FFR_part2.txt

Part 11

MOLTEN-SALT REACTORS

11.
12.
13.
14.
15.
16.
17.

 

H. G. MacPrERsoN, Editor
Oak Ridge N atvonal Laboratory

Introduction

Chemical Aspects of Molten-Fluoride-Salt Reactor Fuels
Construction Materials for Molten-Salt Reactors

Nuclear Aspects of Molten-Salt Reactors

Equipment for Molten-Salt Reactor Heat-Transfer Systems

Aircraft Reactor Experiment

Conceptual Design of a Power Reactor
CONTRIBUTORS

L. G. ALEXANDER H. G. MAcPHERSON

J. W. ALLEN W. D. MANLY

E. S. BerTIs L. A. MANN

F. F. BLANKENSHIP W. B. McDoNALD

W. F. BoubprEAU H. J. METZ

E. J. BREEDING P. PATRIARCA

W. G. CoBB H. F. PoPPENDIEK

W. H. Cooxk J. T. RoBERTS

D. R. CuNEko M. T. RoBINSON

J. H. DEVANN T. K. RocHE

D. A. DougLas H. W. SAVAGE

W. K. ERGEN G. M. SLAUGHTER

W. R. GRIMES E. StorTO

H. INoUYE A. TaBoADA

D. H. JANSEN G. M. TorsoN

G. W. KeLaOoLTZ F. C. VONDERLAGE

B. W. Kinvon - G. D. WHITMAN

M. E. LackEy J. ZASLER
PAUL ERSKINE BROWN
PREFACE

The Oak Ridge National Laboratory, under the sponsorship of the
U. S. Atomic Energy Commission, has engaged in research on molten
salts as materials for use in high-temperature reactors for a number of
years. The technology developed by this work was incorporated in the
Aircraft Reactor Experiment and made available for purposes of civilian
apphcatlon This earlier technology and the new information found in the
civilian power reactor effort is summarized in this part.

So many present and former members of the Laboratory staff have
contributed directly or indirectly to the molten salt work that it should be
regarded as a contribution from the entire Laboratory. The technical
direction of the work was provided by A. M. Weinberg, R. C. Briant,
W. H. Jordan, and S. J. Cromer. In addition to the contributors listed for
the various chapters the editor would like to acknowledge the efforts of
the following people who are currently engaged in the work reported:
R. G. Affel, J. C. Amos, C. J. Barton, C. C. Beusman, W. E. Browning,
S. Cantor, D O. Campbell G. L. Cathers B. H. Clampltt J. A. Conlin,
M. H. Cooper, J. L. Crowley, J. Y. Estabrook, H. A. Friedman, P. A.
Gnadt, A. G. Grindell, H. W. Hoffman, H. Insley, S. Langer, R. E. Mac-
Pherson R. E. Moore, G. J. Nessle, R. F. Newton, W. R. Osborn, F. E.
Romie, C. F. Sales, J. H. Shaffer, G. P. Smith, N. V. Smith, P. G. Smith,
W. L. Snapp, W. K. Stair, R. A. Strehlow, C. D. Susano, R. E. Thoma,
D. B. Trauger, J. J. Tudor, W. T. Ward, G. M. Watson, J. C. White, and
H. C. Young.

The technical reviews at Argonne National Laboratory and Westing-
house Electric Corporation aided in achieving clarity.

The editor and contributors of this part wish to express their apprecia-
tion to A. W. Savolainen for her assistance in preparing the text in its
final form.

Oak Ridge, Tennessee H. G. MacPherson, Editor
June 1958
CHAPTER 11
INTRODUCTION*

The potential utility of a fluid-fueled reactor that can operate at a high
temperature but with a low-pressure system has been recognized for a
long time. Some years ago, R. C. Briant of the Oak Ridge National Lab-
oratory suggested the use of the molten mixture of UF4 and ThF4, together
with the fluorides of the alkali metals and beryllium or zirconium, as the
fluid fuel. Laboratory work with such mixtures led to the operation, in
1954, of an experimental reactor, which was designated the Aircraft Reactor
Experiment (ARE).

Fluoride-salt mixtures suitable for use in power reactors have melting
points in the temperature range 850 to 950°F and are sufficiently compatible
with certain nickel-base alloys to assure long life for reactor components at
temperatures up to 1300°F. Thus the natural, optimum operating tem-
perature for a molten-salt-fueled reactor is such that the molten salt is a
suitable heat source for a modern steam power plant. The principal
advantages of the molten-salt system, other than high temperature, in
comparison with one or more of the other fluid-fuel systems are (1) low-
pressure operation, (2) stability of the liquid under radiation, (3) high
solubility of uranium and thorium (as fluorides) in molten-salt mixtures,
and (4) resistance to corrosion of the structural materials that does not
depend on oxide or other film formation.

The molten-salt system has the usual benefits attributed to fluid-fuel
systems. The principal advantages over solid-fuel-element systems are
(1) a high negative temperature coefficient of reactivity, (2) a lack of radia-
tion damage that can limit fuel burnup, (3) the possibility of continuous
fission-product removal, (4) the avoidance of the expense of fabricating
new fuel elements, and (5) the possibility of adding makeup fuel as needed,
which precludes the need for providing excess reactivity. The high negative
temperature coeflicient and the lack of excess reactivity make possible a
reactor, without control rods, which automatically adjusts its power in re-
sponse to changes of the electrical load. The lack of excess reactivity also
leads to a reactor that is not endangered by nuclear power excursions.

One of the attractive features of the molten-salt system is the variety of
reactor types that can be considered to cover a range of applications. The
present state of the technology suggests that homogeneous reactors which
use a molten salt composed of Bel's and either Li’F or NaF, with UF4 for
fuel and ThF4 for a fertile material, are most suitable for early construction.

*By H. G. MacPherson.
567
968 INTRODUCTION [cHAP. 11

These reactors can be either one or two region and, depending on the size
of the reactor core and the thorium fluoride concentration, can cover a
wide range of fuel inventories, breeding ratios, and fuel reprocessing sched-
ules. The chief virtues of this class of molten-salt reactor are that the design
is based on a well-developed technology and that the use of a simple fuel
cycle contributes to reduced costs.

With further development, the same base salt, that is, the mixture of
BeF2 and Li’F, can be combined with a graphite moderator in a hetero-
geneous arrangement to provide a self-contained Th-U233 system with a
breeding ratio of one. The chief advantage of the molten-salt system over
other liquid systems in pursuing this objective is that it is the only system
in which a soluble thorium compound can be used, and thus the problem
of slurry handling is avoided. The possibility of placing thorium in the
core obviates the necessity of using graphite as a core-shell material.

Plutonium is being investigated as an alternate fuel for the molten-salt
reactor. Although it is too early to describe a plutonium-fueled reactor in
detail, it is highly probable that a suitable PuFs-fueled reactor can be
constructed and operated.

The high melting temperature of the fluoride salts is the principal dif-
ficulty in their use. Steps must be taken to preheat equipment and to keep
the equipment above the melting point of the salt at all times. In addition,
there is more parasitic neutron capture in the salts of the molten-salt
reactor than there is in the heavy water of the heavy-water-moderated
reactors, and thus the breeding ratios are lower. The poorer moderating
ability of the salts requires larger critical masses for molten-salt reactors
than for the aqueous systems. Finally, the molten-salt reactor shares with
all fluid-fuel reactors the problems of certain containment of the fuel, the
reliability of components, and the necessity for techniques of making
repairs remotely. The low pressure of the molten-salt fuel system should
be beneficial with regard to these engineering problems, but to evaluate
them properly will require operating experience with experimental reactors.
CHAPTER 12

CHEMICAL ASPECTS OF MOLTEN-FLUORIDE-SALT
REACTOR FUELS*

The search for a liquid for use at high temperatures and low pressures
in a fluid-fueled reactor led to the choice of either fluorides or chlorides
because of the requirements of radiation stability and solubility of appre-
ciable quantities of uranium and thorium. The chlorides (based on the CI37
isotope) are most suitable for fast reactor use, but the low thermal-neutron
absorption cross section of fluorine makes the fluorides a uniquely desirable
choice for a high-temperature fluid-fueled reactor in the thermal or epi-
thermal neutron region.

Since for most molten-salt reactors considered to date the required con-
centrations of UF4 and ThF4 have been moderately low, the molten-salt
mixtures can be considered, to a first approximation, as base or solvent
salt mixtures, to which the fissionable or fertile fluorides are added. For
the fuel, the relatively small amounts of UF4 required make the correspond-
ing binary or ternary mixtures of the diluents nearly controlling with regard
to physical properties such as the melting point. |

12-1. CHOICE OF BASE OR SOLVENT SALTS

The temperature dependence of the corrosion of nickel-base alloys by
fluoride salts is described in Chapter 13. From the data given there, 1300°F
(704°C) is taken as an upper limit for the molten-salt-to-metal interface
temperature. To provide some leeway for radiation heating of the metal
walls and to provide a safety margin, the maximum bulk temperature of
the molten-salt fuel at the design condition will probably not exceed 1225°F.
In a circulating-fuel reactor, in which heat is extracted from the fuel in an
external heat exchanger, the temperature difference between the inlet and
outlet of the reactor will be at least 100°F. The provision of a margin of
safety of 100°F between minimum operating temperature and melting
point makes salts with melting points above 1025°F of little interest at
present, and therefore this discussion is limited largely to salt mixtures
having melting points no higher than 1022°F (550°C). One of the basic
features desired in the molten-salt reactor is a low pressure in the fuel
system, so only fluorides with a low vapor pressure at the peak operating
temperature (~700°C) are considered.

*By W. R. Grimes, D. R. Cuneo, F. F. Blankenship, G. W. Keilholtz, H. F.
Poppendiek, and M. T. Robinson.

569
570 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [cHAP. 12

 

710°C

 

 

KF

 

—— LiF
454°c”  492°C

F1c. 12-1. The system LiF-NaF-KF [A. G. Bergman and E. P. Dergunov,
Compt. rend. acad. sci. U.R.S.S., 31, 7564 (1941)].

Of the pure fluorides of molten-salt reactor interest, only BeFs meets
the melting-point requirement, and it is too viscous for use in the pure
state. Thus only mixtures of two or more fluoride salts provide useful
melting points and physical properties.

The alkali-metal fluorides and the fluorides of beryllium and zirconium
have been given the most serious attention for reactor use. Lead and bis-
muth fluorides, which might otherwise be useful because of their low neutron
absorption, have been eliminated because they are readily reduced to the
metallic state by structural metals such as iron and chromium.

Binary mixtures of alkali fluorides that have sufficiently low melting
points are an equimolar mixture of KF and LiF, which has a melting point
of 490°C, and a mixture of 60 mole 9, RbF with 40 mole 9, LiF, which has
a melting point of 470°C. Up to 10 mole 9, UF4 can be added to these
alkali fluoride systems without increasing the melting point above the
550°C limit. A melting-point diagram for the ternary system LiF-NaF-KF,
Fig. 12-1, indicates a eutectic with a lower melting point than the melting
points of the simple binary LiF-KF system. This eutectic has interesting
properties as a heat-transfer fluid for molten-salt reactor systems, and
data on its physical properties are given in Tables 12-1 and 12-2. The
KF-LiF and RbF-LiF binaries and their ternary systems with NaF are
the only available systems of the alkali-metal fluorides alone which have
571

CHOICE OF BASE OR SOLVENT SALTS

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572 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [CHAP. 12

Temperature, °C

2NaF-ZrFy

e

b
™
(T

O
Z
™

 

NaF 10 20 30 40 50 60 70 80 90 ZrFy
ZrF4 ,mole %

F1g. 12-2. The system NaF-ZrFy,.

low melting points at low uranium concentrations. They would have
utility as special purpose reactor fuel solvents if no mixtures with better
properties were available.

TABLE 12-2

THERMAL CoNDUCTIVITY OF TYPICAL FLUORIDE MIXTURES

 

 

 

 

Thermal conductivity,
Composition, Btu/ (hr) (ft) (°F)
mole 9,
Solid Liquid
LiF-NaF-KF (46.5-11.5-42) 2.7 2.6
NaF-BeF2 (57-43) 2.4

 

 

 

 

 

Mixtures with melting points in the range of interest may be obtained
over relatively wide limits of concentration if ZrF4 or BeF; is a component
of the system. Phase relationships in the NaF-ZrF4 system are shown in
Fig. 12-2. There is a broad region of low-melting-point compositions that
have between 40 and 55 mole 9, ZrF 4.
12-1]

°C

-

Temperature

700
(2
o
@ 600
2
o
@
E

50
8 0

800

400

300

200 L
LiF

CHOICE OF BASE OR

Fia. 12-3.

SOLVENT SALTS 573

BeFg + Liquid

LigBeFy
+ Liquid

 

BeFy mole % LiBeF3 + BeFg

The system LiF-BeFs.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

900 \ ¥
800 \
NaF + Liquid
700 oe=0rnl Data
a — Na,BeF, 4 LIQUID
2°°74
600 p /’
/ ’
a — NagBeF 4 + NaF \ i - SSIB;E3 BeF, -+ LIQUID
400 BeFy + 8'— NaBeF3
, /Z " o !
a— NCII2BeF4+Bl—NOBeF,3\* \ l
A ! NoBoF 8 — NaBeF4 + LIQUID
300 '~ NagBeF 4 +NaF B~ Nalefs | [BeF5+ 6 — NabeFg
B8 — NaBeF;, +v — Na,BeF , < a’'— NanBeF BeF — NoBeF
1BeFy+7 - NagBefyrs 0284 | BeFp+ - Nabey
200 ¥ — NagBeF4 + NaF = 1 I
NaF 10 20 30 40 50

60 70 80 90 BeF2

BeF,, mole %

Fic. 12-4. The system NaF-BeF.
574 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [cHAP. 12

 

BeF,
Dotted Lines Represent 542
Incompletely Defined \
Phase Boundaries and
Alkemade Lines
The Symbol TC Represents All Temperatures Are in °C
a Compound Whose Exact
Composition Has Not 370
Been Determined NaF-BeF9
TC
356 P 345
< 350
(LiF-BeF9) 28/0 7& 50— 400
400~ 2 NaF (-2 BeFs 4 fl; ].. s
4555 BT “%fifiy’ =X/ /"
5 e 7 /"‘ 2 Na5F°BeF2

 

e, s ] P

grum———= (- -
BT ——— | mBS R
. (/ /
BT SN P
750; “ ‘ (NaF-LiF-BeFp)) ~Z X | 850
800 " 240 F 5";%336&) . 200
[ / \

LiF 800 750 700 649 700 750 800 850 900 950 NaF

F16. 12-5. The system LiF-NaF-BeF's.

The lowest melting binary systems are those containing BeFs and LiF
or NaF. Since BeF'2 offers the best cross section of all the useful diluents,
fuels based on these binary systems are likely to be of highest interest in
thermal reactor designs. |

The binary system LiF-BeF: has melting points below 500°C over the
concentration range from 33 to 80 mole 9, BeF2. The presently accepted
LiF-BeF 2 system diagram presented in Fig. 12-3 differs substantially from
previously published diagrams [1-3]. It is characterized by a single eutectic
between BeF2 and 2LiF - BeF2 that freezes at 356°C and contains 52 mole
% BeFs. The compound 2LiF - BeF2 melts incongruently to LiF and
liquid at 460°C; LiF - BeF'2 is formed by the reaction of solid BeFs and
solid 2LiF - BeF2 below 274°C.

