slurry-to-carbon volume ratio
= 1.00 and bismuth to carbon volume ratio in core = 1.00.

 

 

 

 

 

 

1.06
2
S
o
g
;a 1.05 p— =
Q
o
|

1 04 | 1 | | |

0 0.50 0.75 1.00 1.25 1.50 1.75

Slurry to Carbon Volume Ratio

Fic. 24.3. Breeding vs. slurry-to-carbon volume ratio in blanket for bismuth to
carbon volume ratio = 1.00 and blanket thickness = 3 ft.
24-2] TWO-FLUID REACTOR DESIGN 869

24-2.3 End blanket effects. A series of nuclear calculations were per-
formed to determine the effects of end blanket design upon breeding ratio
and critical fuel concentration. Two extreme blanket designs were con-
sidered. In the most optimistic case, a spherical core, equivalent to a 61-in.-
diameter cylinder, was surrounded by a 3-ft spherical blanket. The pessi-
mistic calculations assumed a cylindrical core with a diameter of 61 in.,
height equal to' 1.5 times the diameter, a 3-ft radial blanket, and no end
blanket. Critical values of fuel concentrations and breeding ratio were
calculated for four power fractions in the blanket for each design.

All calculations were performed for hot, clean conditions with an average
temperature of 900°F. A two-group, multiregion code was used to solve
the diffusion equations, and a 37-group spectral code was used to determine
the two-group nuclear constants. The results of these calculations are
tabulated in Table 24-1. The breeding ratio is decreased 0.20 to 0.25 by
completely eliminating the end blankets. This is due primarily to the
added neutron leakage out the ends of the core, despite the fact that the
core height is increased. Although the critical mass of fuel in the core is
higher without end blankets, the fuel concentration 1s somewhat lower
due to the increased core volume.

TABLE 24-1

CRITICALITY CaLcurAaTIiONS FOR Two-FLuip LMFR
WITH AND WITHOUT END BLANKETS

 

 

 

 

N23/Np; X 10° Ratio of : Blanket
Breeding :
Case blanket power : thickness, Geometry
ratio

Core | Blanket | to total power ft
I 559 152 0.0665 1.053 3.0 Full blanket
11 530 534 0.205 1.051 3.0 ” 7
I11 461 1600 0.445 1.039 3.0 ” »
IV 436 2100 0.515 1.033 3.0 ” ”
A% 403 1050 0.272 0.80 3.0 No end blanket
VI 366 2100 0.425 0.82 3.0 S 7
VII | 347 2808 0.492 0.83 3.0 o ”
VIII | 403 1050 0.272 — 4.0 non ”

 

 

 

 

 

 

 

 

 

The actual core and blanket design is between the two extremes assumed
in these calculations. The blanket can be extended beyond the end bound-
aries of the core, and a graphite reflector can cover the ends of the core
except for the coolant inlet and outlet. Cooling becomes a serious design
[cHAP. 24

LIQUID METAL FUEL REACTOR DESIGN STUDY

Control Rod Drive\fl
['—— i

870

l\?‘—L/Shielding

 

 

 
 
 

 

 

 

    
 

 

 

® o
S £
o= ]
% 3 =
XX O
5~ ®
@
//////////////////////////////////////////////// /////////////////////////
//////// // / // ////////////
AN //// DN ////// ///// // ///
AN //////// /////// ////////
ANNNNN ////// /// ////// /////////
,/%/////M/ // // ML / /
J % // N/N/NN//////// /N// ///7 //// //
_ /A/ // R R ///////
AN //A?/ RS
/ ////lfi// /////// //// //////M///V %
//////// /// //7
//// // //
DN //// NN
%M%fl/fi%fl%@fi/fl%/fi%//fié@%/
m 2 IIIHHHHHI»> 5
o
T - o o
o o M —
v 59 o g
O R S -
0 0 O
s Z
O

Nozzle

Core Inlet

Nozzle

F1c. 24-4. Two-region, externally cooled liquid metal fuel reactor.

problem, if the end reflector is replaced with blanket material. The design
in Fig. 244 is a substantial improvement over no end blanket or reflector.
However, further improvement in breeding ratio could be achieved with

even better end blanket designs.

For a given geometry, coolant-to-

24-2.4 Power level in the blanket.
moderator ratio, and thorium concentration in the blanket, specification of

the fraction of total fissions generated in the blanket establishes a unique
set of values for fuel concentration in the blanket, fuel concentration in the
core, and fissions generated in the core. For simplicity, the power generated

in a region is assumed directly proportional to the fissions in that region.

The data in Table 24-1 indicate that breeding ratio changes very little
with large changes in the fraction of total power generated in the blanket.

This increase in blanket power results in an increased ratio of resonance

to thermal absorptions, a phenomenum which tends to offset the additional

fast neutron leakage out of the blanket as blanket power increases.
24-2] TWO-FLUID REACTOR DESIGN 871

An economic analysis of the effects of changing the blanket power frac-
tion was performed to determine the optimum core-blanket power split
under equilibrium operating conditions. The parameters affecting this
choice are (1) fission-product poison levels in the blanket, (2) fission-
product poison levels in the core, and (3) chemical processing costs.

Fisston-product poisons in the blanket. The chemical processing of the
blanket slurry accomplishes two things:

(1) The removal of bred U233 from the blanket system at a rate necessary
to maintain the U233 concentration in the blanket slurry at some equilibrium
value corresponding to the desired blanket power fraction.

(2) The removal of fission products from the blanket slurry.

If the blanket processing cycle is determined by the minimum removal
rate of U233 for steady-state operation, a corresponding poison level in
the blanket is automatically set. If the blanket chemical processing cycle
is determined by the poison level and is less than the cycle determined by
the above criteria, the bred fuel removed from the blanket must be fed
back into both core and blanket to maintain steady-state fuel concentra-
tions. In this analysis the blanket processing cycle in all cases was assumed
to be based on the minimum removal rate to maintain steady-state U233
concentrations without feeding fuel into the blanket system.

Chemacal processing cycle for blanket slurry. The chemical processing was
assumed to be performed continuously on the reactor site. Unless other-
wise specified, the fluoride volatility process is utilized as described in
Article 24-3.16. The chemical processing cycle for the blanket may be
calculated [3] from the equation

ZME [+ (Zis/ Z2) (b/@)]

o - ()

 

Tp=

 

 

where
T g = blanket processing cycle, days,

Z, = removal efficiency for uranium = 0.25,
Z13 = removal efficiency for protactinium = 0.04,
M2 = mass of fuel in blanket system, kg,
b/a = ratio of Pa233 to U233 in blanket,
B = kg of fuel burned per Mwd = 1.05(1 4+ a23),
P; = total power, 825 Mw,
BR = breeding ratio,
Pp = blanket power, Mw,
872 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

and

0% (cfl)@5 + o
— TB ’

Y13

 

b
a

where

023 (eff) = an effective absorption cross section to account for resonance
and thermal absorption in U233,

% = average thermal flux over the blanket system,

Y13 = decay constant for Pa?33,

The poison level in the blanket depends upon 7'z, and T's is a function
of ME, b/a, breeding ratio, and power fraction in the blanket. All these
variables are interrelated. The ratio b/a is a function of T's, but Tz is a
slowly varying function of b/a due to the low value of Zi3/Z, (0.16).
Breeding ratio is a slowly varying function of fission-product levels in the
blanket due to the heavy loading of fuel and thorium in that region.
The breeding ratio is sensitive to the poison level, and thus to the chemical
processing rate, in the core fuel solution. An iterative calculation procedure
was required to arrive at optimum values of T, fission-product poison
level in the blanket, and the power fraction in the blanket.

For a given chemical processing rate in the blanket, the fission-product
poison level was determined from the data in KAPL 1226 [4]. Relative
poisoning, RP, is defined as the absorptions in fission products per thermal
fission in fuel, while the fission-product poison fraction is the absorptions
in fission products per total absorption in fuel. Xenon and samarium are
treated separately and are not included in the term fission products. The
burnup, F, in a region is defined as the atoms of fuel fissioned per atom
present in the region. The burnup F at time 7' in the blanket is calculated
from

__0.866 T(Pg/P) |

r
My

 

Using this relation, the relative poisoning in the blanket was determined
for each processing cycle from a graph of RP versus F' [4]. The RP curve
used is based upon high cross sections of all fission products with the excep-
tion of a low value for Zr9,

Xenon in the blanket. Xenon is removed from the blanket by the degasser.
Although the removal rate of fission-product gases cannot be determined
until experimental information becomes available, a poison fraction of 0.01
was assumed for Xe!35, |
24-2] TWO-FLUID REACTOR DESIGN 873

Samarium in the blanket. The removal rate of samarium by chemical
processing was neglected. The steady-state ratio of 2°/Z23 yusing ap-

propriate thermal absorption cross sections, is determined by the relation

S
2%

S — 142 X 107166 + 0.0126,

where ¢ = average thermal flux in the region of interest.

Fission-product poisons in the core. 'The level of fission products, FP,
other than xenon and samarium, in the core is determined by the chemical
processing cycle for the core fuel solution. The steady-state value of FP
poisons in the core should be established by an economic balance between
the value of improved breeding ratio and increased chemical processing
costs. The relationship between the core processing cycle, T'., and the rela-
tive poison, RP, in the core may be expressed as

 

d(RP) RP
dF ~— F
and
7= 0.866 TC(PC/Pt)’
M 33
where

d(RP)
dF

 

is the slope of the curve RP versus F [4],

M 55 = total mass of U233 in the core system.

The xenon and samarium poisons in the core are determined as described
for the blanket.

Economic optimization. An optimization study was performed to de-
termine the most economic power split between core and blanket systems
and fission-product poison level for the core during equilibrium operation.
The fuel cost items which vary with these two parameters are (1) bismuth
inventory, (2) fuel inventory, (3) fuel burnup, (4) thorium amortization,
(5) thorium burnup, and (6) chemical processing. Nuclear calculations
specified the fuel concentrations for both core and blanket and breeding
ratios. These values were then used to determine the chemical processing
cycle for the blanket and the pertinent costs.
874 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

2800

 

2600 —
2400 —
2200 |—
2000 —

1800 |—

N23/NBi in Blanket x 106
~ B O
© O o
S o o
|

1000 p—

o
o
o
|

600 —

400 —

200 —

 

 

 

 

Pg/Ps

Fig. 24-5. Fuel concentration in blanket vs. Pg/P; for two-fluid LMFR fully
blanketed sphere.

Nuclear calculations. The values of the parameters investigated were

RP (core) = 0.03, 0.09, 0.15,

Ps/P; = 0.10-0.50.

Since only a relative comparison was needed, all calculations were made
with a spherical core and complete 3-ft spherical blanket. The xenon poison
fraction was taken as 0.01, and the samarium steady-state value was com-
puted for each region in each case.

The fission-product poison level in the blanket cannot be determined
without first knowing the blanket processing cycle. As a first approach,
the breeding ratio for the hot clean conditions was used to determine the
cycle time from which the RP in the blanket was calculated as described
previously. The relative poison levels determined on this basis were as
follows:

 

Pg/P; RP (blanket)
10% 0.029
25% 0.048

50% 0.155
24-2] TWO-FLUID REACTOR DESIGN 875

 

 

 

 

750 | | | l
700 —
@
O
(@
o
o
x 600 —
o
Z RP=0.15
™
N
<
RP=0.09
500 ‘ —
RP=0.03
450 l | | I |
0 10 20 30 40 50 60

Pg/Py

F1a. 24-6. Fuel concentration in core vs. Pp/P; for two-fluid LMFR fully blan-
keted sphere.

All criticality calculations were performed using the specifications out-
lined in Article 24-2.2. Two-group diffusion theory was employed, and a
two-group, multiregion code was used for solving the diffusion equations.
As previously mentioned a 37-group spectral code was used to generate
the two-group coefficients. The critical concentration of fuel in the core
and blanket, breeding ratio, and neutron losses were determined for several
power splits for each relative poison level in the core. The blanket power
fraction values of 10, 33.3, and 509 were used as reference values for com-
parison, and the important nuclear parameters were determined from a set
of parametric curves for these precise values. (Cases actually calculated
corresponded very closely to the desired blanket power in most calcu-
lations.)

- The nuclear parameters corresponding to these power splits are sum-

marized in Table 24-2. Figures 24-5 and 246 show the variation of
N23/Np;i in both the core and blanket as the blanket power fraction
changes. This atom ratio of U233 to bismuth in the blanket ranges from
255 X 1076 to 2420 X 1076 for Pg/P;=0.10 to 0.50. In the core the
N23/Np; ratio decreases approximately 20% over the same range. The
[cHAP. 24

LIQUID METAL FUEL REACTOR DESIGN STUDY

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24-2] TWO-FLUID REACTOR DESIGN 877

 

 

 

2.4
T 1 I 1

" B i [ | ! | !
T i
e
X 221 —
|® ]
~
le-
e 20} —
o
|9/
£
X
2 18} _
ke
c
o
s
QO
Z 1.6} _
- RP=.03
£
o
-
= 14} _
>
g RP=.09 ]
> —_ 1
2 o . . RP_l.15

0 0.1 0.2 " 0.3 0.4 0.5 6.0

 

 

Pg/P;

F1G. 24-7. Average thermal flux in core vs. Pp/P; for two-fluid LMFR based on
a fully blanketed sphere at 825 Mw.

values of the average thermal neutron flux in the core and blanket are
graphed in Figs. 24-7 and 24-8, and BR in Fig. 24-9.

Bismwuth tnventory. The primary system volumes for Pg/P; = 0.33 and
0.50 are based on a six-loop capsule design. Each loop contains a bismuth
inventory of 245 ft3. If 50% of the power is generated in the blanket,
three loops contain blanket slurry and three contain U-Bi core solution.
If one-third of the power originates in the blanket, two loops are devoted
to the blanket system and four to the core system. If only 10% of the total
power is generated in the blanket, a three-loop design is assumed for the
core system, and two small loops of 125 ft3 each are used for the blanket.
The reactor holdup has been estimated from the reactor drawing in Fig.
24—-4. Fuel inventory volumes are summarized in Table 24-3.

Using the value of $2.25/1b of bismuth, 12% annual fixed charges, and a
density of 613.5 Ib/ft3 (9.83 g/cc), the annual bismuth inventory charges
are

C1($/yr) = 165.6 (Ves + Vi),
where

Ve = inventory volume of core system, ft3,
Vs = inventory volume of blanket system, ft3.