The diagram of the NaF-BeF2 system (Fig. 12-4) is similar to that of
the LiF'-BeF; system. The ternary system combining both NaF and LiF
with Bel's, shown in Fig. 12-5, offers a wide variety of low-melting compo-
sitions. Some of these are potentially useful as low-melting heat-transfer
liquids, as well as for reactor fuels.
575

CHOICE OF BASE OR SOLVENT SALTS

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576 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [cHAP. 12

TABLE 124

TaERMAL ConbpucTIiviTY OF TYPICAL FLUORIDE FUELS

 

Thermal conductivity,
Composition, Btu/ (hr) (ft) °F)
mole 9,

 

Solid Liquid

 

 

LiF-NaF-KF-UF,4 (44.5-10.9-43.5-1.1)
NaF-ZrF4,~UF, (50-46-4)
NaF-ZrF,~UF4 (53.5-40-6.5)
NaF-KF-UF, (46.5-26-27.5)

SN
o O
O = == DN
O W W

 

 

 

 

 

 
  
    
 
   

All Temperatures Are in °C

5NaF-2ZrF4 2NaF-ZrF4 3 NaF:2 ZrFy

ZrFy
918
00

Fi1g. 12-6. The system NaF-ZrF,;-UF,.
12-2] FUEL AND BLANKET SOLUTIONS 577

UFy

All Temperatures Are in °C
E—Eutectic
P—Peritectic LiF- 4UF,

:Primary Phase Field

  
 

\“\

 

 

EAN .\
PN T, MLF-4UR
ALiF-UFy A AN
D
Q e,
N\ 7LiF- 6UF,
/\. o \O \
NS 2
e = |LiF.X¢ < \
00\/%\ \7‘3‘0\\ 5 ‘}-:;T\‘ 7=\ E
LiF 2LiF-BeF,” B 400”400 BeF,
A‘\;
3 ]

Fi1a. 12-7. The system LiF-BeF3-UF 4.

12-2. FuEL AND BLANKET SOLUTIONS

12-2.1 Choice of uranium fluoride. Uranium hexafluoride is a highly
volatile compound, and it is obviously unsuitable as a component of a
liquid for use at high temperatures. The compound UO2F3, which is rela-
tively nonvolatile, is a strong oxidant that would be very difficult to con-
tain. Fluorides of pentavalent uranium (UF5,UsFy, etc.) are not thermally
stable [4] and would be prohibitively strong oxidants even if they could be
stabilized in solution. Uranium trifluoride, when pure and under an inert
atmosphere, is stable even at temperatures above 1000°C [4,5]; however,
it is not so stable in molten fluoride solutions [6]. It disproportionates
appreciably in such media by the reaction

4 UF; =3 UF4++4 U9,

at temperatures below 800°C. Small amounts of UF'3 are permissible in the
presence of relatively large concentrations of UF4 and may be beneficial
insofar as corrosion is concerned. It is necessary, however, to use UF4
as the major uraniferous compound in the fuel.
578 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [cHAP. 12

UF4

   
       
   
 

All Temperatures Are in °C

E—Eutectic
P—Peritectic

NaU4F17 —Primary Phase Field

548°

NGF‘2UF4

\

E N\

:\\\\

5NaF-3UF4 4“@

INaF-UF.F \"‘ N
PNl \650)

A VN

R
e\ Nm

X
- NaF E

2NaF-BeF9 NaF-BeF,

Fig. 12-8. The system NaF-BeF2-UF4.

12-2.2 Combination of UF,; with base salts. The fuel for the Aircraft
Reactor Experiment (Chapter 16) was a mixture of UF4 with the NaF-ZrF4
base salt. The ternary diagram for this system is shown in Fig. 12-6. The
compounds ZrF4 and UF4 have very similar unit cell parameters [4] and
are isomorphous. They form a continuous series of solid solutions with a
minimum melting point of 765°C for the solution containing 23 mole 9,
UF4. This minimum is responsible for a broad shallow trough which pene-
trates the ternary diagram to about the 45 mole 9, NaF composition. A
continuous series of solid solutions without a maximum or a minimum
exists between a—3NaF - UF4 and 3NaF - ZrF4; in this solution series the
temperature drops sharply with decreasing ZrF4 concentration. A con-
tinuous solid-solution series without a maximum or a minimum also exists
between the isomorphous congruent compounds 7NaF - 6UF4+ and
7NaF - 6ZrF4; the liquidus decreases with increasing ZrF4 content. These
two solid solutions share a boundary curve over a considerable composition
range. The predominance of the primary phase fields of the three solid
solutions presumably accounts for the complete absence of a ternary
eutectic in this complex system. The liquidus surface over the area below
8 mole 9, UF4 and between 60 and 40 mole 9, NaF is relatively flat. All
fuel compositions within this region have acceptable melting points. Minor
12-2] FUEL AND BLANKET SOLUTIONS 579

ThF,4
1080°C

\060
A000"

950°

900°

850°

 

 

LiF .
845°C ligBeF,  LiBeFs (?) 543°C
475°C 360°C

 

F1ac. 12-9. The system LiF-BeF2-ThFy.

advantages in physical and thermal properties accrue from choosing mix-
tures with minimum ZrF4 content in this composition range. Typical
physical and thermal properties are given in Tables 12-3 and 12—4.

The nuclear studies in Chapter 14 indicate that the combination of
BeFs with NaF or with LiF (provided the separated Li? isotope can be
used) are more suitable as reactor fuels. The diagram of Fig. 12-7 reveals
that melting temperatures below 500°C can be obtained over wide com-
position ranges in the three-component system LiF-BeFs-UF4. The lack
of a low-melting eutectic in the NaF-UF4 binary system is responsible for
melting points below 500°C being available over a considerably smaller
concentration interval in the NaF-BeF2:-UF4 system (Fig. 12-8) than in
its LiF-BeFo—UF4 counterpart.

The four-component system LiF-NalF-BeF.-UF4 has not been com-
pletely diagrammed. It is obvious, however, from examination of Fig. 12-5
that the ternary solvent LiF-NaF-BeF's offers a wide variety of low-melting
compositions; it has been established that considerable quantities (up to
at least 10 mole 9,) of UF4 can be added to this ternary system without
elevation of the melting point to above 500°C.
o980 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [CHAP. 12

    
   
 

   

 

  

 

7 LiF-6 ThF4
45
LiF-2 ThF4
40 1 E—Eutectic
\? P—Peritectic
o 35 Liquidus
s? P—59 Temperatures
& Are in °C
év 30
K E-560
3 LiF+ThF,
E~-570
20 550 %50
15 -
650%00
10 700 0
750 \
5 800 e T
\ : 450
LI VoAV W NN NN N
5 10 15 20 25 30 35 40 45
2 LiF-BeFy P-490

BeF9 (mole %)

Fia. 12-10. The system LiF-BeF3:-ThF4 in the concentration range 50 to 100
mole 9, LiF.

12-2.3 Systems containing thorium fluoride. All the normal compounds
of thorilum are quadrivalent; accordingly, any use of thorium in molten
fluoride melts must be as ThF4. A diagram of the LiF-BeFo—ThF4 ternary
system, which is based solely on thermal data, is shown as Fig. 12-9.
Recent studies in the 50 to 100 mole 9, LiF concentration range have
demonstrated (Fig. 12-10) that the thermal data are qualitatively correct.
Breeder reactor blanket or breeder reactor fuel solvent compositions in
which the maximum ThF4 concentration is restricted to that available in
salts having less than a 550°C liquidus may be chosen from an area of the
phase diagram (Fig. 12-10) in which the upper limits of ThF4 concentra-
tion are obtained in the composition

75 mole 9, LiF-16 mole 9, ThF+-9 mole % BeFq,
69.5 mole 9, LiF-21 mole 9%, ThF 4+9.5 mole 9, BeFq,
68 mole 9, LiF-22 mole 9, ThF4+~10 mole 9, BeFs.

12-2.4 Systems containing Thy and UFs;. The LiF-BeF>-UF4 and the
LiF-BeF2-ThF, ternary systems are very similar; the two eutectics in the
LiF-BeFs-ThF4 system are at temperatures and compositions virtually
identical with those shown by the UF4-bearing system. The very great
12-3] PROPERTIES OF FLUORIDE MIXTURES 581

similarity of these two ternary systems and preliminary examination of
the LiF-Bel's>-ThF4+~UF4 quaternary system suggests that fractional re-
placement of UF4 by ThF4 will have little effect on the freezing tem-
perature over the composition range of interest as reactor fuel.

12-2.5 Systems containing PuF3. The behavior of plutonium fluorides
in molten fluoride mixtures has received considerably less study. Plu-
tonium tetrafluoride will probably prove very soluble, as have UF4 and
ThF4, in suitable fluoride-salt diluents, but is likely to prove too strong an
oxidant to be compatible with presently available structural alloys. The
trifluoride of plutonium dissolves to the extent of 0.25 to 0.45 mole 9 in
LiF-Bel's mixtures containing 25 to 50 mole 9, BeF2. As indicated in
Chapter 14, it is believed that such concentrations are in excess of those
required to fuel a high-temperature plutonium burner.

12-3. PHYSICAL AND THERMAL PROPERTIES OF FLUORIDE MIXTURES

The melting points, heat capacities, and equations for density and vis-
cosity of a range of molten mixtures of possible interest as reactor fuels are
presented above in Tables 12-1 and 12-3, and thermal-conductivity values
are given in Tables 12-2 and 12-4; the methods by which the data were ob-
tained are described here. The temperatures above which the materials
are completely 1in the liquid state were determined in phase equilibrium
studies. The methods used included (1) thermal analysis, (2) differential-
thermal analysis, (3) quenching from high-temperature equilibrium states,
(4) visual observation of the melting process, and (5) phase separation by
filtration at high temperatures. Measurements of density were made by
welghing, with an analytical balance, a plummet suspended in the molten
mixture. Enthalpies, heats of fusion, and heat capacities were determined
from measurements of heat liberated when samples in capsules of Ni or
Inconel were dropped from various temperatures into calorimeters; both
ice calorimeters and large copper-block calorimeters were used. Measure-
ments of the viscosities of the molten salts were made with the use of a
capillary eflux apparatus and a modified Brookfield rotating-cylinder
device; agreement between the measurements made by the two methods
indicated that the numbers obtained were within 4- 109.

Thermal conductivities of the molten mixtures were measured in an
apparatus similar to that described by Lucks and Deem [7], in which the
heating plate is movable so that the thickness of the liquid specimen can
be varied. The uncertainty in these values is probably less than 4 25%.
The variation of the thermal conductivity of a molten fluoride salt with
temperature is relatively small. The conductivities of solid fluoride mix-
tures were measured by use of a steady-state technique in which heat was
passed through a solid slab.
582 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [cHAP. 12

The vapor pressures of PuFg [8], UF4 [9], and ThF 4 are negligibly small
at temperatures that are likely to be practical for reactor operations. Of
the fluoride mixtures likely to be of interest as diluents for high-temperature
reactor fuels, only AlF3, BeF: [9], and ZrF4 [10-12] have appreciable
vapor pressures below 700°C.

Measurements of total pressure in equilibrium with NaF-ZrF4-UF,
melts between 800 and 1000°C with the use of an apparatus similar to that
described by Rodebush and Dixon [13] yielded the data shown in Table
12-5. Sense et al. [14], who used a transport method to evaluate partial

TABLE 12-5

VAPOR PRESSURES OF FLUORIDE MIXTURES CONTAINING ZRF4

 

 

 

 

Composition, Vapor pressure constants™® Vapor pressure
mole 9, o
at 900°C,
mm Hg
NaF ZrF, UF, A B
X103
100 7.792 9.171 0.9
100 12.542 11.360 617
57 43 7.340 7.289 14
50 50 7.635 7.213 32
50 46 4 7.888 7.551 28
53 43 4 7.37 7.105 21

 

 

 

 

 

 

 

 

*For the eqfiation log P (mm Hg) = A — (B/T), where T is in °K.

pressures in the NaF-ZrF4 system, obtained slightly different values for
the vapor pressures and showed that the vapor phase above these liquids
is quite complex. The vapor-pressure values obtained from both investi-
gations are less than 2 mm Hg for the equimolar NaF-ZrF4 mixture at
700°C. However, since the vapor is nearly pure ZrF4, and since ZrF4 does
not melt under low pressures of its vapor, even this modest vapor pressure
leads to engineering difficulties; all lines, equipment, and connections ex-
posed to the vapor must be protected from sublimed ZrF4 “‘snow.”
Measurements made with the Rodebush apparatus have shown that the
vapor pressure above liquids of analogous composition decreases with in-
creasing size of the alkali cation. All these systems show large negative
deviations from Raoult’s law, which are a consequence of the large, posi-
tive, excess, partial-molal entropies of solution of ZrF4. This phenomenon
has been interpreted qualitatively as an effect of substituting nonbridging
583

PROPERTIES OF FLUORIDE MIXTURES

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584 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [CHAP. 12

fluoride ions for fluoride bridges between zirconium ions as the alkali
fluoride concentration is increased in the melt [12].

Vapor pressure data obtained by the transport method for NaF-BeF;
mixtures [15] are shown in Table 12-6, which indicates that the vapor
phases are not pure BeF2. While pressures above LiF-BeF2 must be ex-
pected to be higher than those shown for NaF-BeF2 mixtures, the values
of Table 12-6 suggest that the “‘snow’” problem with BeFs mixtures is
much less severe than with ZrF4 melts.

Physical property values indicate that the molten fluoride salts are, in
general, adequate heat-transfer media. It is apparent, however, from
vapor pressure measurements and from spectrophotometric examination of
analogous chloride systems that such melts have complex structures and
are far from ideal solutions.

12—4. PRoDUCTION AND PURIFICATION OF FLUORIDE MIXTURES

Since commercial fluorides that have a low concentration of the usual
nuclear poisons are available, the production of fluoride mixtures is largely
a purification process designed to minimize corrosion and to ensure the
removal of oxides, oxyfluorides, and sulfur, rather than to improve the
neutron economy. The fluorides are purified by high-temperature treat-
ment with anhydrous HF and Hs gases, and are subsequently stored in
sealed nickel containers under an atmosphere of helium.

12—4.1 Purification equipment. A schematic diagram of the purification
and storage vessels used for preparation of fuel for the Aircraft Reactor
Experiment (Chapter 16) is shown in Fig. 12-11. The reaction vessel in
which the chemical processing is accomplished and the receiver vessel into
which the purified mixture is ultimately transferred are vertical cylindrical
containers of high-purity low-carbon nickel. The top of the reactor vessel
is pierced by a charging port which is capped well above the heated zone
by a Teflon-gasketed flange. The tops of both the receiver and the reaction
vessels are pierced by short risers which terminate in Swagelok fittings,
through which gas lines, thermowells, etc., can be introduced. A transfer
line terminates near the bottom of the reactor vessel and near the top of
the receiver; entry of this tube is effected through copper-gasketed flanges
on l-in.-diameter tubes which pierce the tops of both vessels. This transfer
line contains a filter of micrometallic sintered nickel and a sampler which
collects a specimen of liquid during transfer. Through one of the risers in
the receiver a tube extends to the receiver bottom; this tube, which is
sealed outside the vessel, serves as a means for transfer of the purified
mixture to other equipment.

This assembly is connected to a manifold through which He, H2, HF,
or vacuum can be supplied to either vessel. By a combination of large tube
12-4] PURIFICATION OF FLUORIDE MIXTURES 985

Teflon-Gasketed Flange

4-in. Dia Charging Port

I
)
\—D 3/8-in. Nickel

Transfer Line

 

 

 
  
   
 
 

w Filter

 

 

 

 

 

Reaction Vessel

| .
Receiver Vessel/ p

Fic. 12-11. Diagram of purification and storage system.

 

 

 

 

 

furnaces, resistance heaters, and lagging, sections of the apparatus can be
brought independently to controlled temperatures in excess of 800°C.

12-4.2 Purification processing. The raw materials, in batches of proper
composition, are blended and charged into the reaction vessel. The material
is melted and heated to 700°C under an atmosphere of anhydrous HF to
remove HoO with a minimum of hydrolysis. The HF is replaced with Ho
for a period of 1 hr, during which the temperature is raised to 800°C, to
reduce U5+ and Ut to U%+ (in the case of simulated fuel mixtures), and
sulfur compounds to S~ and extraneous oxidants (Fe** ™, for example) to
586 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [cHAP. 12

lower valence states. The hydrogen, as well as all subsequent reagent gases,
is fed at a rate of about 3 liters/min to the reaction vessel through the re-
ceiver and transfer line and, accordingly, it bubbles up through the molten
charge. The hydrogen is then replaced by anhydrous HF, which serves,
during a 2- to 3-hr period at 800°C, to volatilize H2S and HCI and to con-
vert oxides and oxyfluorides of uranium and zirconium to tetrafluorides at
the expense of dissolution of considerable NiF; into the melt through re-
action of HF with the container. A final 24- to 30-hr treatment at 800°C
with Ho suffices to reduce this NiFs and the contained FeF2 to soluble
metals.

At the conclusion of the purification treatment a pressure of helium
above the salt in the reactor vessel is used to force the melt through the
transfer line with its filter and sampler into the receiver. The metallic
iron and nickel are left in the reactor vessel or on the sintered nickel filter.
The purified melt is permitted to freeze under an atmosphere of helium
in the receiver vessel.

12-5. RADIATION STABILITY OF FLUORIDE MIXTURES

When fission of an active constituent occurs in a molten fluoride solu-
tion, both electromagnetic radiations and particles of very high energy
and intensity originate within the fluid. Local overheating as a consequence
of rapid slowing down of fission fragments by the fluid is probably of little
consequence in a reactor where the liquid is forced to flow turbulently and
where rapid and intimate mixing occurs. Moreover, the bonding in such
liquids is essentially completely ionic. Such a solution, which has neither
covalent bonds to sever nor a lattice to disrupt, should be quite resistant to
damage by particulate or electromagnetic radiation.

More than 100 exposures to reactor radiation of various fluoride mix-
tures containing UF4 in capsules of Inconel have been conducted; in these
tests the fluid was not deliberately agitated. The power level of each test
was fixed by selecting the U235 content of the test mixture. Thermal neu-
tron fluxes have ranged from 10! to 10!* neutrons/(cm?)(sec) and power
levels have varied from 80 to 8000 w/cm3. The capsules have, in general,
been exposed at 1500°F for 300 hr, although several tests have been con-
ducted for 600 to 800 hr. A list of the materials that have been studied is
presented in Table 12-7. Methods of examination of the fuels after irra-
diation have included (1) freezing-point determinations, (2) chemical
analysis, (3) examination with a shielded petrographic microscope, (4) as-
say by mass spectrography, and (5) examination by a gamma-ray spectro-
scope. The condition of the container was checked with a shielded metal-
lograph.

No changes in the fuel, except for the expected burnup of U235 have
been observed as a consequence of irradiation. Corrosion of the Inconel
12-5] RADIATION STABILITY OF FLUORIDE MIXTURES

TaBLE 12-7

MoLTEN SALTS WHIicH HAvE BEEN STUDIED
IN IN-PiLE CaprsULE TEsTS

587

 

 

 

 

 

Composition,
System mole %

NaF-KF-UF4 46.5-26-27.5
NaF-BeF-UF, 25-60-15
NaF-BeFo-UF4 47-51-2
NaF-BeFs-UF4 50-46-4
NaF—ZrF4—UF4 63—-25—-12
NaF-ZrF,~UF, 53.5-40-6.5
NaF-ZrF4-UF, 50-48-2
NaF-ZrFUF; 50-48-2

 

 

TABLE 12-8

DESCRIPTIONS OF INCONEL ForcED-CircuraTioN Loors
OrERATED IN THE LITR aAnp THE MTR

 

Loop designation

 

 

 

 

 

 

 

LITR LITR MTR
Horizontal Vertical Horizontal
NaF-ZrF ,~UF 4 composition,
mole 9 62.5-12.5-25 | 63-25-12 | 53.5-40-6.5

Maximum fission power, w/cm3 400 500 800
Total power, kw - 2.8 10 20
Dilution factor* 180 7.3 5
Maximum fuel temperature, °F 1500 1600 1500
Fuel temperature differential, °F 30 250 155
Fuel Reynolds number 6000 3000 5000
Operating time, hr 645 332 467
Time at full power, hr 475 235 271

 

 

*Ratio of volume of fuel in system to volume of fuel in reactor core.
588 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [CHAP. 12

capsules to a depth of less than 4 mils in 300 hr was found; such corrosion
is comparable to that found in unirradiated control specimens [16]. In
capsules which suffered accidental excursions in temperatures to above
2000°F, grain growth of the Inconel occurred and corrosion to a depth of
12 mils was found. Such Increases in corrosion were almost certainly the
result of the serious overheating rather than a consequence of the radiation
field.

Tests have also been made in which the fissioning fuel is pumped through
a system in which a thermal gradient is maintained in the fluid. These tests
included the Aircraft Reactor Experiment (described in Chapter 16) and
three types of forced-circulation loop tests. A large loop, in which the
pump was outside the reactor shield, was operated in a horizontal beam
hole of the LITR.* A smaller loop was operated in a vertical position in
the LITR lattice with the pump just outside the lattice. A third loop was
operated completely within a beam-hole of the MTR.t The operating con-
ditions for these three loops are given in Table 12-8.

The corrosion that occurred in these loop tests, which were of short
duration and which provided relatively small temperature gradients, was
to a depth of less than 4 mils and, as in the capsule tests, was comparable
to that found in similar tests outside the radiation field [16]. Therefore
it is concluded that within the obvious limitations of the experience up to
the present time there is no effect of radiation on the fuel and no accelera-
tion of corrosion by the radiation field.

12—-6. BEaAvior oF FissioN Propucts

When fission of an active metal occurs in a molten solution of its fluo-
ride, the fission fragments must originate in energy states and ionization
levels very far from those normally encountered. These fragments, how-
ever, quickly lose energy through collisions in the melt and come to equi-
librium as common chemical entities. The valence states which they ulti-
mately assume are determined by the necessity for cation-anion equivalence
in the melt and the requirement that redox equilibrium be established
among components of the melt and constituents of the metallic container.

Structural metals such as Inconel in contact with a molten fluoride so-
lution are not stable to F2, UF5, or UFs. It is clear, therefore, that when
fission of uranium as UF, takes place, the ultimate equilibrium must be
such that four cation equivalents are furnished to satisfy the fluoride ions
released. Thermochemical data, from which the stability of fission-product
fluorides in complex dilute solution could be predicted, are lacking in

*Low Intensity Test Reactor, a tank type research reactor located at Qak
Ridge, Tennessee.
TMaterials Testing Reactor, a tank type research reactor located at Arco, Idaho.
12-6] BEHAVIOR OF FISSION PRODUCTS 589

many cases. No precise definition of the valence state of all fission-product
fluorides can be given; it is, accordingly, not certain whether the fission
process results in oxidation of the container metal as a consequence of de-
positing the more noble fission products in the metallic state.

12-6.1 Fission products of well-defined valence. The noble gases. The
fission products krypton and xenon can exist only as elements. The
solubilities of the noble gases in NaF-ZrFs (53-47 mole %) [17],
NaF-ZrF4+-UF4 (50-46—4 mole %) [17], and LiF-NaF-KF (46.5-11.5-42
mole %) obey Henry’s law, increase with increasing temperature, decrease
with increasing atomic weight of the solute, and vary appreciably with
composition of the solute. The Henry’s law constants and the heats of
solution for the noble gases in the NaF-ZrF4 and LiF-NaF-KF mixtures
are given in Table 12-9. The solubility of krypton in the NaF-ZrF 4 mix-
ture appears to be about 3 X1078 moles/(cm3)(atm).

TABLE 12-9

SoLUBILITIES AT 600°C AND HEATS oF SOLUTION FOR NOBLE
GASES IN MoLTeEN FrLuoriDE MIXTURES

 

 

 

 

In NaF-ZrF4 In LiF-NaF-KF
(53—-47 mole %) (46.5-11.5-42 mole 9)
Gas
Heat of Heat of
K* solution, K* solution,
kcal/mole kcal/mole
X108 X108

Helium 21.6 +1 6.2 11.34+0.7 8.0

Neon 11.3 +£0.3 7.8 4.440.2 8.9
Argon 5.1 £0.15 8.2
- Xenon 1.9440.2 11.1

 

 

 

 

 

 

 

*Henry’s law constant in moles of gas per cubic centimeter of solvent per
atmosphere.

The positive heat of solution ensures that blanketing or sparging of the
fuel with helium or argon in a low-temperature region of the reactor cannot
lead to difficulty due to decreased solubility and bubble formation in higher
temperature regions of the system. Small-scale in-pile tests have revealed
that, as these solubility data suggest, xenon at low concentration is re-
tained in a stagnant melt but is readily removed by sparging with helium.
Only a very small fraction of the anticipated xenon poisoning was observed
590 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS  [CcHAP. 12

during operation of the Aircraft Reactor Experiment, even though the sys-
tem contained no special apparatus for xenon removal [18]. It seems
certain that krypton and xenon isotopes of reasonable half-life can be
readily removed from all practical molten-salt reactors.

Elements of Groups I-A, II-A, II11-B, and IV-B. The fission products
Rb, Cs, Sr, Ba, Zr, Y, and the lanthanides form very stable fluorides; they
should, accordingly, exist in the molten fluoride fuel in their ordinary
valence states. High concentrations of ZrF4 and the alkali and alkaline
earth fluorides can be dissolved in LiF-NaF-KF, LiFs—BeF3, or NaF-ZrF4
mixtures at 600°C. The solubilitids at 600°C of YF3 and of selected rare-
earth fluorides in NaF-ZrF4 (53-47 mole %) and LiF-BeF2 (65-35 mole %)
are shown in Table 12-10. For these materials the solubility increases

TaBLE 12-10

SoLUBILITY OF YF3 AND oF SOME RARE-EARTH FLUORIDES
IN NaAF-ZrF4 anp 1N LiIF-BgFs aT 600°C

 

 

 

 

Solubility, mole 9, MF3
Fluoride )
In NaF-ZrF, In LiF-BekFs
(57-43 mole 9) (62—-38 mole %))
YF; 3.6
LaFg 2.1
CelF'3 2.3 0.48

 

 

 

 

 

about 0.5%/°C and increases slightly with increasing atomic number in
the lanthanide series; the saturating phase is the simple trifluoride. For
solutions containing more than one rare earth the primary phase is a solid
solution of the rare-earth trifluorides; the ratio of rare-earth cations in the
molten solution is virtually identical with the ratio in the precipitated solid
solution. Quite high burnups would be required before a molten fluoride
reactor could saturate its fuel with any of these fission products.

12-6.2 Fission products of uncertain valence. The valence states as-
sumed by the nonmetallic elements Se, Te, Br, and I must depend strongly
on the oxidation potential defined by the container and the fluoride melt,
and the states are not at present well defined. The sparse thermochemical
data suggest that if they were in the pure state the fluorides of Ge, As,
Nb, Mo, Ru, Rh, Pd, Ag, Cd, Sn, and Sb would be reduced to the cor-
responding metal by the chromium in Inconel. While fluorides of some of
12-7] FUEL REPROCESSING 091

these elements may be stabilized in dilute molten solution in the melt, it is
possible that none of this group exists as a compound in the equilibrium
mixture. An appreciable, and probably large, fraction of the niobium and
ruthenium produced in the Aircraft Reactor Experiment was deposited
in or on the Inconel walls of the fluid circuit; a detectable, but probably
small, fraction of the ruthenium was volatilized, presumably as RuF's, from
the melt. -

12-6.3 Oxidizing nature of the fission process. The fission of a mole of
UF4 would yield more equivalents of cation than of anion if the noble gas
1sotopes of half-life greater than 10 min were lost and if all other elements
formed fluorides of their lowest reported valence state. If this were the
case .the system would, presumably, retain cation-anion equivalence by
reduction of fluorides of the most noble fission products to metal and
perhaps by reduction of some U4t to U3*. If, however, all the elements
of uncertain valence state listed in Article 12-6.2 deposit as metals, the
balance would be in the opposite direction. Only about 3.2 equivalents of
combined cations result, and since the number of active anion equivalents
is a minimum of 4 (from the four fluorines of UF4), the deficiency must
be alleviated by oxidation of the container. The evidence from the Aircraft
Reactor Experiment, the in-pile loops, and the in-pile capsules has not
shown the fission process to cause serious oxidation of the container; it is
possible that these experiments burned too little uranium to yield significant
results. If fission of UF4 is shown to be oxidizing, the detrimental effect
could be overcome by deliberate and occasional addition of a reducing agent
to create a small and stable concentration of soluble UF3 in the fuel
mixture.

12-7. FuEL REPROCESSING

Numerous conventional processes such as solvent extraction, selective
precipitation, and preferential ion exchange could be readily applied to
molten fluoride fuels after solution in water. However, these liquids are
readily amenable to remote handling and serve as media in which chemical
reactions can be conducted. Most development efforts have, accordingly,
been concerned with direct and nonaqueous reprocessing methods.

Recovery of uranium from solid fuel elements by dissolution of the
element in a fluoride bath followed by application of anhydrous HF and
subsequent volatilization of the uranium as UFg has been described
[19,20]. The volatilization step accomplishes a good separation from Cs,
Sr, and the rare earths, fair separation from Zr, and-poor separation from
Nb and Ru. The fission products I, Te, and Mo volatilize completely from
the melt. The nonvolatile fission products are discarded in the fluoride
solvent. Further decontamination of the UFg is effected by selective ab-
592 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [cHAP. 12

sorption and desorption on beds of NaF. At 100°C, UF¢ is absorbed on the
bed by the reversible reaction

UFe¢ (g) + 3NaF—=3NaF - UF,

which was first reported by Martin, Albers, and Dust [21]. Niobium ac-
tivity, along with activity attributable to particulate matter, is also
absorbed; ruthenium activity, however, largely passes through the bed.
Subsequent desorption of the UF¢ at temperatures up to 400°C is accom-
plished without desorption of the niobium. The desorbed UFg is passed
through a second NaF bed held at 400°C as a final step and is subsequently
recovered in refrigerated traps. The decontaminations obtained are
greater than 10° for gross beta and gamma emitters, greater than 107 for
Cs, Sr, and lanthanides, greater than 10° for Nb, and about 10% for Ru.
Uranium was recovered from the molten-salt fuel of the Aircraft Reactor
Experiment by this method, and its utility for molten-fluoride fuel systems
or breeder blankets was demonstrated. Recovery of plutonium or thorium,
however, is not possible with this process.

There are numerous possible methods for reprocessing molten-salt fuels.
The behavior of the rare-earth fluorides indicates that some decontamina-
tion of molten-fluoride fuels may be obtained by substitution of CeFs or
LaF's, in a sidestream circuit, for rare earths of higher cross section. It
seems likely that Pul's can be recovered with the rare-earth fluorides and
subsequently separated from them after oxidation to Pul4. Further, it
appears that both selective precipitation of various fission-product ele-
ments and active constituents as oxides, and selective chemisorption of
these materials on solid oxide beds are capable of development into valu-
able separation procedures. Only preliminary studies of these and other
possible processes have been made.
993

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4. J. J. Katz and E. RaBiNnowircH, The Chemastry of Uranitum, National
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5. W. R. GriMEs et al., Oak Ridge National Laboratory. Unpublished.

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11. W. FiscHER, Institut fiir Anorganische Chemie, Technicshe Hochschule,
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12. S. CanToR et al., Vapor Pressures and Derived Thermodynamic Informa-
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13. W. H. RopeBusH and A. L. DixonN, The Vapor Pressures of Metals; A
New Experimental Method, Phys. Rev. 26, 851 (1925).

14. K. A. SENSE et al., Vapor Pressure and Derived Information of the Sodium
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termination of Molecular Complexes Present in the Vapor Phase, J. Phys.
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15. K. A. SENSE et al., Vapor Pressure and Equilibrium Studies of the Sodium
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16. W. D. MANLY et al., Metallurgical Problems in Molten Fluoride Systems,
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of Atomic Energy, Geneva, 1958. |

17. W. R. GriMmEs et al., Solubility of Noble Gases in Molten Fluorides.
I. In Mixtures of NaF—ZrF4 (53-47 mole 9;) and NaF—ZrF4+—UF4 (50-46-4
mole %), J. Phys. Chem. (in press).
594 CHEMICAL ASPECTS: MOLTEN-FLUORIDE-SALT FUELS [CHAP. 12

18. E. S. BeTtis et al., The Aircraft Reactor Experiment—Operation, Nuclear
Sct. and Eng. 2, 841 (1957).

19. G. I. CatuEers, Uranium Recovery for Spent Fuel by Dissolution in Fused
Salt and Fluorination, Nuclear Sct. and Eng. 2, 768 (1957).