Fuel inventory. Five days’ holdup of fuel from both blanket and core is
assumed for the chemical processing plant. Pa2?33 is held up for 135 days to
allow for decay to U233, Approximately 3% of the Pa?33 remains after
135 days and is discarded with the fission-product waste. This loss, while
 

 

 

 

878 LIQUID METAL FUEL REACTOR DESIGN STUDY
| | | [ i

T 16} —
o

=

[- 2]

o~ 1.4 — —
1o

2

5 121 _
0

£

5

c 101 —
o

3

z

T 08} —
£

£

o RP=.03
§’ 0.6 - T
o RP=.09
< RP=.15

0.4 | I | I I
0 0.1 0.2 0.3 0.4 0.5 0.6

Pg/P;

[cHAP. 24

Fiac. 24-8. Average thermal flux in blanket vs. Pp/P, for two-fluid LMFR
based on a fully blanketed sphere at 8256 Mw.

1.03

 

|

1.02 }—

1.01

Breeding Ratio

0.96

0.95

 

0.94

:

RP=.09 |
RP=0.15

 

 

0 .10

F1a. 24-9. Breeding ratio vs.

.20 .30 .40 .50 .60
PB/Pt

Ppg/P, for two-fluid LMFR fully blanketed sphere.
24-2] TWO-FLUID REACTOR DESIGN 879

TABLE 24-3

INvENTORY VOLUMES IN Two-Fruip LMFR

 

Pp/P:=10.10 P:/P,=0.333 P:/P,=0.50

 

 

Core system:

Reactor 275 ft3 275 ft3 275 ft3

External system 1640 980 735

Subtotal 1915 1255 1010

Blanket system:

Reactor 495 495 495

External system 250 490 735

Subtotal 745 985 1230
Total 2660 2240 2240

 

 

 

 

 

 

quite small, has been included with the fuel inventory charges, which may
be expressed as

Cz($/yr)—626M23<1+ >+M <_|_ >+

(1 + 135Z13> + 30+ 132,000 Mss Z1s,
T
This equation assumes a 30-kg inventory of U233 feed material external
to the reactor. The economic assumptions used in this equation are 49,
fuel lease charges and a U233 price of $15.65/¢.
Fuel burnup. The annual cost of the net U233 fyel burned in an 825-Mw
reactor, assuming an 809, plant factor, is

Cs ($/yr) = 3.96 X 108 (1'+ a23)(1 — BR).

Thorium amortization charges. Assuming a cost of $42/kg for thorium
and an annual amortization rate of 159, based on a 20-yr life, the annual
amortization charges for the thorium are

Cs ($/yr) = 6.3 Moso.

Thorium burnup. The thorium replacement costs due to burnup are cal-
culated according to the equation

C5 ($/yr) = 10,620 (1 +a23) BR.
880

 

 

 

 

 

LIQUID METAL FUEL REACTOR DESIGN STUDY
3000 I . | T I ] T I T 1RP ]
| Blanket RP;:O9—
—-—~Core RP = .03
1000] |
0 = - ——RP=.15-
S - -
Q — =
o — —
£ — .
= - ——————RP =09
o —
-4 = 7
S ]
o | _
£
2 —
g | i
[-
100 - —
= —_— _0a
50 L2l ] i | ] | | l‘.‘—lkp _103 ]
0.10 .20 .30 .40 .50 .60

Pg/Py

[cHAP. 24

Fiag. 24-10. Chemical processing cycles vs. blanket pOWer, based on a blanketed
sphere with total reactor power of 825 Mw and the removal efficiencies of Z, = 0.25,

Zy3=0.04, 285 = 0.10, Z§p = 1.00.

5 x 10°

106

Annual Processing Cost, $

102

 

TR

 

T T oo Tt

Total Annual Cost

Capital Equipment

Operating Cost

 
    
 

Building Cost

L Ll | L o Lot bbbl

    
 
   
 
 
 

 

 

1

3 6 10 30 60 100
Plant Liquid Throughput, Cu Ft Per Day

200

Fic. 24-11. Annual fluoride volatility processing cost vs. plant throughput for
825-Mw-two-fluid LMFR.
24-2] TWO-FLUID REACTOR DESIGN 881

 

 
 
   
 
 
  

 

 

 

8x10 1 | T VT T T 7 ' LI L L L R B
6 |- L
//
4+ -~
-~
v - - 4
U; — // /:
> -7
S ] Total Annual Cost /’ 7
m 2 [ —
-E //’
a e ]
8 = e
o Capital Investment Return
a6
3 F T B
2 s b - - Operating Cost _
£
g L. P -
6 —— ]
4“05] e b bbbl : L Lot b Ll

3 6 10 30 60 100
Plant Throughput, kg Thorium Per Day

F1a. 24-12. Annual aqueous processing costs vs. plant throughput for 825-Mw-
two-fluid LMFR.

Chemical processing costs. The chemical processing cycle time for the
blanket is determined by the Pp/P; ratio and the breeding ratio, as dis-
cussed in previous paragraphs. The processing rate for the core system is
determined by the method also described previously; see Fig. 24-10. The
total throughput to the fluoride volatility chemical separations plant is
simply:

Vbs

Throughput (ft3/day) = Tg

  

The annual processing charges based on fluoride volatility can be read
directly from Fig. 24-11, a plot of annual charges versus plant through-
put.

As a matter of comparison, the chemical processing charges were also
computed for each case, assuming on-site aqueous processing methods.
The capacity and cost of an aqueous processing plant are determined by
the amount of thorium per day which must be processed. The core solu-
tion processing does not enter into the cost unless the ratio of fuel to thorium
presents criticality problems in the process equipment. This situation is
likely to occur for the higher power levels in the blanket. This analysis did
not take this possibility into account, however, and annual aqueous
processing costs were taken directly from Fig. 24-12. This design plant
capacity is 35 kg/day of thorium feed.

Results of optimization. The bismuth inventory is shghtly greater for
the case of Pg/P; = 0.10 than for the other two cases, because of the added
primary system volume. Fuel inventory charges are not very sensitive to
882 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

 

 

  
   
      

 

 

 

.40
| l l l I
361 RP=.157]
RP= .09

32 RP= 03]
s .28 —
X Fuel Inventory
2 24— —
3
a 20— ———— . —
3 :
g Bi Inventory
2
=121 RP =.15—
3 Thorium Charges
s 08} RP = .09 |
©
* 04l RP = .03—

Fuel Burn-Up
—.04} —
—.08 | | | ] |
0 .10 .20 .30 .40 .50 .60

Pp/P,

Fi1c. 24-13. Relative fuel costs vs. blanket power for two-fluid LMFR based on a
fully blanketed sphere operating at 825-Mw with a plant factor of 809%.

2.00

 

1.80}—

1.60 —

Aqueous On-Site

1.40 —

1.20

1.00}—

Chemical Processing costs, Mills/KWH

0.80 |—
Fluoride

Volatility
RP = 0.09

 

 

0.40 | | | I I
0 0.10 20 30 40 50 60

 

Fig. 24-14. Chemical processing costs vs. blanket power. The cycle times are
based on a blanketed spherical reactor with a total heat power of 825-Mw.
24-2] TWOQ-FLUID REACTOR DESIGN 883

 

    
  
 

 

 

 

24
| I I | l
22 |- —
X 20 - _
3
x
%
218 |- —
% .
o On-Site Aqueous
§ 16 I Processing _
M.
3 RP =.03
5
2 1.4 | RP =.15-
RP = .09
RP=.15
Fluoride RP = .09
1.2 - Volofili.fy RP=.03 |
Processing
o | | | | |
0 .10 .20 .30 .40 .50 .60
PB/Pi

Fia. 24-15. Relative fuel costs vs. blanket power for a blanketed spherical re-
actor operating at a total power of 825 Mw with a plant factor of 809.

the relative poison level in the core, but they increase sharply with an in-
crease in power level (Fig. 24-13). Thorium charges increase linearly with
blanket system slurry volume, and fuel burnup charges increase as Pg/P;
increases, as shown in Fig. 24-13.

Chemical processing costs drop rapidly as the power fraction in the
blanket increases. The increased processing rate required to maintain a
steady-state fission-product relative poison level in the core of 0.03 results
in a processing cost much higher than required for RP values greater than
0.09. The aqueous processing costs appear to become essentially equal to
fluoride volatility costs at a value of 509, for Pp/P;. Further analysis
would be required to determine the validity of the aqueous processing cost
curve for low throughput and high Ng23/Ng2 ratios encountered in the
cases of high blanket power. The chemical processing costs are tabulated
in Table 24—4 and shown graphically in Fig. 24-14.

The results of the economic comparisons are summarized in Table 24-5
and are graphed in Fig. 24-15. (RP on the graphs refers to the relative
poison level of the fission products in the core.) Figure 24-15 shows that
for all values of RP a minimum fuel cost occurs for a Pg/P; of approxi-
mately 0.33.

24-2.5 Selection of a reference design. The optimization study indi-
cated that the most economic reactor design should produce one-third of
[cHAP. 24

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886 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

the total power in the blanket system and that the relative poison in the
core due to fission products should be approximately 0.09. However, sev-
eral effects must be considered in relating the optimum reactor to the ac-
tual operating reactor. A geometry more realistic than the fully blanketed
sphere must be considered in establishing new specifications; effects of
higher uranium isotopes, Pa losses, and control rods on breeding ratio must
be taken into account; and a new chemical processing cycle for the blanket,
along with a new fission-product poison level in the blanket, must be cal-
culated based upon the adjusted breeding ratio.

Geometry effects. The inability to wrap a blanket around the ends of
the core requires an adjustment to the parameters for the reference design
based on the calculations with a full blanket. The axial leakage out of a
bare ended core and a blanket with a height 1.5 times its diameter was cal-
culated to be 0.18 neutron per absorption in fuel. An extension of the
blanket length and the addition of partial end graphite reflectors are esti-
mated to reduce the end leakage to one-half this value. The total neutron
leakage, both fast and thermal, out of the partially blanketed reactor is
estimated at 0.17 neutron per absorption in fuel.

The added length of core and blanket will slightly increase the critical
mass, but the required N23/Ng; ratio will decrease slightly. In order to
be conservative in the fuel inventory costs, however, the critical values of
N23/Ngi for the fully blanketed sphere are assumed for both core and
blanket.

Breeding ratio. Higher uranium isotopes. The higher uranium isotopes,
primarily U234 U235 and U236, continue to build up in both the core and
blanket fuels throughout reactor life, since they cannot be separated in the
.chemical plant. The relative poison due to these isotopes, however, rises
rapidly at first with the buildup of U236 but increases very slowly there-
after. The return from U235 fissions almost balances for losses to U%3* and
U236 [4]. An average poison fraction of 0.01 for the reactor is used for the
reference design.

Protactinium losses. The equilibrium Pa?33 concentration can be com-
puted from the relationship

b

NE="
13 a

AT B
N23)

using an effective thermal absorption cross section of Pa?33 based on the
calculated neutron spectrum in the blanket. The relative absorptions of
the Pa233 are very small (0.005), but they are included. é

Control rods. The self-regulating properties of an LMFR have not been
established at this time. An allowance of 0.01 in relative absorptions is
included to account for the possibility of using a regulating rod and a small
24-2] TWO-FLUID REACTOR DESIGN

REFERENCE DESIGN SPECIFICATIONS

SPECIFICATIONS FOR KEQUILIBRIUM OPERATION

Core:
Thermal power
Electric power
Diameter, inches
Height, inches
Fuel
Vei/Ve
N23/Nsi
Mass of U233 in system, kg
Total volume of fuel, ft3
Breeding ratio, over-all
Chemical processing cycle, days
Volume flow rate through chemical plant, ft3/day
Mass flow rate through chemical plant, g U233 /day
Average thermal flux in active core
Average thermal flux in core system

Blanket:
Thermal power

Electric power
Thickness, ft

Vslurry/ VC
Slurry content:
Thorium (as Th3Bis) 10% wt
Bismuth 90% wt
N23/Np; (atom ratio) 1190 X 106

Mass of U233 in system, kg

Mass of thorium in system, kg

Total volume of fuel, ft3

Chemical processing cycle, days

Volume flow rate through chemical plant, ft3/day
Mass flow rate through chemical plant, kg of Th/day

 

887

550 Mw
210,000 kw
61

91.5

U233

1.22

600 X 10~%
234

1255

0.86

446

2.81

525

1.6 X 1012
6.4 X 1013

275 Mw
105,000 kw
3

0.5

328
27,900
985
200
491
140

amount of shim control for normal operation. Safety rods are included in

the reference design but do not affect neutron economy.

Fission-product porsons. The adjustment of breeding ratio to correspond
to the effects outlined above changes the required chemical processing
cycle for the blanket system. This change in 7' also changes the equilibrium
888 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

value of fission products in the blanmket. Proper adjustments result in a
blanket processing cycles of 200 days (assuming Z, = 0.25) and a fission-
product poison fraction in the blanket of 0.039 (RP in blanket = 0.15).

Neutron balance. The neutron losses proportional to one absorption
in U233 are listed below:

Absorptions in: U233 1.000
Th 0.860

C 0.025

Bi 0.050
Xel3d 0.010
Sm!49 0.017
Fission products 0.073
Higher isotopes 0.010
Control rod 0.010
Pa233 0.005
Leakage 0.170
Total 2.230

24-3. SyYSTEMS DESIGN

24-3.1 General. Systems design covers all of the reactor plant external
to the reactor, except for chemical processing. The reactor plant includes
the steam generator, but not the steam system or its auxiliaries. The
principal purpose of the systems is to transport heat from the reactor and
generate steam. They also provide supporting functions, such as shield
cooling, uranium addition, etc.

The primary system consists of six heat transport loops, each consisting
of a pump, a heat exchanger, check valve, and interconnecting piping. The
hot-leg temperature is 1050°F; the cold-leg temperature 750°F. In each
of the intermediate heat exchangers, heat is transferred from the bismuth
to the intermediate fluid, sodium. There are six intermediate heat transport
loops, each containing a pump, steam generator, and interconnecting
piping. The hot-leg temperature is 1010°F; the cold-leg temperature
680°F. Steam is produced at 2100 psia, 1000°F.

Selection of the above parameters was a problem involving consideration
of the steam plant as well as the reactor plant. The primary system
temperatures were first fixed by using the largest AT considered likely
to prove practical.