20. F. R. Bruck et al., in Progress tn Nuclear Energy, Series I1I, Process
Chemistry, Vol. I. New York: McGraw-Hill Book Co., Inc., 1956.

21. Von H. MaRrTIN et al., Double Fluorides of Uranium Hexafluoride, Z.
anorg. u. allgem. Chem. 265, 128 (1951).
CHAPTER 13

CONSTRUCTION MATERIALS FOR MOLTEN-SALT REACTORS*

13—-1. SURVEY OF SUITABLE MATERIALS

A molten-salt reactor system requires structural materials which will
effectively resist corrosion by the fluoride salt mixtures utilized in the core
and blanket regions. Evaluation tests of various materials in fluoride salt
systems have indicated that nickel-base alloys are, in general, superior to
other commercial alloys for the containment of these salts under dynamic
flow conditions. In order to select the alloy best suited to this application,
an extensive program of corrosion tests was carried out on the available
commercial nickel-base alloys, particularly Inconel, which typifies the
chromium-containing alloys, and Hastelloy B, which is representative of
the molybdenum-containing alloys.

Alloys contalning appreciable quantities of chromium are attacked by
molten salts, mainly by the removal of chromium from hot-leg sections
through reaction with UFy, if present, and with other oxidizing impurities
in the salt. The removal of chromium is accompanied by the formation of
subsurface voids in the metal. The depth of void formation depends
strongly on the operating temperatures of the system and on the com-
position of the salt mixture.

On the other hand, Hastelloy B, in which the chromium is replaced with
molybdenum, shows excellent compatibility with fluoride salts at tempera-
tures in excess of 1600°F. Unfortunately, Hastelloy B cannot be used
as a structural material in high-temperature systems because of its age-
hardening characteristics, poor fabricability, and oxidation resistance.

The information gained in the testing of Hastelloy B and Inconel led
to the development of an alloy, designated INOR-8, which combines the
better properties of both alloys for molten-salt reactor construction. The
approximate compositions of the three alloys, Inconel, Hastelloy B, and
INOR-S8, are given in Table 13-1.

INOR-8 has excellent corrosion resistance to molten fluoride salts at
temperatures considerably above those expected in molten-salt reactor
service; further, no measurable attack has been observed thus far in tests
at reactor operating temperatures of 1200 to 1300°F. The mechanical
properties of INOR-8 at operating temperatures are superior to those of
many stainless steels and are virtually unaffected by long-time exposure

*By W. D. Manly, J. W. Allen, W. H. Cook, J. H. DeVan, D. A. Douglas,
H. Inouye, D. H. Jansen, P. Patriarca, T. K. Roche, G. M. Slaughter, A. Taboada,
and G. M. Tolson.

295
 

 

 

 

 

 

 

 

 

 

  

 

 

 

 

l

  
  

596 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [cHAP. 13
TaBLE 13-1
CoMPOSITIONS OF POTENTIAL STRUCTURAL MATERIALS
Quantity in alloy, w/o
Components

Inconel INOR-8 Hastelloy B
Chromium 14-17 6-8 1 (max)
Iron 6-10 5 (max) 4-7
Molybdenum 15-18 26-30
Manganese 1 (max) 0.8 (max) 1.0 (max)
Carbon 0.15 (max) 0.04-0.08 0.05 (max)
Silicon 0.5 0.35 (max) 1.0 (max)
Sulfur 0.01 0.01 (max) 0.03 (max)
Copper 0.5 0.35 (max)
Cobalt 0.2 (max) 2.5 (max)
Nickel 72 (min) Balance Balance

_»g _ | Hot-Leg
Samples
Clam- % o
shell
Heaters fig

 

 

 

1) Cold-Leg

Samples

 

 

 

 

Fig. 13-1. Diagram of a standard thermal-convection loop, showing locations
at which metallographic sections are taken after operation.
13-1] SURVEY OF SUITABLE MATERIALS 597

D
Cold Sectio
\ /
/ §
& /
FI/ Heated Sections 5
1 | 1 | ]°

= h— /Freeze Valve

Heater Connecting Lugs
/ .
\/ Fill Tank

Fic. 13-2. Diagram of forced-circulation loop for corrosion testing.

 

  

 

 

to salts. The material is structurally stable in the operating temperature
range, and the oxidation rate is less than 2 mils in 100,000 hr. No difficulty
is encountered in fabricating standard shapes when the commercial prac-
tices established for nickel-base alloys are used. Tubing, plates, bars,
forgings, and castings of INOR-8 have been made successfully by several
major metal manufacturing companies, and some of these companies are
prepared to supply it on a commercial basis. Welding procedures have
been established, and a good history of reliability of welds exists. The
material has been found to be easily weldable with rod of the same com-
position.

Inconel is, of course, an alternate choice for the primary-circuit struc-
tural material, and much information is available on its compatibility with
molten salts and sodium. Although probably adequate, Inconel does not
have the degrée of flexibility that INOR-8 has in corrosion resistance to
different salt systems, and its lower strength at reactor operating tempera-
tures would require heavier structural components.

A considerable nuclear advantage would exist in a reactor with an
uncanned graphite moderator exposed to the molten salts. Long-time
exposure of graphite to a molten salt results in the salt penetrating the
available pores, but it is probable, with the “impermeable’ types of
098 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS [cHAP. 13

graphite now being developed, that the degree of salt penetration en-
countered can be tolerated. The attack of the graphite by the salt and the
carburization of the metal container seem to be negligible if the temperature
is kept below 1300°F. More tests are needed to finally establish the com-
patibility of graphite-salt-alloy systems.

Finally, a survey has been made of materials suitable for bearings and
valve seats in molten salts. Cermets, ceramics, and refractory metals
appear to be promising for this application and are presently being in-
vestigated.

13-2. CoRrrOSION OF NICKEL-BASE ALLoYS BY MOLTEN SALTS

13-2.1 Apparatus used for corrosion tests. Nickel-base alloys have been
exposed to flowing molten salts in both thermal-convection loops and in
loops containing pumps for forced circulation of the salts. The thermal-
convection loops are designed as shown in Fig. 13-1. When the bottom
and an adjacent side of the loop are heated, usually with clamshell heaters,
convection forces in the contained fluid establish flow rates of up to 8 ft/min,
depending on the temperature difference between the heated and unheated
portions of the loop. The forced-circulation loops are designed as shown
in Fig. 13-2. Heat is applied to the hot leg of this type of loop by direct
resistance heating of the tubing. Large temperature differences (up te
300°F) are obtained by air-cooling of the cold leg. Reynolds numbers of
up to 10,000 are attainable with 1/2-in.-ID tubing, and somewhat higher
values can be obtained with smaller tubing.

13-2.2 Mechanism of corrosion. Most of the data on corrosion have
been obtained with Inconel, and the theory of the corrosive mechanism
was worked out for this alloy. The corrosion of INOR-8 occurs to a lesser
degree but follows a pattern similar to that observed for Inconel and pre-
sumably the same theory applies.

The formation of subsurface voids is initiated by the oxidation of chro-
mium along exposed surfaces through oxidation-reduction reactions with
impurities or constituents of the molten fluoride-salt mixture. As the sur-
face is depleted in chromium, chromium from the interior diffuses down
the concentration gradient toward the surface. Since diffusion occurs by
a vacancy process and in this particular situation is essentially monodirec-
tional, it is possible to build up an excess number of vacancies in the metal.
These precipitate in areas of disregistry, principally at grain boundaries
and impurities, to form voids. These voids tend to agglomerate and grow in
size with increasing time and temperature. Examinations have demon-
strated that the subsurface voids are not interconnected with each other
or with the surface. Voids of the same type have been found in Inconel
13-2] MOLTEN-SALT CORROSION OF NICKEL-BASE ALLOYS 999

after high-temperature oxidation tests and high-temperature vacuum tests
in which chromium was selectively removed.

The selective removal of chromium by a fluoride-salt mixture depends on
various chemical reactions, for example:

1. Impurities in the melt:
FeFs 4+ Cr = CrFa2+ Fe. (13-1)
2. Oxide films on the metal surface:
2Fe-03 4+ 3CrF4 === 3CrO2 + 4FeFs. (13-2)
3. Co‘nstituents of the fuel:
Cr+ 2UF4 == 2UF3 4 CrFo. (13-3)

The ferric fluoride formed by the reaction of Eq. (13-2) dissolves in the
melt and further attacks the chromium by the reaction of Eq. (13-1).

The time-dependence of void formation in Inconel, as observed both in
thermal-convection and forced-circulation systems, indicates that the at-
tack is initially quite rapid but that it then decreases until a straight-line
relationship exists between depth of void formation and time. This effect
can be explained in terms of the corrosion reactions discussed above. The
initial rapid attack found for both types of loops stems from the reaction
of chromium with impurities in the melt [reactions (13-1) and (13-2)] and
with the UF4 constituent of the salt [reaction (13-3)] to establish a quasi-
equilibrium amount of CrF3: in the salt. At this point attack proceeds
linearly with time and occurs by a mass-transfer mechanism which, al-
though it arises from a different cause, is similar to the phenomenon of
temperature-gradient mass transfer observed in liquid metal corrosion.

In molten fluoride-salt systems, the driving force for mass transfer is a
result of a temperature dependence of the equilibrium constant for the
reaction between chromium and UFs (Eq. 13-3). If nickel and iron are
considered inert diluents for chromium in Inconel, the process can be
simply described. Under rapid circulation, a uniform concentration of UF4,
UF3, and CrFs is maintained throughout the fluid; the concentrations must
satisfy the equilibrium constant

' _ YOiF; ° ’Y urs Ncer, - N2ur,
K,=K, Ky . 134
Y Yor - Y2ur, Ner - N2uF, (134)

where N represents the mole fraction and 7 the activity coefficient of the
indicated component.
600 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [cHAP. 13

 

Fia. 13-3. Hot-leg section from an Inconel thermal-convection loop which cir-
culated the fuel mixture NaF-ZrF4+UF4 (50-46-4 mole 97) for 1000 hr at 1500°F.
(250%)

Under these steady-state conditions, there exists a temperature 7', inter-
mediate between the maximum and minimum temperatures of the loop,
at which the initial composition of the structural metal is at equilibrium
with the fused salt. Since Ky increases with increasing temperature, the
chromium concentration in the alloy surface is diminished at temperatures
higher than T and is augmented at temperatures lower than 7. In some
melts, NaF-LiF-KF-UF4, for example, the equilibrium constant of reac-
tion (13-3) changes sufficiently with temperature under extreme tempera-
ture conditions to cause precipitation of pure chromium crystals in the
cold zone. In other melts, for example NaF-ZrF,—UF4, the temperature-
dependence of the corrosion equilibrium is small, and the equilibrium is
satisfied at all useful temperatures without the formation of crystalline
chromium. In the latter systems the rate of chromium removal from the
salt stream at cold-leg regions is dependent on the rate at which chromium
can diffuse into the cold-leg wall. If the chromium concentration gradient
tends to be small, or if the bulk of the cold-leg surface is held at a relatively
low temperature, the corrosion rate in such systems is almost negligible.

It is -obvious that addition of the equilibrium concentrations of UF3
and CrF2 to molten fluorides prior to circulation in Inconel equipment
would minimize the initial removal of chromium from the alloy by reac-
13-2] MOLTEN-SALT CORROSION OF NICKEL-BASE ALLOYS 601

 

Fia. 13-4. Hot-leg section of Inconel thermal-convection loop which circulated
the fuel mixture NaF-ZrF+UF, (55.3-40.7-4 mole 9,) for 1000 hr at 1250°F.
(250 %)

tion (13-3). (It would not, of course, affect the mass-transfer process which
arises as a consequence of the temperature-dependence of this reaction.)
Deliberate additions of these materials have not been practiced in routine
corrosion tests because (1) the effect at the uranium concentrations nor-
mally employed is small, and (2) the experimental and analytical difficul-
ties are considerable. Addition of more than the equilibrium quantity of
UF5 may lead to deposition of some uranium metal in the equipment walls
through the reaction

4UF4s == 3UF4++ U°. (13-5)

For ultimate use in reactor systems, however, it may be possible to treat
the fuel material with calculated quantities of metallic chromium to pro-
vide the proper UF3 and CrF concentrations at startup.

According to the theory described above, there should be no great dif-
ference in the corrosion found in thermal-convection loops and in forced-
circulation loops. The data are in general agreement with this conclusion
so long as the same maximum metal-salt interface temperature is present
in both types of loop. The results of many tests with both types of loop
are summarized in Table 13-2 without distinguishing between the two
types of loop. The maximum bulk temperature of the salt as it left the
heated section of the loop is given. It is known that the actual metal-salt
interface temperature was not greater than 1300°F in the loops with a
602 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [CHAP. 13

TaBLE 13-2

SuMMARY OF CoORROSION DaTA OBTAINED IN THERMAL-CONVECTION AND
ForcED-CircuLATION Loor TEsts oF INCONEL aND INOR-8
ExposED TO VARIOUS CIRCULATING SALT MIXTURES

 

 

Maxi 1t Ti ¢ Depth of subsurface
Constituents of UF4 or ThF, Loop ¢ ax1muri1 S8 1m(z'o void formation at
base salts content material emp eor ature, °p erl?r ton, hottest part of loop,
in.
NaF-ZrF, 1 mole 9, UF4 Inconel 1250 1000 <0.001
1 mole 9, UF, Inconel 1270 6300 0-0.0025
4 mole 9, UF, Inconel 1250 1000 0.002
4 mole 9, UF4 Inconel 1500 1000 0.007-0.010 )
4 mole 9, UF, INOR-8 1500 1000 . 0.002-0.003
0 Inconel 1500 1000 0.002-0.003
NaF-BeF; 1 mole 9, UF, Inconel 1250 1000 0.001
0 Inconel 1500 500 0.004-0.010
3 mole 9, UF, Inconel 1500 500 0.008-0.014
1 mole 9, UF4 . INOR-8 1250 6300 0.001
LiF-BeF; 1 mole 9, UF, Inconel 1250 1000 0.001-0.002
3 mole %, UF4 Inconel 1500 500 0.012-0.020
1 mole 9, UF, INOR-8 1250 1000 0
NaF-LiF-BeF; 0 Inconel 1125 1000 0.002
0 Inconel 1500 500 0.003-0.005
3 mole 9, UF, Inconel 1500 500 0.008-0.013
NaF-LiF-KF 0 Inconel 1125 1000 0.001
2.5 mole 9, UF, Inconel 1500 500 0.017
0 INOR-8 1250 1340 0
2.5 mole 9, UF, INOR-8 1500 1000 0.001-0.003
LiF 29 mole 9, ThF, Inconel 1250 1000 0-0.0015
NaF-BeF, 7 mole 9, ThF, INOR-8 1250 1000 0

 

 

 

 

 

 

 

maximum salt temperature of 1250°F, and was between 1600 and 1650°F
for the loop with a maximum salt temperature of 1500°F,

The data in Table 13-2 are grouped by types of base salt because the
salt has a definite effect on the measured attack of Inconel at 1500°F. The
salts that contain BeF2 are somewhat more corrosive than those containing
ZrF4, and the presence of LiF, except in combination with NaF, seems to
accelerate corrosion.

At the temperature of interest in molten-salt reactors, that is, 1250°F, the
same trend of relative corrosiveness of the different salts may exist for
Inconel, but the low rates of attack observed in tests preclude a conclusive
decision on this point. Similarly, if there is any preferential effect of the
base salts on INOR-8, the small amounts of attack tend to hide it.