The temperature approach of the intermediate heat exchanger was set at
40°F, resulting in a sodium hot-leg temperature of 1010°F. To provide the
close approach necessary for steam temperature stability, the steam
temperature was set at 1000°F. A steam pressure of 2100 psig was picked
to correspond with 1000°F.
24-3] SYSTEMS DESIGN 889

Shifting the sodium cold-leg temperature redistributes heat-transfer
surface between the intermediate heat exchanger and the steam generator.
However, it seems desirable to favor making the intermediate heat ex-
changer small to cut down on fuel inventory. For this reason, the sodium
cold-leg temperature was established at 680°F.

24-3.2 Plant arrangement. Plant arrangement starts with positioning
the primary system relative to the reactor, and this is determined by seven
principal considerations: (1) reactor design, (2) plant operation, (3) main-
tenance, (4) operational limitations of major components, (5) structural
integrity of piping, (6) economics, and (7) safety.

A preliminary analysis of the two reactor concepts, single-fluid and
two-fluid, resulted in the decision to use three external loops for the single-
fluid and six for the two-fluid reactor. For both these alternates the main-
tenance philosophy selected was that of removal and replacement by hori-
zontal transfer of a complete primary loop upon failure of any major
component in the loop [5]. Thus, for arrangement purposes, the primary
loops assume the shape of a rail-mounted horizontal containment vessel,
or capsule, sized to contain all loop components. The height of the capsules
relative to the reactor is dictated by an economic balance between height
or elevation costs and pump net positive suction head.

The arrangement for the two-fluid reactor with six primary loops is
shown in Figs. 24-16 and 24-17.

In plan, the primary loops were located radially around the reactor,
Fig. 24-16. A minimum length of interconnecting pipe between the reactor
and the loops was used because of high fuel inventory costs. This latter
consideration ruled out shielding of any appreciable thickness between the
reactor and the loops. Maintenance access doors and other shielding
around the outside of the loops was sized for source conditions 6 to 8 hr
after shutdown of the reactor to permit access by maintenance personnel
at that time into the annular area.

With the primary loop arrangement established, the next problem was
location of the intermediate system. Since this system is the connecting
link between the primary systems and the steam turbines, it must be
located between them. The turbine is above ground level for gravity drain-
age of condenser cooling water, and the primary loops are below ground
level for economy of shield costs. The path taken by the intermediate
system can be either a high-level path, immediately up from the primary
system, or a low-level path, immediately down from the primary system,
and then horizontally to an area outside the primary system area.

The intermediate system in this arrangement follows the high-level
route to the steam plant. Sodium lines are brought straight up to an
annular area around the reactor maintenance chamber. Since access to
890 LIQUID METAL FUEL REACTOR DESIGN STUDY [cEAP. 24

 

- o

‘ Turbo-Generator Building (
Sodium Cold

Dump ; Component

Storage

    

 

' Containment

       
 
   

  
  
 
 
 

Machine
Shop

v

  

@o
&V'
\Q
&
&
&

/

\
-
1

  
    

  

Decontamination
Area

Sodium
Dump
Tanks

   
    
  
 
  

 

p and
fiVenhlahng

Stack : System .

Removal
and Equipment Storage

L L ]

First Floor Plan

Fic. 24-16. LMFR-6: Capsulate loop conceptual plant layout.

Reactor

 

Pyroprocessing and Aqueous

this chamber will not be permitted during reactor operation, a heavy shield
wall is not required around the chamber.

Within this annulus are the sodium pumps and the steam generators.
Final layout of this equipment will require considerable ingenuity, but it
is feasible. Steam lines will cross the roof of the reactor building to the
turbine building.

Because the primary loop hot maintenance shop for this concept serves
such specialized functions, its usefulness for maintenance of chemical
processing equipment is doubtful. Accordingly, the chemical processing
facilities for this two-fluid six-loop plant, together with its supporting hot
and conventional laboratories, fuel addition and other systems, are located
in a separate building.

The turbine building is of conventional construction and will be in all
essential respects identical for both plants.

Startup heating switch gear, gas heating and cooling systems for the
reactor and dump tanks, inert gas storage systems, control rooms, and
other auxiliaries are located relative to the above systems as logically as
possible in the light of their functional requirements.

With respect to contamination control the basic philosophy is (1) con-
trolled access to areas having different order of magnitude activity levels
and (2) controlled circulation of ventilating air to assure flow from low-
24-3] SYSTEMS DESIGN 891

Reactor
Service
Capsule Area

Sodium Area — Sodium
Systems v Systems Shielding Windows

Mechanical Arm

5 Ton Crane

   
  
   
  
 
 

 

Rails

/Crane Rail

 

 

 

 

 

et T T

 

 

  

Dry Hot
Storage

 

Tank Pit \Transfer Car

Magnesium Sump
Fuel Dump Room

Wash  Tanks

Fia. 24-17. LMFR-6: Capsulate loop conceptual plant elevation.

to high-level activity areas. For guidance in achieving these objectives
a rough scale of activity levels has been proposed, as follows:

Class # 1—conventional steam turbine plant, personnel monitored.

Class # 2—uncontaminated areas of nuclear plant, personnel monitored.

Class # 3—potentially contaminated areas, personnel closely monitored;
e.g., shield cooling, reactor and dump tank heating and
cooling, hot shop operating area.

Class #4—low activity, accessible by closely monitored personnel only
under favorable conditions; e.g., exhaust blower room, hot
chemical laboratory.

Class # 5—medium to high activity, accessible by closely monitored
personnel only after executing standard decontamination
procedures; e.g., hot maintenance shop.

Class #6—high activity, no access during life of plant except after
extended shutdown and special decontamination; e.g., chem-
ical processing and chemical hot cell.

Class # 7—very high activity, no access by personnel during or after
life of the plant; e.g., primary loop and reactor areas.

24-3.3 Primary system. The LMFR primary system is designed to re-
move up to 825 Mw of heat from the reactor. The primary system consists
of six separate heat transport loops.

The fuel stream enters the bottom portion of the reactor vessel at a
minimum bulk temperature of 750°F, and flows upward through the core,
where fissions within the fuel cause the fluid to undergo a temperature rise
of 300°F, resulting in a maximum fuel temperature of 1050°F. Upon
leaving the core, the fluid passes upward to a degassing area, where volatile
fission products are removed from the fuel stream. The reactor discharge
consists of a header which splits the fuel flow into the primary heat-
transport loops.

.
892 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

The primary loop piping, 20 in. in diameter, is sized to obtain a maximum
fuel velocity of 10 fps.

From the degassing area discharge, each fuel stream flows to the suction
of a variable speed centrifugal sump type pump. Each pump is de-
signed to deliver about 9000 gpm at 20-ft head of pumped fluid. To obtain
a reasonable pump speed, the net positive suction head requirement is
11.5 ft. A gas pressure (helium) is maintained over the pump sump to
prevent flooding of upper parts (motor windings, cooling system, etc.) of
the pump.

From pump discharge the fuel stream flows to the tube side of a U-tube
U-shell intermediate heat exchanger in which the fuel stream gives up heat
to the intermediate fluid, sodium. Upon discharge from the intermediate
heat exchanger, the fuel solution flows into the reactor.

To meet safety requirements, the reactor and the major components of
the primary loops are enclosed within containment vessels. The contain-
ment vessel which houses the pump and heat exchanger of each primary
loop is a cylindrical capsule, 20 ft in diameter by 30 ft long, including 2:1
elliptical heads. The capsule is equipped with access holes such that cer-
taln maintenance operations may be performed [5]. The reactor con-
tainment vessel is a right circular cylinder 30 ft in diameter, with a hemi-
spherical top head and a flat bottom head. Access holes are provided in the
vessel for maintenance operations.

Each heat-transport loop is provided with four dump tanks which re-
ceive the loop and a portion of the reactor volumes. The tanks are sized and
arranged to prevent a fast chain reaction. The primary loops are filled from
the dump tanks by means of small electromagnetic pumps. These pumps
also promise a means for agitation of the fuel.

Two dump lines, each with a maintainable valve, connect each loop with
the dump tanks.

The bismuth charge system consists of a bismuth melt tank, filter,
valves, and piping to the dump tanks.

The proposed material of construction exposed to primary fluid is 21%
Cr-19% Mo steel.

24-3.4 Intermediate system. The intermediate system, which also con-
sists of six separate heat-transfer loops, utilizes sodium as the heat-transfer
medium. All material of construction of the intermediate system, except
the steam generator, is 21% Cr-19% Mo steel. The steam generator, which
is designed for high-pressure, high-temperature service, is constructed of
type-304 stainless steel. The intermediate piping (12-in. schedule-30) is
sized for a maximum sodium velocity of 17 fps.

Sodium flowing at 11,000 gpm enters the shell side of the intermediate
heat exchanger (which is a U-tube, U-shell unit containing 2400 5/8-in.-OD
24-3] SYSTEMS DESIGN 893

tubes with an average length of 21 ft) at 680°F, flows countercurrent to the
fuel stream, and exits from the heat exchanger. Sodium flows to the suction
of a variable speed centrifugal sump type pump. Each intermediate pump is
designed to deliver 11,000 gpm at 180-ft head.

From pump discharge sodium flows to the shell side of the steam gen-
erator. The steam generator is a U—-tube, U-shell “‘once-through’ type
unit which is constructed of type-304 stainless steel. The unit consists of
530 1/2-in.-OD tubes with an average length of 65 ft. The shell OD is
29 in. and the over-all length is 68 ft.

Sodium flows countercurrent to superheated steam, boiling water, and
feedwater in the steam generator and gives up heat which produces
1,100,000 1b/hr of superheated steam at 2250 psig and 1000°F.

From the steam generator sodium flows to the intermediate heat ex-
changer inlet to complete the cycle.

In addition to the components listed above, auxiliary components are
necessary to obtain proper function of the intermediate system. An ex-
pansion tank is located at the highest point of each intermediate loop. This
tank serves as a cushion for pressure surges, a surge vessel for thermal ex-
pansion of sodium, and suction head for the pumps. The lowest point of
each intermediate loop is connected by pipe and dump valves to a sodium
dump tank which receives the inventory of the respective loop. Sodium is
replaced in the loop by a small electromagnetic pump which takes suction
from the bottom of the dump tank. A plugging indicator and a cold trap
are provided to determine sodium oxide concentration and to maintain
the oxide concentration at low levels.

In the event fission-product “hangup” occurs in the intermediate heat
exchanger, fission product activity will generate heat within the metal.
To prevent excessive metal temperatures, cooling must be provided when
the unit is drained. This cooling is accomplished by providing removable
sections of insulation which, when removed, will permit heat to be dissi-
pated by radiation, conduction, and convection heat transfer. Flow control
of the intermediate system will be by the variable speed pump drives. This
method of control should provide a reasonably constant steam temperature
and pressure.

24-3.5 Reactor heating and cooling system. The reactor must be pre-
heated prior to operation and for outgassing purposes. The required tem-
perature for outgassing the graphite is 1000°F. To achieve preheating,
hot helium gas will be circulated through the close-fitting jacket or double
containment which creates an annulus surrounding the reactor vessel.
During the preheating phase, helium gas will be pumped from one of two
blowers, pass through an electric resistance heater, be introduced at the
bottom of the annulus, pass up around the reactor vessel giving up its
894 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

transported heat to the cooler surface, and return by ducting from the
upper end of the containment to the blower suction.

When, for any reason, it becomes desirable to shut down the reactor and
dump the primary system, reactor cooling must be provided to remove
decay heat generated by fission-product hangup within the reactor after
dump. This is necessary to prevent internal temperatures from exceeding
design limits. The system as just described provides cooling by opening a
valve to bring a finned tube helium-to-water heat exchanger into the cycle
and by closing a stop valve to remove the gas heater from the gas flow
path.

Helium system design pressure and temperature will be 5 psig and 1050°F.
The entire loop is of all-welded construction to minimize helium leakage
and leakage of volatile fission products should a rupture of the reactor
vessel or piping give volatile fission products access to this loop.

24-3.6 Dump tank heating and cooling. When fuel is drained from the
primary system into the dump tanks, fission-product decay produces heat
within the fuel which must be removed to prevent dump tank metal
temperatures from exceeding design limits.

Cooling 1s accomplished by circulating helium at 140 psig through a nar-
row annulus around each dump tank. Helium which has removed heat from
the dump tanks passes through a finned-tube heat exchanger and gives up
heat to river water. Circulation of helium is provided by six 14,000 cfm
blowers, each rated to provide a head of 18-in. water. Three standby
blowers are also provided. Helium piping is arranged such that four dump
tanks are serviced by one blower.

To preheat the dump tanks and to maintain their temperature at a level
such that fuel precipitation does not occur, electric heaters are paralleled
with the heat exchanger such that the same piping system serves for heating
or cooling. The heaters or heat exchangers may be brought into or taken
off the cycle by valving. |

24-3.7 Startup heating system. Prior to power operation, the LMFR
heat-transport system must be preheated to about 800°F. The reactor and
the primary dump tanks are preheated by electric furnaces and circulating
helium. The remainder of the heat-transport systems, i.e., primary pipe,
intermediate heat exchanger, intermediate piping, dump tanks, expansion
tanks, steam generator, and the steam system pipe and components, are
preheated by induction heaters.

Since 21 Croloy and stainless steel are nonmagnetic, a thin sheet of car-
bon steel will be required under areas where induction heaters are applied.
24-3] SYSTEMS DESIGN 895

24-3.8 Primary inert gas system. Inert gasis used in the LMFR primary
system to cover all free liquid metal surfaces and to provide a gas seal
within the pumps. |

Helium, by virtue of its very low activation cross section and inertness,
is utilized as the cover and seal gas for the primary system. It is stored
at 200 psig in a storage tank and is piped via pressure-reducing valves to
the pump, dump tanks, and reactor. Since relatively small quantities of
helium will be used, it is expected that waste helium will be discharged via
the off-gas system to the stack.

Since commercial helium is sufficiently pure for use in an LMFR, no
purification will be required.

24-3.9 Intermediate inert gas system. Nitrogen is used in the LMFR
intermediate system to cover all free sodium surfaces and to provide a gas
seal in the pump. It is stored at 200 psig in a storage tank and is piped via
pressure-reducing valves to the pumps, expansion tanks, and the dump
tanks. Used nitrogen is discharged to the stack.

Commercial nitrogen must be purified prior to use in the intermediate
system. Purification is accomplished by bubbling nitrogen through several
tanks containing NaK.

24-3.10 Shield cooling. The concrete surrounding the primary cells
serves as a shield from the neutrons and gammas leaving the primary fluid.
In the absorption of these neutrons and gammas, considerable heat is
generated within the concrete. To hold temperatures and thermal gradients
within the concrete to reasonable limits, a cooling system must be utilized.
This cooling system consists of panel coils embedded about 6 in. within the
concrete shield. High-purity water, flowing inside the panel coils, removes
heat from the concrete and prevents temperature damage to the concrete.

The closed, high-purity loop which rejects heat to river water is designed
for a maximum heat load of 6 Mw. One pump of 900-gpm capacity pro-
vides circulation for the closed water loop. Flow control valves proportion
the flow to the various panels such that panel coil outlet temperatures are
equal.

A dump tank for the closed loop (about 300 ft3) is located beneath the
panel coils, so that the coils may be gravity drained. Water 1s returned to
the closed loop by means of gas pressure. In the event 1t 1s necessary to
dispose of the water in the closed loop, it may be drained from the dump
tank to the radioactive waste disposal system.

24-3.11 Reactor cell cooling. Instruments located within the reactor
containment vessel must be kept at a relatively low ambient temperature.
To maintain the ambient temperature, a “fin fan” cooling unit is attached
896 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

to the containment vessel. Helium, which fills the containment vessel, is
circulated by a blower located within the cooling unit. The circulating
helium removes heat from the containment vessel and transports it to the
finned coil, where it is transferred to water which is taken from and returned
to the closed shield cooling circuit.

24-3.12 Capsule and reactor room cooling. The containment capsules
and the reactor are located in a large room. The ventilation requirements
of this area are dependent upon heat losses from the primary loop contain-
ment capsules.

Ventilation is provided by locating air intake louvers at several points
around the room. An air fan provides circulation of air around the capsules
and removes heat, which is discharged to the stack. A radiation monitor
continuously monitors the air discharge. In the event radiation tolerance
levels are exceeded by the air discharge, the cooling air will be recirculated
to the reactor and capsule room until the source of radiation is determined.

24-3.13 Raw water system. The raw water system is the final waste
heat sink for the entire plant. River water, which is screened and treated,
is piped beneath the turbine-generator building. The systems which require
river water, i.e., turbine condenser, shield cooling, reactor cooling, dump
tank and pump cooling, take suction from this pipe and discharge to a
similar one which returns the heated water to the river. Where possible,
river water flows tube side in heat exchangers, to facilitate cleaning.

24-3.14 Instrumentation and control. The purpose of the control system
in this plant is to provide safe and stable operation while following the
loads imposed by the utility system. The plant follows the turbo generator.
Loads on the turbo generator are set by the utility.

A load change will appear in the steam system as a change in throttle
valve position and, therefore, a change in steam flow and pressure. The
feedwater controllers at the inlet to the steam generators will sense these
changes and operate to maintain steam pressure constant. The steam flow
could also provide an anticipatory signal to the primary and intermediate
system pumps to change their speed to suit the load.

The reactor will have a negative temperature coefficient of reactivity.
Thus, it will try to maintain its average temperature constant during load
changes. The temperature will change from time to time as reactivity
changes. To take advantage of the negative temperature coefficient, the
average temperature of the reactor will be set at a constant value.

Programming of flow rate in the primary and intermediate loop is un-
certain. Cost estimates for pumps and control equipment were based on
the premise that speed of the pumps would be varied. This might be neces-
sary to avoid thermal stresses during transients.
24-3] SYSTEMS DESIGN 897

   
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 
   
 

 

 

 

Blanket Pa D Tank
-—_Liquid a Decay Tanks
Salt
Core _ o HF
Liquid o | Hydro
Fluor'n’r |  Salt
To NFPN__ Y o
Process -
Waste
From NFPN F
Process 2
Fluor'n'r
_ Blanket Je
- UFs
quUld Return H2 Reduct'n
Mg, Zr, U
Salt + UF4 l

 

 

 

 

> Electr'y’c
Core Reduct'n
‘Liquid Return

 

 

 

Fig. 24-18. Fluoride volatility processing of core and blanket.

24-3.15 Maintenance. The maintenance of the reactor and primary
system components will be completely remote, because of the high levels of
radioactivity of the circulating fuel. The entire plant and reactor system
are arranged for remote maintenance [5].

24-3.16 Chemical processing. The pyro process chosen for this economic
study is the fluoride volatility method applied to a two-region reactor.
Work of adapting this process to bismuth fuel processing is presently under
way at Argonne National Laboratory. Figure 24-18 presents the main
steps in this process. As shown, the process may be used for either blanket
or core liquid. When the plant is processing core liquid the basic steps in
this process are (1) hydrofluorination to oxidize uranium and some fission
products, (2) transfer of the oxidized material to a fused salt phase, (3)
Auorination of the salt carrying the uranium and fission products for sepa-
ration of uranium as volatile UFg, (4) reduction of the UFg to UF4 by
H, in a fused salt phase, and (5) reduction of UF4 to uranium metal and
transfer into the metal phase (bismuth).

The volatility method can be conveniently used to process a thorium
bismuthide blanket. The process must be preceded by a phase separation
step which separates the thorium bismuthide solids from the liquid carrier
bismuth (Fig. 24-19). The modification of the core liquid process flowsheet
is as follows: (1) salt effluent from the hydrofluorination step must be
stored in order to achieve Pa decay to uranium, and (2) the bismuth liquid
is returned to the blanket head end process without the addition of uranium.
898 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Stripped Bi
_ [ From Volatility Plant
Th
Slurry \
ene————- .
Mixer
Feed Slurry
I Recycle
%‘ Heat
Pulser
Crystallizer
Phase
Separator
Slurry Return
. To Blanket
U Rich
Bi Liquid

F1a. 24-19. Head end processing, bismuthide slurry.

Certain of the fission products are not removed by volatility processing.
These may be removed by zinc precipitation (Fig. 24-20). This process
requires that the bismuth feed be free of uranium, and the volatility plant
provides such a bismuth feed stream.

The head end process transfers bred uranium, protactinium, and fission
products out of the solid phase portion of the slurry and into the liquid
phase. After this step the two phases are partially separated. A liquid
portion transferred to the volatility plant carries bred uranium, protac-
tinium, and fission products with it for stripping with HF. The stripped
liquid bismuth is returned to the head end plant for mixing with fresh slurry
feed. The head end process is not 1009% efficient; i.e., the uranium and
protactinium are not completely removed from the slurry before reconsti-
tution and return to the blanket region. This problem has been examined
in some detail and was taken into account in determining economics.

24-3.17 Turbine generator plant. A flow of 3,330,000 1b/hr of super-
heated steam at 2100 psi and 1000°F is delivered to the turbine. The
generator has a gross output of 333,000 electrical kw, and the condenser
removes 1.677 X 10° Btu/hr at 1.5 in. of mercury absolute, thus giving a
gross heat rate of 8450 Btu/kwh. About 18,000 kw of electrical power is
used for the various pumps and auxiliary systems in the plant, making
the net output 315,000 kw. Therefore, the net heat rate is 8940 Btu/kwh,
which corresponds to an efficiency of 38.2%.
24-3] SYSTEMS DESIGN 899

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

From Zn Rich, NFPN
Volatility Plus Bi
Plant
Zinc
C - .
Bi Rich | ~trator Zinc
Plus Zn \ ‘ Waste
Zn
Zn
Crystal—
lizer
Still

 

 

 

I S i
|

_ Bi Return To J

"~ Volatility Plant

 

 

 

Fic. 24-20. NFPN fission-product removal.

At full load there are 1,825,500 lb/hr of steam leaving the turbine and
being condensed in the condenser. Also, 113,800 lIb/hr of water from
various parts of the cycle are being cooled by the condenser. The total
head load on the condenser is 1.677 X 109 Btu/hr. The condenser cooling
water enters one water box at 70°F and leaves the other at 80°F.

24-3.18 Off-gas system. The actual design and efficiency of any con-
ceptual degasser are as yet unknown quantities, and a knowledge of these
important details will have to wait until in-pile loops have provided sufti-
cient data.

The off-gas system will consist of a cooler followed by a series of storage
bottles. Gaseous fission products that have been separated from the liquid
bismuth in the degasser are first sent through a cooler which offers a resi-
dence time of about a day, or enough for most of the short-lived isotopes
to decay. From the cooler, the gasses are compressed into storage bottles,
each capable of holding 30 days’ accumulation. The storage bottles will
each be 4.25 ft3 in volume, and at 212°F and 60 psia at the time of dis-
connection from the compressor.

Some sweep gas may be included in the above gas stream, but the present
design philosophy indicates that no extra sweep gas should be required ;
however, if some sweep gas is required for efficient degasser operation, this
gas could be obtained by a recycle of previously removed gas. This recycle
sweep stream would most probably be taken from the storage bottles after
sufficient cooling.
900 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

The gas in the storage bottles may be vented to the atmosphere after
90 days of storage, since then only the Kr8% activity is still present in ap-
preciable amounts, and this can be released provided there is sufficient di-
lution. However, the most probable course of action will be to process the
stored off-gas through a gas separation system, where the valuable Kr85
will be recovered.

24—-4. SINGLE-FLUiD REACTOR DESIGN

24-4.1 General description. The single-fluid LMFR concept has been
investigated to determine the characteristics and economic attractiveness
of this design. In general, the core consists of a large array of solid modera-
tor blocks stacked to provide the desirable geometry of a cylinder. Vertical
cylindrical channels are drilled through the moderator to allow circulation
of the liquid metal slurry containing both the fuel and fertile material.
The fission heat generated in the fuel-coolant stream is transported by
forced convection to heat exchangers external to the reactor vessel. The
unique feature of this concept is that only one coolant, the slurry, is used
for removing heat from all parts of the reactor. The desired slurry-to-
moderator ratio is achieved by selecting the appropriate combination of
channel size and spacing.

24-4.2 General specifications. The general specifications for the system
affecting reactor design are tabulated below:
Power 825 Mw (thermal)

315 Mw (electrical)
Slurry temperature:

T; 750°F

T out 1050°F
Maximum slurry velocity 10 fps
Fuel U235 or U233
Fertile material Thorium
Moderator material Graphite or BeO
Slurry carrier Bismuth or lead
Slurry-to-moderator ratio Variable
Fertile material content in slurry Variable
Core geometry Cylinder

Core size Variable
24-4] SINGLE-FLUID REACTOR DESIGN 901

24-4.3 Parametric study. A parametric study was performed to deter-
mine the optimum nuclear parameters for a single-fluid concept. The
variable parameters investigated and their range of values are:

Slurry-to-graphite ratio, V,/V.= 0.05 to 1.0,
Fertile material content, g/kg of Bi= 0 to 80,
Equivalent bare reactor diameter, D, ft = 10 to 20.

The choice of fuel for the first full-scalé LMFR will depend upon the
availability of U233, which is much more attractive than U?3> because of
better neutron economy, and a sufficient quantity for fueling an LMFR
may be available in 10 to 15 yr. In the early stages of this study, how-
ever, U235 was arbitrarily chosen as the fuel for the parametric study.
The selection of the reference design should be valid for either fuel.

In each case the critical concentration and conversion ratio were de-
termined by multigroup diffusion theory, using 37 neutron energy groups.
To handle the large number of calculations, a digital computer was used
once the range of values for the parameters was established by a series of
criticality calculations by hand.

The use of BeO as a moderator has the advantage of reducing the core
size because of improved slowing-down power compared with graphite.
Critical size, fuel concentration, and breeding ratio were determined for
one case, using BeO as moderator.

Since the cost of bismuth as a primary coolant is between $3,000,000
and $4,000,000, the inventory charges are a significant fraction of the
total fuel costs. One case was calculated using lead as a coolant in order to
compare the increase in inventory charges due to the use of bismuth with
the loss in conversion ratio due to the absorptions in the lead.

Basis of nuclear calculations. To obtain comparative results, the follow-
ing specifications were assumed for all cases:

Average temperature 862°F

Graphite density 1.80 g/cc
Bismuth density 9.83 g/cc
Geometry Cylinder (H = D)

For consistency and ease of comparison, all calculations used equivalent
bare reactor dimensions, except the calculation of reflector savings as a
function of reflector thickness.
TABLE 24-6
SUMMARY OF SINGLE-FLUID NUCLEAR CALCULATIONS

 

 

 

 

Thorium, Bare equiv. Initial

Case g/kg Bi, V./V, core size, conversion Nzs/ NfiBi’

Woe D=H, ft ratio x10
11144 0 0.5 14 0.00 153
11154 0 0.7 14 0.00 134
11164 0 1.0 14 0.00 120
11234 15 0.3 14 0.53 458
11232 15 0.3 10 0.43 621
11244 15 0.5 14 0.608 451
11254 15 0.7 14 0.65 481
11324 30 0.2 14 0.625 774
11325 30 0.2 17 0.666 706
11326 30 0.2 20 0.695 666
11334 30 0.3 14 0.692 772
11335 30 0.3 17 0.733 708
11336 30 0.3 20 0.760 671
11342 30 0.5 10 0.63 1181
11344 30 0.5 14 0.746 870
11345 30 0.5 17 0.788 794
11346 30 0.5 20 0.814 751
11424 50 0.2 14 0.735 1199
11425 50 0.2 17 0.780 1099
11426 50 0.2 20 0.801 1054
11434 50 0.3 14 0.788 1304
11435 50 0.3 17 0.817 1203
11436 50 0.3 20 0.843 1144
11444 50 0.5 14 0.787 1757
11445 50 0.5 17 0.827 1598
11446 50 0.5 20 0.854 1504
11514 80 0.05 14 0.565 2099
11524 80 0.2 14 0.804 1990
11525 80 0.2 17 0.840 1853
11526 80 0.2 20 0.865 1769
11534 80 0.3 14 0.816 2452
11535 80 0.3 17 0.849 2274
11536 80 0.3 20 0.875 2160
11544 80 0.5 14 0.747 4471
11545 80 0.5 17 0.788 3959
11546 80 0.5 20 0.851 3684
11435* 50 0.3 17 0.720 1531
11431t 50 0.3 8 0.680 1223
11433t 50 0.3 12 0.765 1032
113441% 30 0.5 14 0.880 678

 

 

 

 

 

 

*Lead coolant TBeO moderator 10233 fyel

 
24-4] SINGLE-FLUID REACTOR DESIGN 903

The resonance integral of the fertile material is a function of the scatter-
ing per atom, size of fuel channel, and lattice spacing. The channel size
and lattice spacing, however, are not specified; therefore, the lattice
resonance parameters are not known. A maximum value of the effective
resonance integral is the homogeneous value based on the scattering in
the core mixture per atom of fertile material. A minimum value of the
resonance integral is the homogeneous value based on scattering in the
slurry per fertile atom. For the cases using thorium, a value of Ro2 (ef-
fective resonance integral) was chosen between the maximum and mini-
mum values, and the calculated uncertainties are 4 209, in the N25/Np;
ratio and &+ 3.39, in the conversion ratio.