As expected from the theory, the corrosion depends sharply on the UF4
concentration. Studies of the nuclear properties of molten-salt power
reactors have indicated (see Chapter 14) that the UF4 content of the fuel
will usually be less than 1 mole %, and therefore the corrosiveness of salts

 
13-2] MOLTEN-SALT CORROSION OF NICKEL-BASE ALLOYS 603

 

Fig. 13-5. Hot-leg section of Inconel thermal-convection loop which circulated
the fuel mixture LiF-BeF3s-UF4 (62-37-1 mole 9,) for 1000 hr at 1250°F. (250X)

with higher UF4 concentrations, such as those described in Table 13-2,
will be avoided.

The extreme effect of temperature is also clearly indicated in Table 13-2.
In general, the corrosion rates are three to six times higher at 1500°F than
at 1250°F. This effect is further emphasized in the photomicrographs
presented in Figs. 13-3 and 13-4, which offer a comparison of metallo-
graphic specimens of Inconel that were exposed to similar salts of the NaF-
7rF~UF4 system at 1500°F and at 1250°F. A metallographic specimen of
Inconel that was exposed at 1250°F to the salt proposed for fueling of the
molten-salt power reactor is shown in Fig. 13-5.

The effect of sodium on the structural materials of interest has also been
extensively studied, since sodium is proposed for use as the intermediate
heat-transfer medium. Corrosion problems inherent in the utilization of
sodium for heat-transfer purposes do not involve so much the deterioration
of the metal surfaces as the tendency for components of the container
material to be transported from hot to cold regions and to form plugs of
deposited material in the cold region. As in the case of the corrosion by the
salt mixture, the mass transfer in sodium-containing systems 1s extremely
dependent on the maximum system operating temperature. The results of
604 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS [cHAP. 13

numerous tests indicate that the nickel-base alloys, such as Inconel and
INOR-S8, are satisfactory containers for sodium at temperatures below
1300°F, and that above 1300°F the austenitic stainless steels are preferable.

13-3. FaBricaTioN orF INOR-8

13-3.1 Casting. Normal melting procedures, such as induction or elec-
tric furnace melting, are suitable for preparing INOR-8. Specialized tech-
niques, such as melting under vacuum or consumable-electrode melting,
have also been used without difficulty. Since the major alloying constitu-
ents do not have high vapor pressures and are relatively inert, melting losses
are negligible, and thus the specified chemical composition can be obtained
through the use of standard melting techniques. Preliminary studies indi-
cate that intricately shaped components can be cast from this material.

13-3.2 Hot forging. The temperature range of forgeability of INOR-8
1s 1800 to 2250°F. This wide range permits operations such as hammer
and press forging with a minimum number of reheats between passes and
substantial reductions without cracking. The production of hollow shells
for the manufacture of tubing has been accomplished by extruding forged
and drilled billets at 2150°F with glass as a lubricant. Successful extru-
sions have been made on commercial presses at extrusion ratios of up to
14:1. Forging recoveries of up to 90% of the ingot weight have been re-
ported by one vendor.

13-3.3 Cold-forming. In the fully annealed condition, the ductility of
the alloy ranges between 40 and 50% elongation for a 2-in. gage length.
Thus, cold-forming operations, such as tube reducing, rolling, and wire
drawing, can be accomplished with normal production schedules. The ef-
fects of cold-forming on the ultimate tensile strength, yield strength, and
elongation are shown in Fig. 13-6.

Forgeability studies have shown that variations in the carbon content
have an effect on the cold-forming of the alloy. Slight variations of other
components, in general, have no significant effects. The solid solubility of
carbon in the alloy is about 0.01%. Carbon present in excess of this amount
precipitates as discrete particles of (Ni,Mo)sC throughout the matrix;
the particles dissolve sparingly even at the high annealing temperature
of 2150°F. Thus cold-working of the alloy causes these particles to align
in the direction of elongation and, if they are present in sufficient quantity,
they form continuous stringers of carbides. The lines of weakness caused
by the stringers are sufficient to propagate longitudinal fractures in tubular
products during fabrication. The upper limit of the carbon content for
tubing is about 0.10%, and for other products it appears to be greater
than 0.20%. The carbon content of the alloy is controllable to about 0.02%
in the range below 0.10%. |
13-3] FABRICATION OF INOR-8 605

(x103)
25o1||[||||1oo

 

    
  
    
 

200 |—
Ultimate Tensile Strength

80

60
0.2% Offset Yield Strength

Strength (psi)

100

Elongation
50

Per Cent Elongation (2-in. Gage Length)

 

 

 

| | l L1 L1 | 0
0 10 20 30 40 50 60 70 80 90

Reduction in thickness (%)

 

Fig. 13-6. Work hardening curves for INOR-8 annealed 1 hr at 2150°F before
reduction.

13-3.4 Welding. The parts of the reactor system are joined by welding,
and therefore the integrity of the system is in large measure dependent
on the reliability of the welds. During the welding of thick sections, the
material will be subjected to a high degree of restraint, and consequently
both the base metal and the weld metal must not be susceptible to cracking,
embrittlement, or other undesirable features.

Extensive tests of weld specimens have been made. The circular-groove
test, which accurately predicted the weldability of conventional materials
with known welding characteristics, was found to give reliable results for
nickel-base alloys. In the circular-groove test, an inert-gas-shielded tung-
sten-arc weld pass is made by fusion welding (i.e., the weld metal contains
no filler metal) in a circular groove machined into a plate of the base metal.
The presence or absence of cracks in the weld metal is then observed.
Test samples of two heats of INOR-8 alloys, together with samples of four
other alloys for comparison, are shown in Fig. 13-7. As may be seen, the
restraint of the weld metal caused complete circumferential cracking in
INOR-8 heat 8284, which contained 0.04%, B, whereas there are no cracks
in INOR-8 heat 30-38, which differed from heat 8284 primarily in the
606 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [CHAP. 13

NOR 8 HEAT 30-38

 

STAINLESS STEEL

F1a. 13-7. Circular-groove tests of weld metal cracking.

WELD DEPOSIT

 

7/ T~

 

 

 

 

 

 

 

 

 

 

 

s 12 in
o
// N
777 .
’ /o in.
117 2!
[ I/\4 .
BACKING STRIP —~ /o-in-THICK PLATE

F1a. 13-8. Weld test plate design showing method of obtaining specimen.
13-3] FABRICATION OF INOR-8 607

 

Fic. 13-9. Weld in slot of vacuum-melted ingot.

absence of boron. Two other INOR-8 heats that did not contain boron
similarly did not crack when subjected to the circular-groove test.

In order to further study the effect of boron in INOR-8 heats, several
3-b vacuum-melted ingots with nominal boron contents of up to 0.10%
were prepared, slotted, and welded as shown in Fig. 13-9. All ingots with
0.029, or more boron cracked in this test.

A procedure specification for the welding of INOR-8 tubing is available
that is based on the results of these cracking tests and examinations of
numerous successful welds. The integrity of a joint, which is a measure
of the quality of a weld, is determined through visual, radiographic, and
metallographic examinations and mechanical tests at room and service
temperatures. It has been established through such examinations and
tests that sound joints can be made in INOR-8 tubing that contains less
than 0.029, boron.

Weld test plates of the type shown in Fig. 13-8 have also been used for
studying the mechanical properties of welded joints. Such test plates were
side-bend tested in the apparatus illustrated in Fig. 13-10. The results of
the tests, presented in Table 13-3, indicate excellent weld metal ductility.
For example, the ductility of heat M—5 material is greater than 409, at
temperatures up to and including 1500°F.
608 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [CHAP. 13

Hydraulic Cylinder

Specimen
Thermocouple

       
  

—~—— Bend

Quartz Rod e Specimen

 

 

Control
Thermocouple

  

¥ a -
e e ne Ay el ¥ dnfonv e
: T S P > TR

 

 

    
 

° ) }:;A

1\74

  
 

Heating Element—
Connections
(220v or 110v)

£ -

  

Quartz Mirror

10123435
Scale in Inches

F1a. 13-10. Apparatus for bend tests at high temperatures.

13-3.5 Brazing. Welded and back-brazed tube-to-tube sheet joints are
normally used in the fabrication of heat exchangers for molten salt service.
The back-brazing operation serves to remove the notch inherent in con-
ventional tube-to-tube sheet joints, and the braze material minimizes the
possibility of leakage through a weld failure that might be created by ther-
mal stresses in service.

The nickel-base brazing alloys listed in Table 13—4 have been shown to
be satisfactory in contact with the salt mixture LiF-KF-NaF-UF. in
tests conducted at 1500°F for 100 hr. Further, two precious metal-base
brazing alloys, 829, Au-189, Ni and 809, Au-209, Cu, were unattacked
in the LiF-KF-NaF-UF4 salt after 2000 hr at 1200°F. These two precious
609

FABRICATION OF INOR-8

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610 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS

TABLE 134

NickEL-BASE BraziNnGg ArrLoys ror USE IN

HeAaT EXCHANGER FABRICATION

 

 

 

 

 

 

 

 

 

Brazing alloy content, w/o
Components
Alloy 52 Alloy 91 Alloy 93
Nickel 91.2 91.3 93.3
Silicon 4.5 4.5 3.5
Boron 2.9 2.9 1.9
Iron and carbon Balance Balance Balance

 

 

[cHAP. 13

metal alloys were also tested in the LiF-BeFs—UF4 mixture and again were
not attacked.

13-3.6 Nondestructive testing. An ultrasonic inspection technique is
available for the detection of flaws in plate, piping, and tubing. The water-
immersed pulse-echo ultrasound equipment has been adapted to high-
speed use. KEddy-current, dye-penetrant, and radiographic inspection
methods are also used as required. The inspected materials have included
Inconel, austenitic stainless steel, INOR-8, and the Hastelloy and other
nickel-molybdenum-base alloys.

Methods are being developed for the nondestructive testing of weld-
ments during initial construction and after replacement by remote means
in a high-intensity radiation field, such as that which will be present if
malntenance work is required after operation of a molten-salt reactor.
The ultrasonic technique appears to be best suited to semiautomatic and
remote operation and of any of the applicable methods, it will probably
be the least affected by radiation. Studies have indicated that the diffi-
culties encountered due to the high ultrasonic attenuation of the weld
structures in the ultrasonic inspection of Inconel welds and welds of some
of the austenitic stainless steels are not present in the inspection of INOR-8
welds. In addition, the troublesome large variations in ultrasonic attenua-
tion common to Inconel and austenitic stainless steel welds are less severe
in INOR-8 welds. The mechanical equipment designed for the remote
welding operation will be useful for the inspection operation.

In the routine inspection of reactor-grade construction materials, a tube,
pipe, plate, or rod is rejected if a void is detected that is larger than 59
of the thickness of the part being inspected. In the inspection of a weld,
the integrity of the weld must be better than 959, of that of the base metal.
13-4] MECHANICAL AND THERMAL PROPERTIES OF INOR-S8 611

Typical rejection rates for Inconel and INOR-8 are given below:

 

 

 

 

Rejection rate (9))
Item
Inconel INOR-8

Tubing 17 20
Pipe 12 14
Plate 8 8
Rod 5 5
Welds 14 14

 

 

 

 

 

The rejection rates for INOR-8 are expected to decline as more experience
is gained in fabrication.

13—4. M ECHANICAL AND THERMAL ProrPeErRTIES OF INOR-8

13—4.1 Elasticity. A typical stress-strain curve for INOR-8 at 1200°F
is shown in Fig. 13-11. Data from similar curves obtained from tests at
room temperature up to 1400°F are summarized in Fig. 13-12 to show
changes in tensile strength, yield strength, and ductility as a function of
temperature. The temperature dependence of the Young’s modulus of this
material is illustrated in Fig. 13-13.

13—4.2 Plasticity. A series of relaxation tests of INOR-8 at 1200 and
1300°F has indicated that creep will be an important design consideration
for reactors operating in this temperature range. The rate at which the
stress must be relaxed in order to maintain a constant elastic strain at
1300°F is shown in Fig. 13-14, and similar data.for 1200°F are presented
in Fig. 13-15. The time lapse before the material becomes plastic 1s about
1 hr at 1300°F and about 10 hr at 1200°F. The time period during which
the material behaves elastically becomes much longer at lower tempera-
tures, and below some temperature, as yet undetermined, the metal will
continue to behave elastically indefinitely.

It is possible to summarize the creep data by comparing the times to
1.09, total strain, as a function of stress, in the data shown in Fig. 13-16.
The reproducibility of creep data for this material is indicated by the
separate curves shown in Fig. 13-17. It may be seen that quite good corre-
lation between the creep curves is obtained at the lower stress values.
Some scatter in time to rupture occurs at 25,000 psi, a stress which corre-
sponds to the 0.29, offset yield strength at this temperature. Such scatter
is to be expected at this high stress level.
612 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS

 

35,000 | 7
30,000 /
25,000 |—

20,000 —

Stress (psi)

15,000 —

10,000 —

5,000 —

 

o | |

/

1

| | |

 

 

0.0006 0.0012 0.0018

0.0024 0.0030 0.0036

Elongation (in./in.)

Fia. 13-11. Stress-strain relationships for INOR-8 at 1200°F. Initial slope (rep-
resented by dashed line at left) is equivalent to a static modulus of elasticity in
tension of 25,200,000 psi. The dashed line at right is the curve for plastic deforma-
tion of 0.002 in/in; its intersection with the stress-strain curve indicates a yield
strength of 25,800 psi for 0.29, offset. Ultimate tensile strength, 73,895 psi; gage
length, 3.25 in.; material used was from heat 3038.

[cHAP. 13

The tensile strengths of several metals are compared with the tensile
strength of INOR-8 at 1300°F in the following tabulation, and the creep
properties of the several alloys at 1.09, strain are compared in Fig. 13—18.

 

Material

Tensile strength at
1300°F, psi

 

 

18-8 stainless steel
Cr-Mo steel (69, Cr)
Hastelloy B
Hastelloy C

Inconel

INOR-8

 

 

40,000
20,000
70,000

100,000
60,000
65,000

 

 

The test results indicate that the elastic and plastic strengths of INOR-8
are near the top of the range of strength properties of the several alloys
13-4] MECHANICAL AND THERMAL PROPERTIES OF INOR-8 613

 

     
  
   

 

 

 

 

(x103)
110 e
| I I [ | I | | -
Tensile Strength
100 — 3
90 — &
80 — 8
g 70 —R .
o Elongation, % e
@ £
g 5
A 60 g &
50 —H g
40 — <
Yield Strength
30 }— — 9
20 | | | I | | ] | &

N
0 2 4 6 8 10 12 14 16 18x102)

Temperature, °F

F1gc. 13-12. Tensile properties of INOR-8 as a function of temperature.

(x109)
34

 

30

28 |—

26 —

24 |-

Young’s Modulus (psi)

 

22 —

ol L 1 1 1 1
0 200 400 600 800 1000 1200 1400

 

 

 

Temperature (°F)

F1a. 13-13. Young’s modulus for INOR-8.
614 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS [cuAP. 13

(x103)
28 T T TTTTT T TTTTTT T T 171111

0.2% Constant Strain

 

16 — 0.1% Constant Strain —

0.05% Constant Strain

Stress (psi)

 

 

Lt e EER Lt
0.1 0.2 0.5 1.0 2 5 10 20 50 100
Time (hr)

 

Fic. 13-14. Relaxation of INOR-8 at 1300°F at various constant strains.