Results of nuclear calculations. The results of the parametric study are
summarized in Table 24-6 for all cases. The critical concentrations and
conversion ratios for the cases using thorium as the fertile material are
graphed in Figs. 24-21 through 24-25.

The notation used on all graphs have the following definitions:

N25/Ng; = atom ratio of U235 to bismuth.

Wos = thorium concentration in grams of Th232/gBi.
V./V. = volume ratio of slurry to graphite in core.
D = the equivalent bare core diameter in feet.

In all cases, the fuel concentration increases with an increase in fertile
material, Woe (Fig. 24-24). An increase in V,/V, increases the thorium
content, reduces the slowing-down power, increases the average energy of
the neutron spectrum in the core, and increases the thorium absorptions.
As a result of these effects, the critical fuel concentration in the fluid fuel,
N35/Ng; ratio, increases as Vs/V, increases (Fig. 24-25).

0.9 -

 

0.8

 

Conversion Ratio
o
N

 

. ’ ~ Vs/Ve=0.2
)/ / —.—Vs/Vc=0.3
/ ——=-Vs/Vc=0.5

0.6 - / ............ VS/Vc=O.7 ]

. | | | | |
20 30 40 50 60 70 80

W g9, Grams Thorium/kg Bi

 

 

 

Fig. 24-21. Conversion ratio vs. thorium concentration for a single-fluid LMFR.
904 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

0.9

 

Conversion Ratio
o o
u o

o
o

 

 

 

 

0.5
0.2 0.3 0.4 0.5 0.6 0.7

Vs/Vc

Fia. 24-22. Conversion ratio vs. slurry-to-graphite volume ratio for a single-
fluid LMFR.

 

0.90 l I

0.85

o
o
o

Conversion Ratio
o
9
(3

o
N
o

 

 

 

 

 

0.65 }— - Vs/Vc=0.3 —
// -———Vs/Vc=0.5
_— —-—Vs/Vc=0.2
060 | | | | |
14 15 16 17 18 19 20

Bare Diameter, Feet

Fie. 24-23. Conversion ratio vs. diameter for a bare single-fluid LMFR.
 

 

 

 

 

 

24-4] SINGLE-FLUID REACTOL DESIGN 905

U Vs/Vc=0.2 ,
o — D=1
'; 4000 ———— Vs/Vc=0.3 D= 22,

5 3600 — Vs/Ve=0.5

Z © Vs/Vec=07

W 32001—

z\

2 2800—

5 D=14
T 2400(— D=17'
S D =20’
< 2000 D=14
'.';: 1D=17’
E 1600|— D =20’
8

3 800 }—

400
10 20 30 40 50 60 70 80

Wgo: Grams Thorium/kg Bismuth

F1a. 24-24. Critical fuel concentration vs. thorium concentration for single-fluid

LMFR.

5000

1000

Fuel to Bismuth Atom Ratio,N,./N,. x 106
25" ""Bi

100

 

 
    
 

R " b=14" |
— Wo2= 80 D=17" -
| Wo2 = 50 D=20" i
| ———-—Wp2=30
—cmma—ee Wpp =15 R
_.D=14" |
4’:’—D= ]7'

1

PPN

 

| | | | | |

D=14'
——— e = ______-_——-—‘--——
— -
\-.\Q.N--D = ] 4’

| gt

 

 

0.4
Vs/Vc

0.1

0.7

Fic. 24-25. Critical fuel concentration vs. slurry-to-graphite volume ratio for a

single-fluid LMFR.
906 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

The conversion ratio is highly dependent upon the No2/N 25 ratio, the
average energy of the neutron spectrum, and the reactor size. Figure 24-21
shows that for larger values of V,/V ., the conversion ratio passes through a
maximum as thorium concentration increases; however, for smaller values
of V/V., the conversion ratio increases continuously as V,/V,. increases
over the range of interest; i.e., the maximum value of the conversion
ratio shifts to higher values of W2 as the V,/V, ratio decreases. Likewise,
the curves of conversion ratio versus V,/V, go through a maximum, with
the maximum value occurring at increasingly higher values of V,/V,. as
W2 increases (Fig. 24-22).

An increase in core diameter simply reduces the neutron leakage. As a
result, the conversion ratio increases as the diameter increases. An in-
crease in D from 14 to 20 ft increases the CR approximately 0.09 (Fig.
24-23).

Case 11435 was recalculated using lead instead of bismuth as the coolant
fluid. The conversion ratio decreased by 0.10, and the critical N25/Np;
ratio increased from 1203 X 107% to 1531 X 1076,

Beryllium oxide, BeO, was used as moderator in another variation of
Case 11435. This calculation, case 11433, for a diameter of 12 ft, requires
an N25/Np; ratio of 1032 X 107¢ and yields the slightly lower conversion
ratio of 0.77.

The worth of a pure graphite reflector was calculated for Case 11435.
The reflector savings as a function of reflector thickness are shown in
Fig. 24-26. The reflector savings are approximately equal to the reflector
thicknesses for reflectors less than 2 ft thick.

The values of conversion ratio and Ng5/Np; ratio calculated in this
parametric study are for hot, clean reactor conditions, and they are used
for comparative purposes only. The effects of fission-product poisons,
control rods, and Pa?33 losses have not been included.

24-4.4 Economic optimization. The selection of parameters for a refer-
ence design must be based upon economics. An economic optimization
was accomplished by computing relative energy costs based on those
variable costs which depend upon the parameters selected. The costs
which are dependent upon the nuclear parameters are (1) bismuth in-
ventory, (2) fuel inventory, (3) fuel burnup, (4) thorium inventory,
(5) thorlum burnup, (6) reactor core and vessel, and (7) chemical process-
ing costs.

Reactor cost. Since the range of reactor sizes varies from 10 to 20 ft,
reactor cost 1s an important variable. Reactor vessel, graphite, and erec-
tion costs have been estimated for several sizes; to these is added $167,000
for three control rods and miscellaneous hardware. Contingency and en-
gineering of 449, were also assumed. A breakdown of these costs is listed
in Table 24-7.
I

SINGLE-FLUID REACTOR DESIGN 907

80 I

 

! I l l | I
60 |- —
40 |— —

20 — ]

Reflector Savings, SR’ In cm

 

 

l l | | | | | |
0 10 20 30 40 50 60 70 80 90
Reflector Thickness, cm

Fie. 24-26. Reflector savings vs. reflector thickness for a single-fluid LMFR.
These data are obtained from case 11435, where V,V,= 0.3, W2 = 50 g/kg, and
Dp=17.

 

 

 

 

 

 

 

 

 

 

 

 

TABLE 24-7
EstiMATED SINGLE-FLUID REAcTOR COST

Size, Reactor | Graphite | Misc. | Erection| Total Total $/yr
ft : cost

10 160,000 350,000 | 167,000 | 24,000 701,000 | 1,009,440 | 151,400
14 | 380,000 | 970,000 | 167,000 | 30,000 | 1,547,000 | 2,227,680 | 334,152
17 570,000 | 1,700,000 | 167,000 | 35,000 2,472,000 | 3,559,680 | 533,952
20 900,000 | 2,800,000 { 167,000 | 40,000 3,907,000 | 5,627,080 | 843,912

 

 

Bismuth inventory charges. The bismuth inventory is determined by the
primary system volume external to the reactor vessel, the volume of bis-
muth in the core, the volume of bismuth external to the core but inside the
reactor vessel, and the holdup external to the reactor system. The primary
system external to the reactor vessel is made up of three heat-exchanger
loops containing a total volume of 1640 ft3. The volume of bismuth in
the core 1s

Vs/Ve

VeI v, 7,

Vgi= » where V, = core volume
The volume of bismuth external to the core and inside the reactor vessel
is tabulated in Table 24-8.

No additional holdup is included to account for temperature expansion
during startup, fuel feed system, and other sources of bismuth inventory.
The assumption used throughout this study that the volume of bismuth is
equal to the volume of slurry accounts for an additional 3 to 10% excess

bismuth due to the ThO2 content of the slurry.
908 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

TABLE 24-8

BismMmutH INVENTORY IN REACTOR VESSEL
EXTERNAL TO CORE

 

 

 

Core diameter, ft Bi inventory, ft3
10 550
14 600
17 650
20 700

 

 

 

 

The density of bismuth is taken as 9.83 g/cc, and the price is assumed
to be $2.25/lb. Bismuth is a nondepreciating capital investment with a
12% annual amortization rate. The annual bismuth inventory charges
may be represented by the equation

Vs VC
Cu(8/y1) = 0.12225) |V, /0247, o

where
Vp = total primary system volume except core, ft3,

pBi = density of bismuth, 1b/ft3.

Fuel wnventory charges. The annual lease charges on the U235 are as-
sumed to be 4%. Treating Pa233 as fuel, the annual fuel inventory charges
can be expressed as

C2($/yr) = 0.04VosM o5+ VasM a3 + VisMys,

where
Vas = Vi3 = value of U233 as fuel,

Va5 = value of U232 as fuel, $17,760/kg,

M ; = average mass of element j in entire reactor system
during life of plant.

To simplify the work in the absence of information concerning average
values of fuel mass, the total mass of fuel was considered to be the hot, clean
critical loading at startup. The value of M 5 is taken as the initial value
with M 23 and M;3 taken as zero.
24-4] SINGLE-FLUID REACTOR DESIGN 909

Fuel burnup costs. Using U235 ag fuel, the yearly burnup costs are

C3($/yr) = 17.76(292) PB(1 — CR),

where
P = power, 825 Mw,

B = grams of fuel burned per MwD, 1.25

CR = average conversion ratio.

The initial value of the conversion ratio is used, since only relative costs
are needed.

Thorvum burnup costs. Thorium is periodically replenished in the reactor
to maintain the desired concentration in the slurry. The thorium burnup
costs may be expressed as

C4($/yr) = Vo2 PBCR(292),

where Vo2 = value of thorium, $42/kg. These costs are very small, ap-
proximately $10,000/yr, and are neglected.

Chemucal processing costs. The chemical processing is assumed to use
solvent extraction aqueous chemistry in a central processing plant. The
irradiated fuel is removed from the reactor on a batch processing cycle.
The processing costs are represented by

Cp = 292 [95.875 MT(E + 4795 M;;S ) + 250&9 00 + 596] ’

 

where

M2 = total thorium inventory kg,
23(T) = M 23+ M3 at time T after loading of fuel charge, kg,

T = chemical processing cycle time, days.

Results of economic optimization. Since chemical processing costs are
very sensitive to the chemical processing cycle time and the optimum cycle
time may vary with reactor design, the relative energy cost of each reactor
design was determined neglecting the chemical processing costs. The results
of this study are tabulated in Table 24-9 and are shown graphically in
Figs. 24-27 and 24-28.

The pure burner, W2 = 0, shows costs more than twice as high as several
of the more attractive concepts (Fig. 24-28). In general, the minimum
910 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

 

 

   
    

 
 

 
 

 

 

2.8 1
T T T T T | | 'Vs/Vc==0.5
26 ——- D=20'
\ T D=17'
2.4 | D=14 Vs/Ve=0.5
\ Vs/Vc=0.5
a 2.2 —\ //// S/ Cc 0
8 \\ s s
.‘é’ \ Vs/Vc=0.3
18 \ 4
- \\ Vs/Vec=0.2
& < Vs/Vec=0.3
- N\ < s
' NN — //Vs/‘/c=03
~ RN - VS/VC:O 2
141 S~ T \:_,\/\:‘é’—‘—:*—./ -:-._.—__;3;_"‘ / Vs/Vc=0.2
1.2 |
10 20 30 40 50 60 70 80 90

W2, Grams Thorium/KG Bismuth

Fic. 24-27. Relative cost vs. thorium concentration for a single-fluid LMFR.

 

 

 

 

3.0 | l I T i l I I 1 I' | ] T I T
- =0
2.8 e oo Wo2=9
;7027
26 - ————D=20 / —
_‘_—D=]7, //

2.4 - ____D=~|4’ /// W02 =80 ]
%90 | 2 D=10 // //W02=80 i
O // ,

.g 20 — /// //./ —
c /s
3 s |- \ / // _
\ W02 = » /yx02igg
1.6 02 —_—
N 2:? —~ /Woz 50

1.4 |— \'\_ = " W02 30

1.2 |— . _"'Wog 30 |

1.0 i I i I s I I J { l 1 l 1 l i

0 .10 .20 .30 40 .50 .60 .70 .80
Vs/Vc

Fic. 24-28. Relative cost vs. slurry-to-graphite volume ratio for a single-fluid
LMFR.

costs are achieved with thorium loadings corresponding to Woe = 20 to
50 g/kg. The most attractive designs do not have the highest values of
conversion ratio.

In many cases the additional fuel inventory charges and reactor vessel
costs corresponding to higher conversion ratios more than offset the reduc-
tion in fuel burnup costs. The economically optimum reactor is neither a
burner nor a converter with maximum conversion ratio, but somewhat
between these extremes.

Using a cost of 18¢/1b for lead as a coolant, comparisons of lead versus
Bi as a coolant were made for Case 11435. The annual fixed charges on
lead were only $44,000 compared with $478,000 for bismuth in this case;
24-4] SINGLE-FLUID REACTOR DESIGN 911

however, the increased fuel burnup and inventory charges associated with
the lead coolant resulted in a net increase of $288,000/yr or 0.14 mills/kwh
in the fuel cost. BeO is not feasible as a moderator material for this concept
because of its high cost. Fixed charges on the BeO alone add almost a
mill/kwh to the fuel cost.

The six most attractive cases were selected and the chemical processing
costs computed for several processing cycle times. The total costs tabulated
in Table 24-10 are based on a 3000-day cycle time, and other costs are from
Table 24-9. Since the aqueous processing costs are dependent upon the
total thorium inventory to be processed, chemical processing costs penalize
the designs with heavy thorium loadings. In Tables 24-9 and 24-10, the
total costs are reduced to mills/kwh by using an electrical power output
of 315 Mw with an 809 plant factor.