(x103)
28 l T T T 11711 l T T T 171711 l T T 1111

 

24 — 0.1% Constant Strain

 

20 —

o
[

  
 

0.05% Constant Strain

Stress (psi)

~
I
I

* Discontinued Test

 

 

0 | L | I EEEE | L1t
1 2 5 10 20 50 100 200 500 1000
Time (hr)

 

Fic. 13-15. Relaxation of INOR-8 at 1200°F at various constant strains.
13-4] MECHANICAL AND THERMAL PROPERTIES OF INOR-8 615

 

   
 
     

 

 

 

 

 

 

 
  
    
   

 

 

 

109 -
orreir o rreer v P eyt v b T ri=g
5 /1100°F =
2 T TN
]04 — ~ = —
— ~ =
-~ . FE 1300°F/ ~o =
‘% 5 — ~ ~ -
§ e ——
s 2 —
W
Extrapolated ~
]03 — :~
5 |— =
| 1yr 10 yr
L | "
102 L e b b Preree L Eriime b el
1 10 100 1000 10,000 100,000
Time (hr) to 1% Strain
Fic. 13-16. Creep data for INOR-S8.
100 = T T T TTTTT T T T 1T -
- ¥* —]
50 - —
— Results for Three Samples -
Stressed at 25,000 psi
20— —
2
L 1.0 —
o — —
& |- _]
0.5 [ —
B Results for Two Samples N
- Stressed at 20,000 psi —
0.2 —
o A 11 L L L Ll
10 20 50 100 200 500 1000 2000

Time (hr)

F1g. 13-17. Creep-rupture data for INOR-S8.
616 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [cHAP. 13

 

R b DT I EEE

—
—
—
ed
—
—

Stress (psi)

 

IRERERN
L 1i1ll]

p—

()
I

 

 

2 L1 it [ L 111l L 1 111

10 20 50 100 200 500 1000 2000 5000 10,000
Time to 1% Strain at 1300°F (hr)

 

F1c. 13-18. Comparison of the creep properties of several alloys.

commonly considered for high-temperature use. Since INOR-8 was de-
signed to avoid the defects inherent in these other metals, it 1s apparent
that the undesirable aspects have been eliminated without any serious
loss in strength.

13—4.3 Aging characteristics. Numerous secondary phases that are ca-
pable of embrittling a nickel-base alloy can exist in the Ni-Mo-Cr-Fe-C
system, but no brittle phase exists if the alloy contains less than 209, Mo,
89, Cr, and 59, Fe. INOR-8, which contains only 15 to 189, Mo, consists
principally of two phases: the nickel-rich solid solution and a complex car-
bide with the approximate composition (Ni, Mo)¢C. Studies of the effect
of the carbides on creep strength have shown that the highest strength
exists when a continuous network of carbides surrounds the grains. Tests
have shown that carbide precipitation does not cause significant embrittle-
ment at temperatures up to 1480°F. Aging for 500 hr at various tempera-
tures, as shown in Fig. 13—-19, improves the tensile properties of the alloy.
The tensile properties at room temperature, as shown in Table 13-5, are
virtually unaffected by aging.
INOR-8

 

 

 

 

 

 

 

     
 

 

 

 

13-4] MECHANICAL AND THERMAL PROPERTIES OF
(x103)
110
| ! |
100 |— -
= 90
&
= 80
@
g 70
A
2 60
§ Annealed 1 hr at 2100°F
=0 - Aged 500 hr at Test Temperature o
40 — —
0.2% Yield Point
30 w————__—_——_-fi
60
—_— I |
50 —
2
< 40 Elongati —
® 30—
o
c
2 20 - —
Wl
10 — —
0 l . |
1000 1100 1200 1300 1400

Test Temperature (°F)

617

F1c. 13-19. Effect of aging on high-temperature tensile properties of INOR-8.

 

 

 

TABLE 13-5
REesuLts oF RooM-TEMPERATURE EMBRITTLEMENT TESTS oF INOR-8
Ultimate tensile | Yield point at | Elongation,
Heat treatment strength, psi 0.29, offset, psi %
Annealed* 114,400 44,700 50
Annealed and aged 500 hr at
1000°F 112,000 42 500 53
Annealed and aged 500 hr at
1100°F 112,600 44,000 51
Annealed and aged 500 hr at
1200°F 112,300 44,700 51
Annealed and aged 500 hr at
1300°F 112,000 44 500 49
Annealed and aged 500 hr at
1400°F 112,400 43,900 50

 

 

 

 

 

 

*0.045-in. sheet, annealed 1 hr at 2100°F and tested at a strain rate of 0.05 in/min.
6018 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [cHAP. 13

13—4.4 Thermal conductivity and coefficient of linear thermal expansion.
Values of the thermal conductivity and coefficient of linear thermal ex-
pansion are given in Tables 13—6 and 13-7.

TABLE 13-6

CoMPARISON OF THERMAL Conpuctivity VALUES ForR INOR-8
AND INCONEL AT SEVERAL TEMPERATURES

 

 

 

 

 

 

 

 

 

Thermal conductivity, Btu/(ft2)(sec) (°F/ft)
Temperature,
°F
INOR-S8 Inconel
212 5.56 9.44
392 6.77 9.92
572 " 11.16 10.40
752 12.10 10.89
933 14.27 11.61
1112 16.21 12.10
1292 18.15 12.58
TaBLE 13-7

CoErFICIENT OF LINEAR ExpansioN or INOR-8
FOR SEVERAL TEMPERATURE RANGES

 

 

 

Temperature range, °F Coeffimenii;n(;f(ilrllx;??;‘ )expansmn,
X106
70-400 5 76
70-600 6.93
70-800 6. 58
70-1000 6.89
70-1600 8 10
70-1800 8 39

 

 

 

 
13-5] OXIDATION RESISTANCE 619

13-5. OXIDATION RESISTANCE

The oxidation resistance of nickel-molybdenum alloys depends on the
service temperature, the temperature cycle, the molybdenum content, and
the chromium content. The oxidation rate of the binary nickel-molybdenum
alloy passes through a maximum for the alloy containing 159, Mo, and the
scale formed by the oxidation is NiMoO4 and NiO. Upon thermal cycling
from above 1400°F to below 660°F, the NiMoO4 undergoes a phase trans-
formation which causes the protective scale on the oxidized metal to spall.
Subsequent temperature cycles then result in an accelerated oxidation
rate. Similarly, the oxidation rate of nickel-molybdenum alloys containing
chromium passes through a maximum for alloys containing between 2 and
69, Cr. Alloys containing more than 69, Cr are insensitive to thermal
cycling and the molybdenum content because the oxide scale is pre-
dominantly stable CroOs. An abrupt decrease, by a factor of about 40, in
the oxidation rate at 1800°F is observed when the chromium content is
increased from 5.9 to 6.29.

The oxidation resistance of INOR-8 is excellent, and continuous opera-
tion at temperatures up to 1800°F is feasible. Intermittent use at tempera-
tures as high as 1900°F could be tolerated. For temperatures up to 1200°F,
the oxidation rate is not measurable; it is essentially nonexistent after
1000 hr of exposure in static air. It is estimated that oxidation of 0.001 to
0.002 in. would occur in 100,000 hr of operation at 1200°F. The effect of
temperature on the oxidation rate of the alloy is shown in Table 13-8.

TABLE 13-8

OxIDATION RATE oF INOR-8 AT VarioUs TEMPERATURES*

 

 

 

 

 

 

 

 

. : o
Test temperature, Weight gain, mg/cm Shape of
°F rate curve
In 100 hr In 1000 hr

1200 0.00 0.00 Cubic or logarithmic

1600 0.25 0.67¢ Cubic

1800 0.48 1.5% Parabolic

1900 0.52 2.07 Parabolic

2000 2.70 28.27 Linear

 

*3.7 mg/cm2 = 0.001 in. of oxidation.

tExtrapolated from data obtained after 170 hr at temperature.

 
620 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS [cHAP. 13

 

Fia. 13-20. Components of a duplex heat exchanger fabricated of Inconel clad
with type-316 stainless steel.

13-6. FABRICATION OF A DUPLEX TuBING HEAT EXCHANGER

The compatibility of INOR-8 and sodium is adequate in the temperature
range presently contemplated for molten-salt reactor heat-exchanger oper-
ation. At higher temperatures, mass transfer could become a problem, and
therefore the fabrication of duplex tubing has been investigated. Satis-
factory duplex tubing has been made that consists of Inconel clad with
type—316 stainless steel, and components for a duplex heat exchanger have
been fabricated, as shown in Fig. 13-20.

The fabrication of duplex tubing is accomplished by coextrusion of
billets of the two alloys. The high temperature and pressure used result
in the formation of a metallurgical bond between the two alloys. In sub-
sequent reduction steps the bonded composite behaves as one material.
The ratios of the alloys that comprise the composite are controllable to
within 39,,. The uniformity and bond integrity obtained in this process are
illustrated in Fig. 13-21.

The problem of welding INOR-8-stainless steel duplex tubing is being
studied. Experiments have indicated that proper selection of alloy ratios
and weld design will assure welds that will be satisfactory in high-tempera-
ture service.

To determine whether interdiffusion of the alloys would result in a con-
tinuous brittle layer at the interface, tests were made in the temperature
range 1300 to 1800°F. As expected, a new phase appeared at the interface
between INOR—-8 and the stainless steel which increased in depth along
the grain boundaries with increases in the temperature. The interface of a
duplex sheet held at 1300°F for 500 hr is shown in Fig. 13-22. Tests of
this sheet showed an ultimate tensile strength of 94,400 psi, a 0.29, offset
yield strength of 36,800 psi, and an elongation of 519,. Creep tests of the
13-6] FABRICATION OF A DUPLEX TUBING HEAT EXCHANGER 621

 

  

Ly R

 

F1c. 13-21. Duplex tubing consisting of Inconel over type-316 stainless steel.
Etchant: glyceria regia.

sheet showed that the diffusion resulted in an increase in the creep re-
sistance with no significant loss of ductility.

Thus no major difficulties would be expected in the construction of an
INOR-8-stainless steel heat exchanger. The construction experience thus
far has involved only the 20-tube heat exchanger shown in Fig. 13-20.
622 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS [cEAP. 13

Unstressed

Stainless
Steel

 

Fre. 13-22. Unstressed and stressed specimens of INOR-8 clad with type-316
stainless steel after 500 hr at 1300°F. Etchant: electrolytic H2SO4 (29 solution).
(100X%)

 

 

A o1 NN g

 

F1a. 13-23. CCN graphite (A) before and (B) after exposure for 1000 hr to NaF-
ZrF4+-UF4 (50-46-4 mole 9) at 1300°F as an insert in the hot leg of a thermal-
convection loop. Nominal bulk density of graphite specimen: 1.9 g/cm3.
13-8] GRAPHITE WITH MOLTEN SALTS AND NICKEL-BASE ALLOYS 623

13—-7. AvaiLaBiLiTy oF INOR-8

Two production heats of INOR-8 of 10,000 1b each and numerous smaller
heats of up to 5000 Ib have been melted and fabricated into various shapes
by normal production methods. Evaluation of these commercial products
has shown them to have properties similar to those of the laboratory heats
prepared for material selection. Purchase orders are filled by the vendors
in one to six months, and the costs range from $2.00 per pound in ingot
form to $10.00 per pound for cold-drawn welding wire. The costs of tubing,
plate, and bar products depend to a large extent on the specifications of the
finished products.

13-8. CoMPATIBILITY OF GRAPHITE WITH MOLTEN SALTS AND
NickeL-BAsE ALLOYS

If graphite could be used as a moderator in direct contact with a molten
salt, it would make possible a molten-salt reactor with a breeding ratio in
excess of one (see Chapter 14). Problems that might restrict the usefulness
of this approach are possible reactions of graphite and the fuel salt, pene-
tration of the pores of the graphite by the fuel, and carburization of the
nickel-alloy container.

Many molten fluoride salts have been melted and handled in graphite
crucibles, and in these short-term uses the graphite is inert to the salt.
Tests at temperatures up to 1800°F with the ternary salt mixture
NaF-ZrF4+UF4 gave no indication of the decomposition of the fluoride
and no gas evolution so long as the graphite was free from a silicon im-
purity.

Longer-time tests of graphite immersed in fluoride salts have shown
greater indications of penetration of the graphite by salts, and it must be
assumed that the salt will eventually penetrate the available pores in the
graphite. The “impermeable’’ grades of graphite available experimentally
show greatly reduced penetration, and a sample of high-density, bonded,
natural graphite (Degussa) showed very little penetration. Although
quantitative figures are not available, it is likely that the extent of pene-
tration of “‘impermeable’ graphite grades can be tolerated.

Although these penetration tests showed no visible effects of attack of
the graphite by the salt, analyses of the salt for carbon showed that at
1500°F more than 1% carbon may be picked up in 100 hr. The carbon
pickup appears to be sensitive to temperature, however, inasmuch as
only 0.025% carbon was found in the salt after a 1000-hr exposure at
1300°F.

In some instances coatings have been found on the graphite after ex-
posure to the salt in Inconel containers, as illustrated in Fig. 13-23. A
cross section through the coating is shown in Fig. 13-24. The coating was
624 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [CHAP. 13

 

F1c. 13-24. Cross sections of samples shown in Fig. 13-23. (A) Before exposure;
(B) after exposure. Note the thin film of Cr3sCz on the surface in (B). The black
areas in (A) are pores. In (B) the pores are filled with salt. (100X)

found to be nearly pure chromium carbide, CrsCs. The source of the
chromium was the Inconel container. -

In the tests run thus far, no positive indication has been found of car-
burization of the nickel-alloy containers exposed to molten salts and
graphite at the temperatures at present contemplated for power reactors
(<1300°F). The carburization effect seems to be quite temperature sensi-
tive, however, since tests at 1500°F showed carburization of Hastelloy B
to a depth of 0.003 in. in 500 hr of exposure to NaF-ZrF+UF4 containing
graphite. A test of Inconel and graphite in a thermal-convection loop in
which the maximum bulk temperature of the fluoride salt was 1500°F gave
a maximum carburization depth of 0.05 in. in 500 hr. In this case, however,
the temperature of the metal-salt interface where the carburization oc-
curred was considerably higher than 1500°F, probably about 1650°F.

A mixture of sodium and graphite is known to be a good carburizing
agent, and tests with it have confirmed the large effect of temperature on
the carburization of both Inconel and INOR-8, as shown in Table 13-9.
13-10] SUMMARY OF MATERIAL PROBLEMS 625

TABLE 13-9

ErrFEcT OF TEMPERATURE ON CARBURIZATION OF
INcoNEL AND INOR-8 18y 100 HR

 

 

 

Alloy Temperature, °F Depth of carburization, in.
Inconel 1500 0.009
1200 0
INOR-8 1500 0.010
1200 0

 

 

 

 

 

Many additional tests are being performed with a variety of molten
fluoride salts to measure both penetration of the graphite and carburization
of INOR-8. The effects of carburization on the mechanical properties will
be determined.

13-9. MATERIALS FOR VALVE SEATS AND BEARING SURFACES

Nearly all metals, alloys, and hard-facing materials tend to undergo
solid-phase bonding when held together under pressure in molten fluoride
salts at temperatures above 1000°F. Such bonding tends to make the
startup of hydrodynamic bearings difficult or impossible, and it reduces
the chance of opening a valve that has been closed for any length of time.
Screening tests in a search for nonbonding materials that will stand up
under the molten salt environment have indicated that the most promising
materials are TiC-Ni and WC—-Co types of cermets with nickel or cobalt
contents of less than 35 w/o, tungsten, and molybdenum. The tests, in
general, have been of less than 1000-hr duration, so the useful lives of these
materials have not yet been determined.

13-10. SUMMARY OF MATERIAL PROBLEMS

Although much experimental work remains to be done before the con-
struction of a complete power reactor system can begin, it is apparent that
considerable progress has been achieved in solving the material problems
of the reactor core. A strong, stable, and corrosion-resistant alloy with
good welding and forming characteristics is available. Production tech-
niques have been developed, and the alloy has been produced in com-
mercial quantities by several alloy vendors. Finally, it appears that even
at the peak operating temperature, no serious effect on the alloy occurs
when the molten salt it contains is in direct contact with graphite.
CHAPTER 14
NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS*

The ability of certain molten salts to dissolve uranium and thorium. salts
in quantities of reactor interest made possible the consideration of fluid-
fueled reactors with thorium in the fuel, without the danger of nuclear ac-
cidents as a result of the settling of a slurry. This additional degree .of
freedom has been exploited in the study of molten-salt reactors.