24-4.5 Selection of a reference design. Using the data presented in
Table 24-10, a design was selected for further study. It is important to
realize that when chemical processing costs are included in the comparison
of energy costs, there is little difference in the cheapest four or five cases.
The relative attractiveness of these cases depends very heavily on the
economic ground rules. Kven a change in chemical processing cycle may
change the relative order of the cases. With the wide range of freedom for
choice of nuclear parameters in this concept, the economic optimum can
be chosen to correspond to any set of basic assumptions on economics.
For example, an increase in fuel price would emphasize higher conversion
ratios. The design selected for further study was Case 11344.

Time study. The nuclear performance of the reference design, Case 11344,
was determined using a thorium lifetime program written for the digital
computer. These calculations provided information concerning the varia-
tions of fission-product poisons, breeding ratio, and critical fuel mass as
functions of reactor operating time. This information then made possible
the choice of an optimum fuel processing cycle and the determination of
over-all fuel cost for the operating reactor.

Basis of time study. The reference design calculations used U233 as fuel
for both the initial charge and feed material. Since the contemplated con-
struction date for an LMFR is 10 yr in the future, the assumption that
U233 fuel will be available seems reasonable, and data based on U233 allows
comparison with previous work [2].

The reference design on which the time studies were based has a graphite
side reflector 1.5 ft thick, an active core diameter of 11 ft, and a core
height of 14 ft. The average core temperature is 900°F. The nuclear con-
stants used in the two-group criticality and isotope buildup calculations
were determined by using a 40-group spectral code.
[cHAP. 24

LIQUID METAL FUEL REACTOR DESIGN STUDY

912

 

 

 

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[cHAP. 24

LIQUID METAL FUEL REACTOR DESIGN STUDY

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24-4] SINGLE-FLUID REACTOR DESIGN 915

The neutron poisons due to fission products and higher uranium iso-
topes were calculated using the data by W. L. Robba et al. [4]. The xenon
poisoning (absorptions in Xel3% to absorptions in fuel) was held at 0.01
throughout life, and Sm!4° was allowed to reach steady state. The other
fission-product poisoning corresponds to poison data labeled ‘“‘less Xe and
Sm with high cross sections except low Zr?3.”” Due to lack of information,
no resonance absorption by the fission products was considered. The
neutron flux averaged over the entire primary system volume was used in
all isotope and poison buildup computation, since this is a circulating fuel
reactor.

Fuel was added at frequent time intervals to maintain ks = 1.01
(assuming 19, rod holddown). Thorium was added to the core with the
fuel to maintain a constant thorium loading.

Results of time study. The study was carried to 2000 days of full-power
operation. The mass of U233 fuel and the buildup of Pa?33 are shown in
Fig. 24-29, and the buildup of fission product poisons (other than Xel35
and Sm'49) along with breeding ratio are graphed in Fig. 24-30. The fission-
product poisons vary in an almost linear manner for burnups corresponding
to 2000 to 6000 days. Other calculations have indicated that extrapolations
(represented by dashed lines on Figs. 24-29 and 24-30) to 5840 days, the
expected life of the plant, are reasonable.

The quantities necessary to evaluate the chemical processing costs for
various processing cycles are average values of fuel mass and breeding ratio
(M 23, Mys, and BR). The average value of M3 is approximately the
steady-state value; M 23 and BR are shown in Figs. 24-29 and 24-30.

Selection of chemical processing cycle. The fuel costs which are dependent
upon the chemical processing cycle are fuel inventory, fuel burnup, and
processing charges. These charges were computed using formulas similar
to those described in Article 24-4.4 but using data appropriate to U233 fuel.
Equations giving costs in dollars per full power day are

Fuel inventory: Cs ($/day) = 2.143 (M 'zii— M;i3)
Fuel burnup: C3 ($/day) = 15,250 (1 — BR)

Chemical processing:

Cp (8/day) = 95.875 2202 1 4795 MtT) 4 200000

The results of these calculations are tabulated in Table 24-11 and

graphed in Fig. 24-31. This analysis indicates an economic optimum

processing cycle of approximately 4000 full-power days. However, only a

small penalty of slightly more than $200/day (less than 0.03 mills/kwh) is

incurred by operating the reactor for its complete life (5840 full-power days)
before sending the fuel to a chemical separations plant.

—+ 596.

 
 

 

 

 

 

 

 

 

 

 

916 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24
800
700
o
-
w" 600
™
S
500 _J Fi1a. 24-29. Mass of U233
vs. reactor operating time
50 D M3 ‘f for a single-fluid reactor
| | | | | operating at 825 Mw.
0 1 2 3 4 5
Reactor Operating Time, In Full Power Days x 10-3
0.9 l | | |
08 |- - -
%07 |- Bt
S
5 T
303 |- .
Z
g - =’ <
T -
2 0.2 |~ 'o’ —
”"'

Fia. 24-30. Neutron ’,” Fission Product Poison
losses vs. reactor operating 011" B
time for a single-fluid re-
actor operating at 825 Mw. 1 l l l

0 1 2 3 4 5

Reactor Operating Time, In Full Power Days x 10-3

 

 

 

 

 

11000 , , I , ! 1.4
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O
o
~
o
S 9000 |-
o T
S 8000 |- 2  Fia. 24-31. Varia-
2 Z ble fuel cost for
- 825-Mw single fluid
* 7000 |- LMFR vs. chemi-
—10-2 cal processing cycle
(aqueous batch proc-
6000 | L ' | ' ess).
0 1000 2000 3000 4000 5000 6000

Cycle Time, T, Full Power Days
SINGLE-FLUID REACTOR DESIGN 917

24-4]

 

 

 

 

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[cHAP. 24

LIQUID METAL FUEL REACTOR DESIGN STUDY

918

Shielding

-
(.

Fuel Out

_—_OOOOOOOdOdOsdhdSs-
?///////////////////////////////////////////////////////
///fi///// MMM TR
AN /////////////////////////////////////////////////////
/////////////////////////////////////////////////////
////////////,/////,////////// NI IO
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Graphite

’

Fuel Passage

Fuel In

Fig. 24-32. Single-region, externally cooled liquid metal fuel reactor.

Specifications of reference design. The single-region reactor design is 1l-
lustrated in Fig. 24-32. The core is constructed of large blocks of high

density reimpregnated graphite, with 1.5- to 2.0-in.-diameter axial holes

for the passage of fuel slurry.

The graphite is supported by a number of

made for three or four liquid metal control rods, if experience indicates

compensated molybdenum rods and a bottom support plate. Provision is
they are necessary.

The reactor vessel is constructed of 219, Cr-19;, Mo steel, 2% inches

thick, designed for a temperature of 1150°F and maximum pressure of

120 psi. Three 28-in.-diameter pipes carry the fluid into the reactor at the
bottom and leave at the top. The entire reactor vessel is doubly contained

A drain line to the fuel

dump tanks is also provided. The free space above the reactor core is used

by a relatively thin-walled containment vessel.

The reference core

as the degasser to remove volatile fission products.

design has the following specifications:
24-4] SINGLE-FLUID REACTOR DESIGN 919

Power:

Thermal power
Net electrical power
Station efficiency

Materials:

Fuel

Fertile material

Moderator

Reflector

Coolant

Coolant-to-moderator ratio, V,/V,
Thorium concentration, W»

Geometry:

Core radius
Core height
Reflector thickness
Number of primary coolant loops
Fuel-slurry volume:
Coolant loops 1640 ft3

 

Reactor core 443
Reactor vessel 600
Total

Chemical processing cycle

Nuclear data:

Startup
Mass U233 546 kg
Mass Pa233 0
Mass 233 046
Initial average core thermal flux
Breeding ratio 0.87
Poison fraction 0
Mass of bismuth 1,646,000 1b

Mass of thorium 22,400 kg

825 Mw
315,000 kw
38.29,

U233
Thorium
High-density graphite
High-density graphite
UO2-ThO2-Bi Slurry
0.5

30 g/kg Bi

5.5 ft
14.0 {t
1.5 ft

2683 ft3

4000 days

4000 days Average
725 kg 675 kg

30 30
755 705
3 X 10
0.725 0.75
0.216
920 LIQUID METAL FUEL REACTOR DESIGN STUDY [cHAP. 24

24-5. EcoNowMmics

Economic considerations were essential to the optimization studies re-
quired to establish the reference designs presented in Sections 24-2 and
24-4. An important objective of this study is the economic comparison
of energy costs for the single-fluid and the two-fluid externally cooled
LMFR. A brief summary of energy costs for the optimum design in each
concept is presented in Table 24-12.

TABLE 2412

Enercy Cost (809, PrANT FACTOR)

 

Mills/kwh

 

Single-fluid Two-fluid

 

 

LMFR LMFR
Fixed charges on total capital investment 4.09 4.31
Nuclear fuel and inventory costs 1.24 1.41
Maintenance 1.18 1.05
Operation 0.38 0.38
Interest on working capital 0.04 0.04
Total energy cost, mills/kwh 6.9 7.2

 

 

 

 

 

24-5.1 Fixed charges on capital investment. Direct construction cosis.
The estimated costs of equipment, installation of equipment, and con-
struction are based on the plant layouts for the two reference designs eval-
uated in this study. Construction and erection costs of all items, as well as
direct materials costs for those components manufactured by the Babcock
& Wilcox Company, were developed by B&W estimators. Delivered costs
of equipment supplied by manufacturers other than B&W were taken from
vendors’ quotations.

A summary of direct construction costs for each reference design 1is
tabulated in Tables 24-13 and 24-14.

Total capital investment. The total capital investments are summarized
according to account numbers in Tables 24-15 and 24-16.

24-5.2 Maintenance and operation. In computing energy costs, the
fixed charges on maintenance equipment and spare parts are included in
the maintenance costs, while fixed charges on buildings used for main-
tenance are included in fixed charges on capital investment.
24-5] ECONOMICS 921

24-5.3 Fuel costs. The fuel costs as presented in this report include
(1) bismuth inventory, (2) fuel inventory, (3) fuel burnup, (4) thorium
inventory, (5) thorium burnup, and (6) chemical processing. Sodium in-
ventory 1s not included, since it is used as coolant fluid for the inter-
mediate system and does not contain fuel. Fuel costs are summarized in
Table 24-17.

24-5.4 Summary of energy costs. The energy costs in mills /kwh, based
upon an electric output of 315,000 kw and a plant factor of 80%, are
tabulated in Table 24-18 for various categories.
[cHAP. 24

LIQUID METAL FUEL REACTOR DESIGN STUDY

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24-5] ECONOMICS 929

TABLE 24-18

Unit ENERGY COSTS

 

 

 

 

Cost, mills/kwh
Item
Single-fluid Two-fluid
Land and land rights 0.03 0.03
Structures and improvements (less chemical
processing facilities) 0.69 0.72
Equipment (less maintenance equipment and
spares):
Reactor vessel and internals 0.19 0.14
Primary and blanket system 0.21 0.27
Intermediate system 0.36 0.44
Feedwater heating system 0.26 0.27
Instrumentation and controls 0.29 0.36
Miscellaneous equipment and Na inventory 0.11 0.10
Auxiliary systems 0.22 0.27
Station equipment 0.14 0.14
Accessory electric equipment 0.28 0.28
Turbine generator equipment 1.29 1.29
Miscellaneous power plant equipment 0.02 0.02
Fuel costs (includes chemical processing fa-
cilities) 1.24 1.41
Plant operation 0.38 0.38
Maintenance (includes maintenance equip-
ment and spares) 1.18 1.05
Interest on working capital 0.04 0.04
Total 6.93 7.21

 

 

 

 

 

REFERENCES

1. BaBcock & WiLcox Co., 1958. USAEC Report BAW-1046.

2. BaBcock & Wincox Co., Liquid Metal Fuel Reactor; Technical Feasibility
Report, USAEC Report BAW-2(Del.), June 30, 1955.

3. BaBcock & Wircox Co., 1958. USAEC Report BAW-1048.

4. W. L. RoBBaA et al., Fission-product Buildup in Long-burning Thermal
Reactors, Nucleonics 13(12), 30-33 (1955).

5. BaBcock & Wincox Co., A Review and Evaluation of Maintenance Concepts
for Liquid Metal Fuel Reactors, USAEC Report BAW-1047, March 1958.
CHAPTER 25
ADDITIONAL LIQUID METAL REACTORS

In this chapter three other types of Liquid Metal Fuel Reactors will be
discussed. The first of these is the Liquid Metal Fuel Gas-Cooled Reactor.
In principle this reactor is similar to the LMFR previously discussed, but
it has many features that are different; for example, it has a noncirculating
fuel, and the heat is removed by cooling with helium under pressure.
Advantages and disadvantages of this design over the circulating fuel
LMFR will be discussed in the following pages.

The second reactor discussed in this chapter is the LAMPRE. This is a
molten plutonium fueled reactor which is under development at the Los
Alamos Scientific Laboratory. Although only in its beginning stages of de-
velopment, it is conceived as a high temperature (650°C) fast breeder re-
actor utilizing plutonium as the fuel.

The third type of reactor is based on a liquid metal-UQ; slurry fuel.

25-1. Liquip METAL FueL Gas-CooLeEp REACTOR*

25~1.1 Introduction and objectives of concept. The Liquid Metal Fuel
Gas-Cooled Reactor (LMF-GCR) design is unique in that it combines
inert gas cooling with the advantageous liquid fuel approach. The LMF-
GCR concept has a high degree of design flexibility. It is a high-tempera-
ture, high-efficiency system that may be designed as a thermal converter,
uranium thermal breeder, or plutonium fast breeder; that may produce
heat, electric energy, or propulsive power; and that may power either a
steam or a gas turbine.

The fundamental principle of the LMF-GCR is the utilization of an
internally cooled fixed moderator-heat exchanger element with fluid fuel
center. The fuel is circulated slowly through the core to assure proper
mixing and to facilitate fuel addition. The core is cooled by gas that is
pumped through it in passages that are separated by a suitable high-
temperature material from the fuel channels. The many well-known
advantages of fluid fuels are thereby gained without the penalties of
circulating great quantities of corrosive, highly radioactive fuel-coolant
solution and of tying up large amounts of expensive fuel outside the core.

*American Nuclear Power Associates: Raytheon Manufacturing Co., Waltham,
Mass.; Burns and Roe, Inc., New York City; The Griscom-Russell Co., Massillon,
Ohio; Clark Bros. Co., Olean, New York; Orange and Rockland Utilities, Inc.,
Nyack, New York. Reference design by Raytheon Manufacturing Co. This sec-
tion is based largely on contributions from W. A. Robba, Raytheon Manufac-
turing Co.