Mixtures of the fluorides of alkali metals and zirconium or beryllium, as
discussed in Chapter 12, possess the most desirable combination of low
neutron absorption, high solubility of uranium and thorium compounds,
chemical inertness at high temperatures, and thermal and radiation sta-
bility. The following comparison of the capture cross sections of the alkali
metals reveals that Li? containing 0.019], Li® has a cross section at 0.0795 ev
and 1150°F that is a factor of 4 lower than that of sodium, which also has
a relatively low cross section:

Element Cross section, barns
Li? (containing 0.019%, Li®) 0.073
Sodium 0.290
Potassium 1.13
Rubidium 0.401
Cesium 29

The capture cross section of beryllium is also satisfactorily low at all
neutron energies, and therefore mixtures of LiF and BeF2, which have
satisfactory melting points, viscosities, and solubilities for UF4 and ThF 4,
were selected for investigation in the reactor physics study.

Mixtures of NaF, ZrF4, and UF4 were studied previously, and such a
fuel was successfully used in the Aircraft Reactor Experiment (see Chap-
ters 12 and 16). Inconel was shown to be reasonably resistant to corrosion
by this mixture at- 1500°F, and there is reason to expect that Inconel
equipment would have a life of at least several years at 1200°F. As a fuel
for a central-station power reactor, however, the NaF-ZrF4 system has
several serious disadvantages. The sodium capture cross section is less
favorable than that of Li17?. More important, recent data [1] indicate that
the capture cross section of zirconium is quite high in the epithermal and
intermediate neutron energy ranges. In comparison with the LiF-BeF2
system, the NaF-ZrF4 system has inferior heat-transfer characteristics.

*By L. G. Alexander.
626
NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS 027

Finally, the INOR alloys (see Chapter 13) show promise of being as resistant
to the beryllium salts as to the zirconium salts, and therefore there is no
compelling reason for selecting the NaF-ZrF4 system.

Reactor calculations were performed by means of the Univac* program
Ocusol [2], a modification of the Eyewash program [3], and the Oraclef
program Sorghum. Ocusol is a 31-group, multiregion, spherically symmet-
ric, age-diffusion code. The group-averaged cross sections for the various
elements of interest that were used were based on the latest available data
[4]. Where data were lacking, reasonable interpolations based on resonance
theory were made. The estimated cross sections were made to agree with
measured resonance integrals where available. Saturation and Doppler
broadening of the resonances in thorium as a function of concentration
were estimated. Inelastic scattering in thorium and fluorine was taken
into account crudely by adjusting the value of £o;; however, the Ocusol
code does not provide for group skipping or anisotropy of scattering.

Sorghum is a 31-group, two-region, zero-dimensional, burnout code.
The group-diffusion equations were integrated over the core to remove
the spatial dependency. The spectrum was computed, in terms of a
space-averaged group flux, from group scattering and leakage parameters
taken from an Ocusol calculation. A critical calculation requires about
1 min on the Oracle; changes in concentration of 14 elements during a
specified time can then be computed in about 1 sec. The major assumption
involved is that the group scattering and leakage probabilities do not
change appreciably with changes in core composition as burnup progresses.
This assumption has been verified to a satisfactory degree of approximation.

The molten salts may be used as homogeneous moderators or simply as
fuel carriers in heterogeneous reactors. Although, as discussed below,
graphite-moderated heterogeneous reactors have certain potential advan-
tages, their technical feasibility depends upon the compatibility of fuel,
graphite, and metal, which has not as yet been established. For this rea-
son, the homogeneous reactors, although inferior in nuclear performance,
have been given greatest attention.

A preliminary study indicated that if the integrity of the core vessel
could be guaranteed, the nuclear economy of two-region reactors would
probably be superior to that of bare and reflected one-region reactors. The
two-region reactors were, accordingly, studied in detail. Although entrance
and exit conditions dictate other than a spherical shape, it was necessary,
for the calculations, to use a model comprising the following concentric

*Universal Automatic Computer at New York University, Institute of Mathe-
matics.

1Oak Ridge Automatic Computer and Logical Engine at Oak Ridge National
Laboratory.
628 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

spherical regions: (1) the core, (2) an INOR-8 core vessel 1/3 in. thick,
(3) a blanket approximately 2 ft thick, and (4) an INOR-8 reactor vessel
2/3 in. thick. The diameter of the core and the concentration of thorium
in the core were selected as independent variables. The primary dependent
variables were the critical concentration of the fuel (U235 U233 or Pu239),
and the distribution of the neutron absorptions among the various atomic
species in the reactor. From these, the critical mass, critical inventory,
regeneration ratio, burnup rate, etc. can be readily calculated, as described
in the following section.

14-1. HoMmoGENEOUS REACTORS FUELED wiTe U235

While the isotope U233 would be a superior fuel in molten fluoride-salt
reactors (see Section 14-2), it is unfortunately not available in quantity.
Any realistic appraisal of the immediate capabilities of these reactors must
be based on the use of U235,

The study of homogeneous reactors was divided into two phases: (1) the
mapping of the nuclear characteristics of the initial (i.e., “clean’) states
as a function of core diameter and thorium concentration, and (2) the
analysis of the subsequent performance of selected initial states with
various processing schemes and rates. The detailed results of these studies
are given in the following paragraphs. Briefly, it was found that regenera-
tion ratios of up to 0.65 can be obtained with moderate investment in U235
(less than 1000 kg) and that, if the fission products are removed (Article
14-1.2) at a rate such that the equilibrium inventory is equal to one year’s
production, the regeneration ratio can be maintained above 0.5 for at
least 20 years.

14-1.1 Initial states. A complete parametric study of molten fluoride-
salt reactors having diameters in the range of 4 to 10 ft and thorium con-
centrations in the fuel ranging from O to 1 mole 9, ThF4 was performed.
In these reactors, the basic fuel salt (fuel salt No. 1) was a mixture of
31 mole 9, BeF2 and 69 mole 9, LiF, which has a density of about 2.0 g/cc
at 1150°F. The core vessel was composed of INOR-8. The blanket fluid
(blanket salt No. 1) was a mixture of 25 mole 9, ThF4 and 75 mole 9}, LiF,
which has a density of about 4.3 g/cc at 1150°F. In order to shorten the
calculations in this series, the reactor vessel was neglected, since the re-
sultant error was small. These reactors contained no fission products or
nonfissionable isotopes of uranium other than U238,

A summary of the results is presented in Table 14-1, in which the neutron
balance is presented in terms of neutrons absorbed in a given element per
neutron absorbed in U235 (both by fission and the n—y reaction). The
sum of the absorptions is therefore equal to 7, the number of neutrons
produced by fission per neutron absorbed in fuel. Further, the sum of the
14-1] HOMOGENEOUS REACTORS FUELED WITH U235 629

(x1019)
40 l

X\ l
30 |— ‘ \: mole % ThF, In
.\ vel Salt

20

 

-n

 

1]

o

_—
"

P
°
4
O
:I

U233 Concentration atoms/cm3

e Calculated Y
Values

/

2 — Interpolations
of Data

/

No ThF4 In Fuel Salt

1 I l 1l
2 4 6 8 10
Core Diameter, ft

)

 

 

 

Fig. 14-1. Initial critical concentration of U235 in two-region, homogeneous,
molten fluoride-salt reactors.

absorptions in U238 and thorium in the fuel, and in thorium in the blanket
salt gives directly the regeneration ratio. The losses to other elements are
penalties imposed on the regeneration ratio by these poisons; 1.e., if the core
vessel could be constructed of some material with a negligible cross sec-
tion, the regeneration ratio could be increased by the amount listed for
capture in the core vessel.

The inventories in these reactors depend in part on the volume of the
fuel in the pipes, pumps, and heat exchangers in the external portion of
the fuel circuit. The inventories listed in Table 14-1 are for systems having
a volume of 339 ft3 external to the core, which corresponds approximately
to a power level of 600 Mw of heat. In these calculations it was assumed
that the heat was transferred to an intermediate coolant composed of the
fluorides of Li, Be, and Na before being transferred to sodium metal. In
more recent designs (see Chapter 17), this intermediate salt loop has been
replaced by a sodium loop, and the external volumes are somewhat less
because of the improved equipment design and layout.

Critical concentration, mass, inventory, and regeneration ratio. The data
in Table 14—1 are more easily comprehended in the form of graphs, such as
Fig. 14-1, which presents the critical concentration in these reactors as a
function of core diameter and thorium concentration in the fuel salt. The
data points represent calculated values, and the lines are reasonable
interpolations. The maximum concentration calculated, about 35 X 10'?
[cHAP. 14

NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS

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14-1] HOMOGENEOUS REACTORS FUELED WITH U235 633

 

0T T1T 1 | l

mole % ThF4 In Fuel Salt

@
1
400 |— e Calculated values /—
— Interpoly °

a Data

o~ \

=

6 300 | —
-e @

- / 075 ¢
" .—\—_ /__-—

o @

= v

T 200 |— ly ]
£ o 0.50 ¢
O ~~

o\\

 

 

 

I\ 0.25
100 . )
N ¢
) ~———No ThF4 In Fuel Salt
ol | | |
04 5 6 7 8 9 10

Core Diameter, ft

Fic. 14-2. Initial critical masses of U235 in two-region, homogeneous, molten
fluoride-salt reactors.

atoms of U235 per cubic centimeter of fuel salt, or about 1 mole 9, UFy, is
an order of magnitude smaller than the maximum permissible concentra-
tion (about 10 mole 9).

The corresponding critical masses are graphed in Fig. 14-2. As may be
seen, the critical mass is a rather complex function of the diameter and the
thorium concentration. The calculated points are shown here also, and the
solid lines represent, it is felt, reliable interpolations. The dashed lines
were drawn where insufficient numbers of points were calculated to define
the curves precisely; however, they are thought to be qualitatively correct.
Since reactors having diameters less than 6 ft are not economically attrac-
tive, only one case with a 4-ft-diameter core was computed.

The critical masses obtained in this study ranged from 40 to 400 kg of
U235, However, the critical inventory in the entire fuel cirecuit is of more
interest to the reactor designer than is the critical mass. The critical in-
ventories corresponding to an external fuel volume of 339 ft3 are therefore

shown in Fig. 14-3. Inventories for other external volumes may be com-
puted from the relation

I=M<1+§DIZ;>’

where D is the core diameter in feet, M is the critical mass taken from
Fig. 14-2, V., is the volume of the external system in cubic feet, and I is
the inventory in kilograms of U235, The inventories plotted in Fig. 14-3
634

2000

NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS

 

1000

b

l

Critical Inventory, kg of U235
W
S
|

 

   

l !

mole % ThFy4 In Fuel Salt

No ThF4 In Fuel Salt

 

I T

 

6 8 10

Core Diameter, ft

[cHAP. 14

F1g. 14-3. Initial critical inventories of U235 in two-region, homogeneous, molten
fluoride-salt reactors. External fuel volume, 339 ft3.

 

 

 

 

0.8
| | | | |
Numbers On Data Points
Are Core Diameters In Feet ( ) mole % ThF4 In Fuel Salt
0.7 — —
0 W10
98 N
2 pro | e (o.n)\fi
né 0.6 — 70’ Te————— 06 b —]
o 1 5 (0.25)
3 5p ¢ .-
& 7.,/\No ThF4 In Fuel Salt
> [ %10
& 0.5 |— 8/0 —
%0
0.4 |— —
0.3 100 | | | | l
0 200 400 600 800

1000 1200

Critical Inventory, kg U235

F1a. 14-4. Initial fuel regeneration in two-region, homogeneous, molten fluoride-
salt reactors fueled with U235, Total power, 600 Mw (heat); external fuel volume,

339 ft3; core and blanket salts No. 1.
14-1] HOMOGENEOUS REACTORS FUELED WITH U?3% 635

 

0.8
| l |
( ) mole % ThF4 In Fuel Salt
0.7 — nEeEy —
- 10 —
o— =
° 9./(0.75) m
s 6 7 (0.50) '
% 0.6 |— ®(0.25) ]
.2
B
c
®
o No ThF 4 In Fuel Salt
205 | —
0.4 |— Numbers On Data Points ]
’ Are Core Diameters In Feet
0.3 | | |

 

 

 

0 200 400 4600 800
Critical Inventory, kg of 235

Fre. 14-5. Maximum initial regeneration ratios in two-region, homogeneous,
molten fluoride-salt reactors fueled with U235, Total power, 600 Mw (heat); ex-
ternal fuel volume, 339 ft3. |

range from slightly above 100 kg in an 8-ft-diameter core with no thorium
present to 1500 kg in a 5-ft-diameter core with 1 mole 9, ThF4 present.

The optimum combination of core diameter and thorium concentration
is, qualitatively, that which minimizes the sum of inventory charges (in-
cluding charges on Li?, Be, and Th) and fuel reprocessing costs. The fuel
costs are directly related to the regeneration ratio, and this varies in a
complex manner with inventory of U235 and thorium concentration, as
shown in Fig. 14—4. It may be seen that at a given thorium concentration,
the regeneration ratio (with one exception) passes through a maximum as
the core diameter is varied between 5 and 10 ft. These maxima increase
with increasing thorium concentration, but the inventory values at which
they occur also Increase.

Plotting the maximum regeneration ratio versus critical inventory
generates the curve shown in Fig. 14-9. It may be seen that a small in-
vestment in U235 (200 kg) will give a regeneration ratio of 0.58, that 400 kg
will give a ratio of 0.66, and that further increases in fuel inventory have
little effect.

The effects of changes in the compositions of the fuel and blanket salts
are indicated in the following description of the results of a series of calcu-
lations for which salts with more favorable melting points and viscosities
were assumed. The BeFs content was raised to 37 mole % in the fuel salt

 
636 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

 

| | !

mole % ThFy in Fuel Salt

Inventory , kg of U235

 

 

 

 

Core Diameter, f

Fia. 14-6. Initial critical inventories of U235 in two-region, homogeneous, molten

fluoride-salt reactors. Total power, 600 Mw (heat); external fuel volume, 339 ft3;
core and blanket salts No. 2.

 

 

0.7
| | l
Numbers On Data Points are
Core Diameters In Feet
8.
o \, \8\
B 7
S 8 .1.\. \. 7 .
= \ ® (1)
O 0.6 — ® 1 ® ]
e o \ 6 \ \‘
> o« ° ° S
& 7o (0.50) (0.75)
[
8e (0.25)
() mole % ThF4 In Fuel Salt
0.5 | | | | |

 

 

o 200 400 600 800 1000 1200
Critical Inventory, U235, kg

F1G. 14-7. Initial fuel regeneration in two-reg
salt reactors fueled with U235, Total

339 ft3; core and blanket salts No. 2

ion, homogeneous, molten fluoride-
power, 600 Mw (heat); external fuel volume,
14-1] HOMOGENEOUS REACTORS FUELED WITH U230 637

(fuel salt No. 2), and the blanket composition (blanket salt No. 2) was
fixed at 13 mole % ThF4, 16 mole % BeFs, and 71 mole % LiF. Blanket
salt No. 2 is a somewhat better reflector than No. 1, and fuel salt No. 2 a
somewhat better moderator. As a result, at a given core diameter and
thorium concentration in the fuel salt, both the critical concentration and
the regeneration ratio are somewhat lower for the No. 2 salts.

Reservations concerning the feasibility of constructing and guaranteeing
the integrity of core vessels in large sizes (10 ft and over), together with
preliminary consideration of inventory charges for large systems, led to
the conclusion that a feasible reactor would probably have a core diameter
lying in the range between 6 and 8 ft. Accordingly, a parametric study in
this range with the No. 2 fuel and blanket salts was performed. In this
study the presence of an outer reactor vessel consisting of 2/3 in. of
INOR-8 was taken into account. The results are presented in Table 14-2
and Figs. 14-6 and 14-7. In general, the nuclear performance is somewhat
better with the No. 2 salt than with the No. 1 salt.