930
25-1] LIQUID METAL FUEL GAS-COOLED REACTOR 931

25-1.2 Reference design characteristics of an LMF-GCR. Materzals.
Internal gas cooling avoids the corrosion and material problems encoun-
tered in reactor concepts that require the circulation of liquid fuels or
coolants as a heat-transport medium. Helium has been selected as the
gas coolant because it is inert and has better heat-transfer properties than
other inert gases. Graphite has been chosen for the moderator and core
element structural material in a thermal reactor, because of its excellent
moderating and high-temperature properties. Its resistance to corrosion
by bismuth has been fairly well established, and the operating temperature
is high enough so that energy storage in the graphite should not be a
problem.

Reference design. A reference design of an LMF-GCR nuclear power
station has been produced. A summary of the design parameters is given
in Table 25-1. It is a graphite-moderated thermal reactor employing highly
enriched uranium-bismuth fuel and helium coolant. The coolant leaves the
core at 1300°F and is circulated through a superheater and steam generator,
where it produces steam at 850 psig, 900°F. Since it is inherently self-
regulating, has little excess reactivity, and is cooled by inert helium, it is
extremely safe.

In order that the capital cost of the first plant be low, the reference
design is for a small plant producing approximately 16,000 kw net electrical
output. However, it is large enough to demonstrate the practicability of
an LMF-GCR and provide operational experience applicable to com-
mercial-size plants. By assuming the feasibility of constructing a 13-t
diameter pressure vessel for a design pressure of 1000 psi, it appears possible
to design a gas-cooled reactor plant having an electrical capacity of 240
Mw.

A U235fyeled thermal reactor was chosen for the design because 1t will
demonstrate the practicability of the LMF-GCR concept in a relatively
simple reactor. A breeder is more complicated because it requires two
similar systems for fuel and blanket solutions.

The reactor building and the general arrangement of components as
conceived in the reference design are shown in Fig. 25-1. The reactor,
primary coolant system, fuel system, and steam generator are enclosed in
a gastight steel containment shell.

The reactor core, reflector, internal fuel and gas piping, and pressure
vessel are shown in Fig. 25-2. The core, consisting of an array of graphite
elements, has an active length of 56 in. and a cross section approximating
a circle of 56-in. diameter. Fig 25-3 is a picture of a sample section of the
core element. The larger rectangular holes are vertical fuel channels that
would be 56 in. long in the reactor. The small crosswise slots are for helium
coolant flow. This graphite element, which separates the two fluids, is
similar to a heat exchanger that conducts heat from the fuel to the gas
932 ADDITIONAL LIQUID METAL REACTORS [cHAP. 25

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F1a. 25-1. Artist’s concept of LMF-GCR nuclear power station.

channel surface, where it is removed by convection into the coolant stream.
The principal problem associated with the LMF-GCR is the development
of an impervious graphite core material that will prevent significant leakage
of bismuth or fission-product gases into the coolant stream, or of helium
into the fuel.

The machining operations required to produce the core section of the
element have been demonstrated to be feasible. The reflector is made up of
various machined graphite shapes. The fuel piping completes the core and
the reactor assembly.

By volume, the core region is approximately 659, graphite, 259, fuel,
and 109, void space for coolant. The fuel solution contains fully enriched
uranium dissolved in bismuth. With these proportions of fuel and moder-
ator, the minimum critical dimensions as calculated for a cylindrical re-
actor are height and diameter of approximately 42 in. For this application,
a larger core size is required in order to have sufficient heat-transfer area.
Since the graphite core elements are a permanent part of the reactor and
are not changed in routine refueling procedure, it is not required that they
be interchangeable. A considerable amount of design flexibility is thereby
achieved, and variations of the fuel channel, moderator, and gas channel
geometry provide control over the nuclear and heat processes.

For the reactor described above, it is calculated that 900 atomic parts
of U23° per million parts of bismuth are necessary for criticality, if there is
no poisoning of the reactor. However, if the effect of xenon and samarium
25-1] LIQUID METAL FUEL GAS-COOLED REACTOR 933
1999 QL

1

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Fic. 25-2. Reactor and pressure vessel assembly.

equilibrium poisoning is included, 1010 ppm of U235 will be required for
criticality.

The buildup of fission products and uranium isotopes as a function of
time was calculated to determine the fuel concentration necessary for
criticality after various time periods of operation. Since the solubility of
uranium in bismuth is limited to 6560 ppm at 965°F, the lowest fuel tem-
perature in the LMF-GCR, the reactor fuel must be replaced or processed
after the poisons build up to such a level that this solubility limit is exceeded
by criticality requirements. With the total fuel inventory in the system
equal to 1.2 times the fuel in the core, the fuel lifetime will be 220 megawatt-
years. This corresponds to an operating period of 4.8 years with a plant
utilization factor of 809.

At the end of the fuel lifetime, the fuel solution will contain 3370 ppm
934

ADDITIONAL LIQUID METAL REACTORS

[cHAP. 25

 

Fia. 25-3. Model section of nuclear core element for LMF-GCR liquid metal
fuel gas-cooled reactor.

 

Liquid Bismuth

 

Charcoal
Trap
-*A
OAAA~O
Bismuth

Charge Tank

Spent Fuel
Shipping Tank

 

  

 

  
 
 
  
 
 
 
 

Bismuth
Condenser

Fuel Feeder
& Sampler

 

 

 

  
   

 

 

 

 

 

 

 

 

 

 

   
 

 

 

 

 

 

850 psig
/ 900°F
Bismuth-Uranium Helium 4
Solution
Degasser Super-
| Electric Heater Heater
I Trap
|
|
|
I
|
Diff. Press. - 1300°F
Controller |
| Helivm
| | |_Cover Gas
i ' Storage
! Steam
| Reactor ea
@ O @ ' Generator
|
I
|
|
|
l
! L ——— Feed
Water

Pump

l
|
|
L.______......_.._.'

 

 

 

 

 

Water

Electromagnetic
\_1 Pump Helium
S 1 ! » Coolant
pen Storage
Fuel Tank Coolant Pressure Control 9

F1a. 25-4. Over-all plant flow diagram.
25-1] LIQUID METAL FUEL GAS-COOLED REACTOR 935

of U235 1960 ppm of U236 and 1230 ppm of U238 which make up the 6560
ppm of uranium allowed by the solubility limit.

In producing the 220 Mw-yr of heat, 98.7 kg of U235 will be either fis-
sioned or transmuted into U236, Since 23.8 kg of U235 remain in the reactor
at the end of fuel lifetime, approximately 809, of the total amount of U235
added to the reactor during its operation will have been “‘burned.”

The systems required in the plant are shown by the flowsheet of IFig. 25-4.
The heat is removed from the reactor by helium at 500 psia, which leaves
the reactor at 1300°F and returns at 900°F. This heat is removed from the
helium in a steam generator that produces superheated steam at 850 psig,
900°F. The steam is utilized by a standard turbine generator plant.

A steam-cycle generating plant was incorporated, since it is highly de-
veloped. A closed-cycle gas turbine, the most probable alternative, has
not yet been developed sufficiently for general utility application, but may
be advantageously combined with the LMF-GCR at some later time. In
such a system, the reactor coolant would serve also as the cycle working
fluid, eliminating the intermediate heat exchanger.

Although the reference LMF-GCR is envisioned as a high-enrichment
reactor, 1t is possible, by changing the parameters, to use fuel of only 209,
enrichment. This low-enrichment reactor would have the advantage of
producing a sizeable fraction of its own fuel by creating Pu?3® through
neutron absorption in U238,

Parametric calculations of low-enrichment reactors have been made
using a two-group, two-region spherical geometry computer code developed
for the IBM 650 digital computer. The results show that to have a fuel
lifetime long enough (about 1 yr) to be of practical value, the dimensions of
the reactor core should be equivalent to a sphere at least 6 ft in diameter.

25-1.3 Fuel and fuel system. Fuel system. The fuel system is completely
separate from the heat-removal system. The main fuel loop flow rate is
approximately 2 to 4 gpm, which is sufficient to provide for uranium makeup
and for gas separation in the degasser.

Fuel flows upward through the reactor core and into the degasser. From
there, the flow goeés down into the sump tank and back into the reactor
inlet. The fuel is pumped electromagnetically and flow is measured by an
orifice or an electromagnetic low meter.

The sump tank acts as a receiver for all the fuel in the loop when the
core is to be drained. To keep the sump tank nearly empty during operation,
the pressure differential between the helium cover gas in the sump tank
and the degasser must be kept equal to the bismuth static head. The
fuel is automatically drained into the sump tank when the pump is de-
energized and the two cover gas lines are connected together. Thus there
are no valves in the primary fuel loop which must be operated in order to
drain the reactor.
936

TABLE 25-1

ADDITIONAL LIQUID METAL REACTORS

[cHAP. 25

SUMMARY OF DESIGN PARAMETERS

 

Over-all plant performance
Reactor core thermal power
Helium blower power
Net electric power generated
Plant efficiency

Thermal data on reactor at full power
Helium pressure
Coolant inlet temperature
Coolant exit temperature
Coolant flow rate
Coolant velocity in core
Number of flow passes
Average thermal power density
Peak thermal power density
Peak to average heat flux ratio (average over life)
Design heat output
Maximum graphite temperature
Maximum fuel temperature
AP/P through reactor

Steam plant data
Pressure
Temperature
Flow rate
Number of extractions
Turbine heat rate
Condenser pressure
Turbine speed
Gross turbine output

Pressure vessel
Material
Outside diameter
Thickness
Over-all length
Weight
Type of closure
Insulation

Core
Neutron energy
Fuel, clean

 

Fuel lifetime

 

57,000 kw
5,530 kw

16,470 kw
28.99%,

500 psia
900°F
1300°F
389,000 Ib/hr
~ 560 fps
1
0.714 Mw/ft3
~0.922 Mw/ft3
~1.29
1.94 x 108 Btu/hr
1650°F
1755°F

4.3%

850 psig
900°F
188,300 Ib/hr
4
9,645 Btu/kwh
1.5in. Hg
3600 rpm
22,000 kw

Stainless steel
94 in.
2 1n.
123 in.
30,000 Ib.
Bolted
4 in. of diatoma-
ceous earth

Thermal
900 ppm of U235
93.59, en-
riched U in Bi
220 Mw-yr

 

 
25-1] LIQUID METAL FUEL GAS~COOLED REACTOR 937

TABLE 25-1 (continued)

 

Reprocessing interval (0.8 plant factor) 4 8yr

Fuel burnup 809 of U235
Moderator 1.9 g/cc graphite
Bismuth in core 11,400 Ib

Bismuth volume fraction 259,

Graphite volume fraction 659%

Void (helium) fraction 109,

Average core radius 28 in.

Core height 56 in.

Core volume 79 .8 ft3

Power 57 Mw

Specific power, average over fuel lifetime ~3700 kw/kg
Power density (based on core volume in liters) 25.2 kw/liter
Average thermal flux (clean) 5.9 X 1014
Average thermal flux (average over life) ~3 X 1014
Average fast flux (clean) ~6 X 1014
Average moderator temperature 800°C
Temperature coefficient, average over fuel lifetime ~0.5 X 10~48k/°C
Critical mass (clean, enriched U) | 5.6 kg

Critical mass (xenon at equilibrium, enriched U) 6.3 kg
Inventory (xenon at equilibrium, enriched U) 7.6 kg
Inventory volume = 1.2 core volume of bismuth 24 ft3

U235 in system at end of fuel lifetime 23.8 kg

Reflector 1.9 g/cc graphite
Reflector thickness 1.5ft

Reflector void fraction 5%,

 

 

 

 

Fuel. Uranium makeup is added to the fuel solution on a day-to-day
basis, thus keeping excess reactivity to a minimum. The operating lifetime
“of the fuel is nearly 5 yr at full power and 80% plant utilization factor. Fuel
burnup may be as much as 80%, and total U235 inventory varies from
about 7 kg at the beginning of fuel life to about 24 kg at the end of fuel life.

The LMF-GCR tends to be self-regulating. Under the influence of its
negative temperature coeflicient, the reactor will tend to operate at the
same average moderator temperature at all power levels. This temperature
will be maintained by controlling the uranium fuel solution concentration.

Spent fuel. After 4 to 5 yr, nonvolatile fission-product poisons and non-
fissionable isotopes of uranium accumulate to such an extent that a new
fuel charge is required. The used fuel is drained into the spent fuel tank and
the reactor fuel loop is then ready to receive a new fuel charge. The spent
fuel is transferred into a number of small, shielded shipping tanks for ship-
ment to a chemical processing plant.
938 ADDITIONAL LIQUID METAL REACTORS [cHAP. 25

25-1.4 Reactor materials. The critical problem associated with the
LMF-GCR is the development of a core element. As a basic core element
material graphite is extremely attractive because it is a very good moder-
ator, possesses excellent high-temperature strength, has unexcelled re-
sistance to thermal shock, is not attacked by bismuth, has a low neutron
absorption cross section, possesses a satisfactorily high thermal conduc-
tivity, and shows evidence that radiation damage is rapidly annealed at
high temperature. Presently available graphite is not impermeable to bis-
muth or gases, as the core element material of the LMF-GCR must be in
order to separate the fuel and coolant satisfactorily. However, recent de-
velopments indicate a chance for success in this area.

The other aspect of core element development is to find a suitable means
for joining the graphite to the upper and lower fuel system headers. The
graphite-to-metal bond must have adequate mechanical strength and be
resistant to corrosion, thermal cycling, and radiation damage. Bonds of
this type have been prepared by means of high-temperature brazing tech-
niques, and the work has shown that numerous additional bonding agents
are available. Preliminary work is encouraging and indicates that with
improvements in bond design, bond techniques, and test methods, solutions
to the bonding problem may be achieved.

Alternate materials as the basic core element structural material are
under investigation as a backup to the graphite development. These
include KT silicon carbide, molybdenum, molybdenum carbide, niobium,
niobium carbide, zirconium carbide, tantalum, and tantalum carbide, all
of which have properties indicating promise for LMF-GCR application.

25-1.5 Plant operation and maintenance. The LMF-GCR is primarily
self-regulating, having a temperature coeflicient of approximately —0.5 X
10—4/°C. Large changes in power output are controlled by varying coolant
flow rate while keeping the gas temperatures approximately constant.
Coolant flow rate will be varied by controlling the helium blower speed, and
by changing the coolant gas density (pressure) with the compressor and
accumulator system.

The main plant and reactor control room will be outside the reactor
containment shell in the steam plant generator building. A full thickness
of shielding wall separates the boiler and blower compartments from the
reactor, and operating personnel will be able to conduct maintenance and
inspection of these items while the reactor is in operation. This wall is
penetrated by the concentric piping which carries the primary gas into and
out of the reactor. To attenuate radiation streaming through the pipe, a
turn is made within the shield.

The core and pressure vessel assembly have been designed so that the
core, and also the reflector if necessary, may be replaced in the event of a
25-2] MOLTEN PLUTONIUM FUEL REACTOR 939

failure. During operation, the core and reflector are supported at the bot-
tom of the pressure vessel. However, the core assembly is attached to the
pressure vessel head so that the two will be lifted together when the head
1s removed. The reflector is also constructed with a metal support structure
so that it can be lifted out of the pressure vessel as a unit.

The fuel loop components and piping are arranged so that maintenance
can be carried out in a safe and reliable manner. Since the parts are rela-
tively inexpensive, it will probably be cheaper to replace than repair them.

25~1.6 Plant capital and power cost. For a 16,000-kw (electrical) LMF-
GCR plant, the cost of power, at an 80% plant utilization factor, is es-
timated at 14.6 mills/kwh, made up of 8.6 mills/kwh for fixed charges,
2.7 mills/kwh for operation and maintenance, and 3.3 mills/kwh for fuel.
The total power cost using a 60% plant utilization factor is 18.4 mills /kwh.
A fixed charge rate of 15% was used.
- The capital cost for a 16,000-kw LMF-GCR nuclear plant has been
estimated at $409/kw of installed capacity. These cost figures are based
on estimates for the important equipment in the plant, and on recent AEC
fuel prices.

25-2. MovTEN PLrutontuM FurL REACTOR*

25-2.1 Introduction. The long-range utility of nuclear power based on
uranium fission depends upon the development of a plutonium-fueled
reactor capable of being refueled by an integral, or associated, breeding
cycle. If full utilization of the energy content in the world’s supply of
uranium is to be accomplished, the more abundant U238 must be converted
into the easily fissionable isotopes of plutonium. The need for this full
utilization is apparent when it is realized that the economically recoverable
U?3 content of uranium ores [1,2] is sufficient to supply projected world
power requirements for only a few decades. Breeding on the plutonium
cycle extends fission power capabilities by a factor of 140, yielding thou-
sands, instead of tens, of years of world energy reserves. |

The high values of the capture-to-fission ratio at thermal and epithermal
neutron energies for the plutonium isotopes preclude these types of reactors
from an integral plutonium breeding cycle system. To obtain an appreci-
able breeding gain, a plutonium-fueled reactor must be either a fast or a
fast-intermediate neutron spectrum device where breeding ratios of the
order of 1.7 may be expected from suitably designed systems. One of the
power-producing reactors of the future must logically be a fast plutonium
breeder.

*This section is based largely on material from Los Alamos Scientific Laboratory,
LA2112, R. M. Kiehn.
940 ADDITIONAL LIQUID METAL REACTORS [cHAP. 25

In order to maintain a fast-neutron spectrum, fuel densities in a plu-
tonium breeder will be high, and coolants must be either molten metals or
salts. The latter characteristic will permit large amounts of power to be
extracted from relatively small volumes, thus obtaining a large specific
power. Hydrogenous and organic coolants are eliminated because of their
attendant neutron moderation properties, high vapor pressures at high
temperatures, and relatively poor resistance to radiation damage. For
efficiency reasons the system temperature should be as high as is compatible
with a long operating life. Therefore, to be in step with modern electrical
generation techniques, this would imply coolant outlet temperatures of the
order of 650°C.

25-2.2 Basic components. Before discussing the Los Alamos Molten
Plutonium Reactor (LAMPRE) proposal in detail, the following resume
will treat some of the possibilities for the three basic components of a
power reactor: the fuel, the container, and the coolant.

Molten plutonium fuels. Plutonium metal melts at 640°C, a temperature
that is somewhat high, but not beyond the bounds of utility. Fortunately,
some alloys of plutonium have significantly lower melting temperatures.
Specifically, eutectic alloys of plutonium with iron, nickel, and cobalt
all have melting temperatures in the vicinity of 400 to 450°C. Ternary
and quaternary alloying agents will further lower these melting tempera-
tures by a few percent. One characteristic of these transition metal alloys
is that they do not dilute the fuel volumetrically to a great extent in their
eutectic compositions.

Other alloys of plutonium which are more dilute in fuel and have not
too unreasonable melting temperatures are the magnesium-plutonium and
bismuth-plutonium alloys. The spatial dilution of fuel atoms alleviates the
high power density problem but, unfortunately, these alloys have melting
temperatures significantly higher than the transition metal alloys.

A compilation of the interesting fuel alloys, their melting points, and
eutectic compositions appears in Table 25-2.

Container materials. A material capable of being fabricated into various
shapes and resistant to high-temperature corrosion by the fuel alloy is a
necessity if practical use is to be made of the low melting temperature
plutonium alloys. Since the transition metals readily form low melting
point alloys with plutonium, the normal constructional materials, steels
and nickel alloys, are eliminated.

The next alternatives, the refractory metals, have been used with meas-
urable success to contain the various alloys of plutonium. Tungsten and
tantalum have been somewhat better containers than molybdenum and
niobium and much better than chromium, vanadium, and titanium. The
requirement of fabricability eliminates several of the refractory metals, such
25-2] MOLTEN PLUTONIUM FUEL REACTOR 941

as tungsten and molybdenum, because of the poor state of their peculiar
welding art.

The limitations of metallurgical knowledge at present lead to the con-
clusion that tantalum will be one of the best container materials for these
plutonium alloys. The high-temperature strength properties and the heat-
transfer properties of tantalum are excellent; moreover, it is weldable. The
parasitic capture cross section of tantalum would be intolerable in an
epithermal or thermal power breeder reactor and, although relatively
large in a fast spectrum, its effect on neutron economy in a fast reactor can
be made small, if not minor, by careful design.

Dynamic corrosion tests indicate that tantalum’s resistance to corrosion
by molten sodium, a possible coolant, will be adequate. Long-term static
corrosion tests (9000 hr at 650°C) indicate that the fuel is compatible with
tantalum at proposed operating temperatures.

Coolant. The desire to obtain a high power density at high temperatures
and low pressures in a high radiation field dictates the use of molten metal
or salt coolant. The list of possibilities is topped by sodium and bismuth.
A few words about the properties of these coolants are probably appro-
priate at this point.

Sodium is advantageous because of its low melting point, good heat-
transfer properties, low pumping power requirement, and because there
has been considerable engineering experience with it. Its poor long-term
corrosion properties when in contact with the better container materials
such as tantalum and its explosive burning property when exposed to water
or moist air are distinct disadvantages. |

Bismuth, on the other hand, does not react explosively with water, nor
does it burn in air. Pumping power requirements some five times larger
than for sodium, its higher melting temperature, and the polonium buildup
problems are disadvantageous factors of a bismuth coolant. However,

 

 

 

TABLE 25-2
FueL ALLoYys
Eutectic Melting Approximate |
Alloy composition, point, ~ density,
a/o0 °C | g/cc

Pu-Fe 9.5 Fe 410 16.8
Pu—Co 10 Co 405 16
Pu-Ni 12.5 Ni 465 16
Pu-Mg 85 Mg 552 3.4
Pu-Bi Noneutectic 271-900

 

 

 

 

 

 
942 ADDITIONAL LIQUID METAL REACTORS [cHAP. 25

the corrosion resistance of tantalum in dynamie, high-temperature bis-
muth is excellent, according to the Ames experiments [3].

25-2.3 LAMPRE. A first step in solving the plutonium power reactor
problem is to prove the feasibility of operating and maintaining a molten
plutonium power reactor core. To this end, the reactor assembly known as
LAMPRE I has been devised. The LAMPRE system has the following

essential features:

Fuel alloy: Molten plutonium-iron
(eutectic composition, 9.5 a/0 Fe)

Container: Tantalum

Reflector: Steel

Shield: Graphite, iron, concrete

Coolant: Sodium ‘

Power: 1 Mw heat

Heat transfer: Internally cooled core
Tube-shell

Heat exchanger
Heat rejected to air

Breeding No breeding blanket

Core. The LAMPRE core consists of three parts: fuel alloy, container,
and coolant. A proposed design, described in detail below, yields a struc-
ture which is approximately 509, by volume fuel alloy, 159, structure, and
359, coolant. The minimum tube separation 1s slightly under 1/16 1n.
At reasonable heat-transfer rates, this configuration is capable of develop-
ing a specific power of better than 250 watts/g. More efficient systems can
utilize 2 similar structure but must dilute the fuel volumetrically to ob-
tain a larger heat-transfer surface per unit of contained fuel. The larger
area-to-volume ratio can be obtained by going to smaller diameter tubes
and /or closer spacing of the tube array. In the tube-shell arrangement,
the fuel is located on the outside of the tubes and the coolant flows through
the tubes. Such a scheme preserves the volumetric integrity of the fuel.
Other radiator-type schemes, which also preserve fuel integrity, are
conceivable.

The over-all assembly will be designed so that the core will be completely
filled during operating conditions. The estimated core height is 6.5 1in.
25-2] MOLTEN PLUTONIUM FUEL REACTOR 943

Tantalum expansion, filling, and draining tubes will be attached to the
core structure. A reference core assembly would be:

Container Tantalum
Tubes (547) 3/16-in. OD, 0.015-in. wall,
hexagonal array
Cage shape Right cylinder,
6.25-in. OD, 6.5-in. height
Headers and shell 0.040 to 0.080 in.
Critical mass 26 kg plutonium alloy

Reflector. No attempt to breed will be carried out in the first LAMPRE
concept. Although the over-all coolant container will be made of stainless
steel, the fast-neutron reflector will be made of steel and will be cooled
by the main sodium stream. The thickness of the radial steel reflector will
be adjusted to be thin enough, neutronwise, to obtain adequate external
reflector control, but will be too thick to allow the thermalized neutrons
returning from the graphite shield to build up a power spike at the core
surface. The core, although slightly coupled to the reflector and shield, will
have a mean fission energy greater than 500 kev, ensuring a high possible
breeding gain. The top and bottom stainless-steel reflector slugs will also
be sodium-cooled and will be essentially “infinitely’’ thick to fast neutrons.
The coolant channels will be drilled or machined into solid slug or disk
castings. |

Control. The control of LAMPRE will be effected by reflector-type
mechanisms. An annular shim control displacing the innermost 4 in. of
shield with aluminum will be used as a coarse criticality adjustment
mechanism. Several replacement cylinders, replacing the inner portions
of aluminum with void, will be used as fine controls. A rotating control
cylinder will be built into the system in anticipation of safety and neutron
kinetics experiments. |

The radial thickness of the steel fast-neutron reflector is adjusted so that
the fast and intermediate neutrons returning to the core from the aluminum
reflector and graphite are worth approximately 109 to the core critical
mass. Displacement of the aluminum reflector effectively reduces the
neutron reflection back to the core, yielding an external, large-effect control
mechanism adequately cooled by aluminum conduction and air convention.

The LAMPRE critical experiments have proved that aluminum-void
replacement mechanisms are effective and operable. The annular shim
has been shown to be almost ineffective at distances greater than 2 in.
944 ADDITIONAL LIQUID METAL REACTORS [cHAP. 25

SODIUM N

 
 
   
   

~—SODIUM QUT

e\ ESSEL
SECOND CONTAINMENT

TOP REFLECTOR

 

 

GAS VENT

SHIM

 

HEAT EX?HANGER

FUEL CONTAINER BAYONE

BOTTOM REFLECTOR

 

 

CONTROL ROD THIMBLE
FUEL ADDITION LINE

 

RADIAL REFLECTOR

 

FUEL RESERVOIR

 

CATCH PAN

LOCATING PIN

Fia. 25-5. The Los Alamos Molten Plutonium Reactor Experiment.

above or below the core height for the geometry. These results have been
incorporated into the LAMPRE design as presented in Fig. 25-5.

25-3. Liquip METAL-URANIUM OXIDE SLURRY REACTORS

There has been some work done at other locations on uranium oxide
slurry reactors. At Knolls Atomic Power Laboratory, a uranium oxide-
bismuth slurry reactor has been explored [4]. In this reactor, the fuel,
consisting of uranium oxide suspension and liquid bismuth, is pumped
through a moderator matrix and then through an external heat exchanger.
The reader will recognize that this is the same as the single-region LMFR
described in the preceding chapter.
' REFERENCES 945

The studies at KAPL were encouraging. A small amount of experimental
work indicated that dispersions of uranium oxide and bismuth can be made.
These workers found that at 500 to 600°C titanium is the best additive for
promoting the wetting of UO2 by bismuth. An 8 w/0 UQOgz-bismuth slurry
was actually pumped with an electromagnetic pump at 450°C.

At Argonne National Laboratory, uranium oxide-NaK slurries have been
studied as possible reactor fuels [5]. This fuel would be suitable for a
fast-breeder reactor. Investigations have been carried out at a maximum
concentration of 4.3 vol. % UO2 in eutectic NaK. Two loops have been
operated at temperatures ranging from 450 to 600°C. A slurry with 4.3
vol. % actually has a very high weight percent, 36.0 w/o.

The tests in the two loops indicated uniform suspension at flow rates of
2 fps. |

The UO2 dropped out of suspension at temperatures above 500°C but
would resuspend at lower temperatures. When a very small amount of
uranium metal was added to the slurry, better wetting of the particles
was obtained and no further settling above 500°C was observed.

Work on the uranium oxide slurries is continuing, and the incorporation
of these results into liquid metal fuel reactors can be expected.

REFERENCES

1. S. GLASSTONE, Principles of Nuclear Reactor Engineering. Princeton, N. J.:
D. Van Nostrand Co., Inc., 1955. (pp. 1-2)

2. P. C. PurNaMm, Energy tn the Future. Princeton, N. J.: D. Van Nostrand
Co., Inc., 1953. (p. 214)

3. R. W. FisuEr and G. R. WinpERs, High Temperature Loop for Circulating
Liquid Metals, in Chemical Engineering Progress Symposium Series, Vol. 53,
No. 20. New York: American Institute of Chemical Engineers, 1957. (pp. 1-6)

4. D. H. AEMANN et al.,, 4 UOo-Bismuth System As a Reactor Fuel, USAEC
Report KAPL-1877, Knolls Atomic Power Laboratory, July 1, 1957.

5. B. M. ABraHAM et al., UO2-NaXK Slurry Studies in Loops to 600°C, Nuclear
Sci. and Eng. 2, 501-512 (1957).