Neutron balances and miscellaneous details. The distributions of the
neutron captures are given in Tables 14-1 and 14-2, where the relative
hardness of the neutron spectrum is indicated by the median fission energies
and the percentages of thermal fissions. It may be seen that losses to Li,
Be, and F in the fuel salt and to the core vessel are substantial, especially
in the more thermal reactors (e.g., Case No. 18). However, in the thermal
reactors, losses by radiative capture in U?3° are relatively low. Increasing
the hardness decreases losses to salt and core vessel sharply (Case No. 5),
but increases the loss to the n—y reaction. It is these opposing trends
which account for the complicated relation between regeneration ratio and
critical inventory exhibited in Figs. 14—4 and 14-7. The numbers given
for capture in the Li and F in the blanket show that these elements are well
shielded by the thorium in the blanket, and the leakage values show that
leakage from the reactor is less than 0.01 neutron per neutron absorbed in
U235 in reactors over 6 ft in diameter. The blanket contributes sub-
stantially to the regeneration of fuel, accounting for not less than one-third
of the total even in the 10-ft-diameter core containing 1 mole % ThF4.

Effect of substitution of sodium for L17. In the event that Li7 should prove
not to be available in quantity, it would be possible to operate the reactor
with mixtures of sodium and beryllium fluorides as the basic fuel salt. The
penalty imposed by sodium in terms of critical inventory and regeneration
ratio is shown in Fig. 14-8, where typical Na—Be systems are compared
with the corresponding Li-Be systems. With no thorium in the core, the
use of sodium increases the critical inventory by a factor of 1.5 (to about
300 kg) and lowers the regeneration ratio by a factor of 2. The regeneration
penalty is less severe, percentagewise, with 1 mole % ThF4 in the fuel
salt; in an 8-ft-diameter core, the inventory rises from 800 kg to 1100 kg
[cHAP. 14

NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS

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640 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

 

0.8 l l |

Li—-Be Salt with 1 mole
]0. % ThF4
. .
06— ¢ Fuel Salt A 8 \z-— —
®

o
£ kNo ThF4 8e-—e
o
Fuel Salt A 10
c
2 0.4 __.8. Fuel Salt B/ s —
g ’ | 67 Na-Be Salt
g 100 d No ThF4 With 1 mole % ThF4
5 £
- 8 Fuel Salt B

|

Numbers On Data Points Are
Core Diameters In Feet

| l |
0 400 800 1200 1600

 

 

 

Critical Inventory, Kg of y23s5

F1a. 14-8. Comparison of regeneration ratio and critical inventory in two-region,
homogeneous, molten fluoride-salt reactors fueled with U235, Fuel salt A: 37 mole
% BeF2 plus 63 mole %, Li’F. Fuel salt B: 46 mole 9, BeF; plus 54 mole 9, NaF.

and the regeneration ratio falls from 0.62 to 0.50. Details of the neutron
balances are given in Table 14-3.

Reactwity coefficients. By means of a series of calculations in which the
thermal base, the core radius, and the density of the fuel salt are varied
independently, the components of the temperature coefficient of reactivity
of a reactor can be estimated as illustrated below for a core 8 ft in diam-
eter and a thorium concentration of 0.75 mole 9} in the fuel salt at 1150°F.
From the expression

k=f(T, p, R),

where k is the multiplication constant, 7' is the mean temperature in the
core, p is the mean density of the fuel salt in the core, and R is the core
radius, it follows that

1dk 1[0k 1/0k\ dR , 1/0k\ dp
Efi—i(5_7_’>p,ze+70<573>p,1’3—7—’+%<55>3,Td_T’ (14-1)

where the term (1/k)(0k/3T),, r represents the fractional change in k due
to a change in the thermal base for slowing down of neutrons, the term
(1/k)(0k/0p)r, r represents the change due to expulsion of fuel from the
core by thermal expansion of the fluid, and the term (1/k)(dk/ OR),, 7
represents the change due to an increase in core volume and fuel holding
641

HOMOGENEOUS REACTORS FUELED WITH U?23°

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642 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

capacity. The coefficient dR/dT may be related to the coefficient for linear
expansion, &, of INOR-8, viz:

dR
d_T = Ro.

Likewise the term dp/dT may be related to the coefficient of cubical ex-
pansion, 3, of the fuel salt:

do _
dT pB.

From the nuclear calculations, the components of the temperature co-
efficient were estimated, as follows:

1/0k _ ~5/0
E(ST)p, _=—(0.13£0.02) X 10~5/°F,

R<6k

k a\R),,, =+ 0.412 £ 0.0005,

p 916) _
k< 3p e T 0.405 4 0.0005.

The linear coefficient of expansion, o, of INOR-8 was estimated to be
(8.01+0.5) X 1075/°F [5], and the coefficient of cubical expansion, 3, of
the fuel was estimated to be (9.889 4 0.005) X 10~ 5/°F from a correlation
of the density given by Powers [6]. Substitution of these values in
Eq. (14-1) gives

a e (3.80 4 0.04) X 10-5/°F

Y| =
s}

for the temperature coefficient of reactivity of the fuel. In this calculation,
the effects of changes with temperature in Doppler broadening and satura-
tion of the resonances in Th and U235 were not taken into account. Since
the effective widths of the resonances would be increased at higher tem-
peratures, the thorium would contribute a reactivity decrease and the
U?3% an increase. These effects are thought to be small, and they tend to
cancel each other.
14-1] HOMOGENEOUS REACTORS FUELED WITH U23° 643

Additional coefficients of interest are those for U23% and thorium. For
the 8-ft-diameter cores,

 

 

 

 

N(U235)< ok ) 14 [0.17N,(U?35) x 10~19]
k (9N<U235) NTh 247N, (U23%) X 10~ 19
and
N(Th)< ok ) _ N(Th) <_ak > dN,(U235)
k. \ON(Th)/vw=s  k \ON(UZ35)/)yw dN(Th)
where
w — 0.0595N (Th) x 10™1®
N (Th = 0.0805 ¢ X 10719,

In these equations, N (U?235) represents the atomic density of U%3° in atoms
per cubic centimeter, N,(U235) is the critical density of U235, and N(Th)
is the density of thorium atoms.

Heat release in core vessel and blanket. The core vessel of a molten-salt
reactor is heated by gamma radiation emanating from the core and blanket
and from within the core vessel itself. Estimates of the gamma heating can
be obtained by detailed analyses of the type illustrated by Alexander and
Mann [7]. The gamma-ray heating in the core vessel of a reactor with an
8-ft-diameter core and 0.5 mole 9% ThF4 in the fuel salt has been estimated
to be the following:

Source Heat release rate, w/cm3
Radioactive decay in core 1.4
Fission, n—y capture, and inelastic 5.2
scattering in core
n—y capture in core vessel 4.5
n—y capture in blanket 0.3

Total 1.4

Estimates of gamma-ray source strengths can be used to provide a crude
estimate of the gamma-ray current entering the blanket. For the 8-ft-
diameter core, the core contributes 45.3 w of gamma energy per square
centimeter to the blanket, and the core vessel contributes 6.8 w/cm?, which,
multiplied by the surface area of the core vessel, gives a total energy
escape into the blanket of 9.7 Mw. Some of this energy will be reflected
into the core, of course, and some will escape from the reactor vessel, and
644 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

therefore the value of 9.7 Mw is an upper limit. To this may be added the
heat released by capture of neutrons in the blanket. From the Oecusol-A
calculation for the 8-ft-diameter core and a fuel salt containing 0.5 mole %
ThF4 it was found that 0.176 of the neutrons would be captured in the
blanket. If an energy release of 7 Mev/capture is assumed, the heat release
at a power level of 600 Mw (heat) is estimated to be 8.6 Mw. The total is
thus 18.3 Mw or, say, 20 4+ 5 Mw, to allow for errors.

No allowance was made for fissions in the blanket. These would add 6
Mw for each 1% of the fissions occurring in the blanket. Thus it appears
that the heat release rate in the blanket might range up to 50 Mw.

14-1.2 Intermediate states. Without reprocessing of fuel salt. The nu-
clear performance of a homogeneous molten-salt reactor changes during
operation at power because of the accumulation of fission products and
nonfissionable isotopes of uranium. It is necessary to add U235 to the fuel
salt to overcome these poisons and, as a result, the neutron spectrum is
hardened and the regeneration ratio decreases because of the accompanying
decrease in 7 for U235 and the increased competition for neutrons by the
poisons relative to thorium. The accumulation of the superior fuel U233
compensates for these effects only in part. The decline in the regeneration
ratio and the increase in the critical inventory during the first year of
operation of three reactors having 8-ft-diameter cores charged, respectively,
with 0.25, 0.75, and 1 mole 9, ThF4 are illustrated in Fig. 14-9. The criti-
cal inventory increases by about 300 kg, and the regeneration ratio falls
about 16%,. The gross burnup of fuel in the reactor charged with 1 mole 7,
ThF4 and operated at 600 Mw with a load factor of 0.809, amounts to
about 0.73 kg/day. The U235 burnup falls from this value as U233 assumes
part of the load. During the first month of operation, the U235 burnup
averages 0.69 kg/day. Overcoming the poisons requires 1.53 kg more and
brings the feed rate to 2.22 kg/day. The initial rate is high because of the
holdup of bred fuel in the form of Pa233. As the concentration of this iso-
tope approaches equilibrium, the U235 feed rate falls rapidly. At the end
of the first year the burnup rate has fallen to 0.62 kg/day and the feed rate
to 1.28 kg/day. At this time U233 contributes about 129, of the fissions.
The reactor contains 893 kg of U235, 70 kg of U233, 7 kg of Pu23®, 62 kg
of U?3%, and 181 kg of fission products. The U236 and the fission products
capture 1.8 and 3.89, of all neutrons and impair the regeneration ratio
by 0.10 units. Details of the inventories and concentrations are given in
Table 144.

With reprocessing of fuel salt. If the fission products were allowed to ac-
cumulate indefinitely, the fuel inventory would become prohibitively large
and the neutron economy would become very poor. However, if the fission
products are removed, as described in Chapter 12, at a rate such that the
14-1] HOMOGENEOUS REACTORS FUELED WITH U235 645

 

 

 

 

0.7
[ |

2

©
-

c —
2

B

o 1 mole % ThFy4 in Fuel Salt

> 0.75

®

2 —

0.50

|

l—-—Processing Begins

 

 

 

 

 

 

Inventory, Kg

   

No Processing ——s]s— Processing Begins

100 |- || mole % ThF4
075

0.50

 

 

l
1 2 3

0 |

Time of Operation, years

Fia. 14-9. Operating performance of two-region, homogeneous, molten flyoride-
salt reactors fueled with U235, Core diameter, 8 ft; total power, 600 Mw (heat);
load factor, 0.80.

equilibrium inventory is, for example, equal to the first year’s production,
then the increase in U235 inventory and the decrease in regeneration ratio
are effectively arrested, as shown in Fig. 14-10. The fuel-addition rate
drops immediately from 1.28 to 0.73 kg/day when processing is started.
At the end of two years, the addition rate is down to 0.50 kg/day, and it
continues to decline slowly to 0.39 kg/day after 20 years of operation.
The nonfissionable isotopes of uranium continue to accumulate, of course,
but these are nearly compensated by the ingrowth of U233, As shown in
Fig. 14-10, the inventory of U235 actually decreases for several years in a
typical case, and then increases only moderately during a lifetime of
20 years.

The rapid increase in critical inventory of U235 during the first year can
be avoided by partial withdrawl of thorium. In Fig. 14-10 the dashed
lines indicate the course of events when thorium is removed at the rate of
1/900 per day. Burnup reduces the thorium concentration by another
646 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

 

 

 

 

 

 

 

 

 

 

 

 

0.7
o
o
£ 0.6 B
5 1 mole % ThF4 In Fuel Salt
o | S=__ —
§7 0.5 S —— Decreasing ThF 4
© T T e——— e
Y | | l I
1200
T T T T T T T 1
™ / |
2: 1000— / B
6 1 mole % ThF4 In Fuel Salt
’ 800
m — —— —
«Q o .
- Decreasing ThF ,
s 600
<z
S r 1 mole % ThF4 In Fuel Salt ]
c
: 233
2ol o e U
== (Accumulated)
0 10 20

Time of Operation, years

F1g. 14-10. Long-term nuclear performance of typical two-region, homogeneous,
molten fluoride-salt reactors fueled with U235, Core diameter, 8 ft; total power,
600 Mw (heat); load factor, 0.80.

1/4300 per day. The U235 inventory rises to 826 kg and then falls, at the
end of eight months, to 587 kg. At this time, the processing rate is in-
creased to 1/240 per day (eight-month cycle), but the thorium is returned
to the core and the thorium concentration falls thereafter only by burnup.
It may be seen that the U23% inventory creeps up slowly and that the re-
generation ratio falls slowly. The increase in U235 inventory could have
been prevented by withdrawing thorium at a small rate; however, the re-
generation ratio would have fallen somewhat more rapidly, and more U235
feed would have been required to compensate for burnup.

14—2. HoMoGENEOUS REACTORS FUELED wiTH U233

Uranium-233 is a superior fuel for use in molten fluoride-salt reactors in
almost every respect. The fission cross section in the intermediate range of
neutron energies is greater than the fission cross sections of U235 and
Pu?39. Thus initial critical inventories are less, and less additional fuel is
required to override poisons. Also, the parasitic cross section is sub-
stantially less, and fewer neutrons are lost to radiative capture. Further,
the radiative captures result in the immediate formation of a fertile iso-
647

HOMOGENEOUS REACTORS FUELED WITH U233

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649

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650 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

tope, U234, The rate of accumulation of U236 is orders of magnitude smaller
than with U235 as a fuel, and buildup of Np237 and Pu?3? is negligible.

The mean neutron energy is rather nearer to thermal in these reactors
than it is in the corresponding U235 cases. Consequently, losses to core
vessel and to core salt tend to be higher. Both losses will be reduced sub-
stantially at higher thorium concentrations.

14-2.1 Initial states. Results from a parametric study of the nuclear
characteristics of two-region, homogeneous, molten fluoride-salt reactors
fueled with U233 are given in Table 14-5. The core diameters considered
range from 3 to 10 ft, and the thorium concentrations range from 0.25 to 1
mole 9,. Although the regeneration ratios are less than unity, they are very
good compared with those obtained with U235, With 1 mole 9, ThF4 in
an 8-ft-diameter core, the U233 inventory was only 196 kg, and the re-
generation ratio was 0.91.

The regeneration ratios and fuel inventories of reactors of various diam-
eters containing 0.25 mole 9, thorium and fueled with U235 or U233 are
compared in Fig. 14-11. The superiority of U233 is obvious.

 

 

 

].O , |
U233

2
S 08— —
c
0
B
c y235
o
o \

0.4 I |

0 200 400 600

 

Critical Inventory, kg of U

Fia. 14-11. Comparison of regeneration ratios in molten-salt reactors containing
0.25 mole 9, ThF4 and U235- or U233-enriched fuel.

14-2.2 Intermediate states. Calculaticns of the long-term performance
of one reactor (Case 51, Table 14-5) with U233 as the fuel are described
below. The core diameter used was 8 ft and the thorium concentration
was 0.75 mole 9%,. The changes in inventory of U233 and regeneration ratio
are listed in Table 14-6. During the first year of operation, the inventory
rises from 129 to 199 kg, and the regeneration ratio falls from 0.82 to 0.71.
If the reprocessing required to hold the concentration of fission products
651

14-2]

HOMOGENEOUS REACTORS FUELED WITH U233

 

 

 

 

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653

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NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS

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656 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

and Np237 constant is begun at this time, the inventory of U233 increases
slowly to 247 kg and the regeneration ratio rises slightly to 0.73 during the
next 19 years. This constitutes a substantial improvement over the per-
formance with U235,

14-3. HoMOoGENEOUS REACTORS FUELED WITH PLUTONIUM

It may be feasible to burn plutonium in molten fluoride-salt reactors.
The solubility of PuF3 in mixtures of LiF and BeF'2 is considerably less
than that of UF4, but is reported to be over 0.2 mole % [8], which may be
sufficient for criticality even in the presence of fission fragments and non-
fissionable isotopes of plutonium but probably limits severely the amount
of ThF, that can be added to the fuel salt. This limitation, coupled with
the condition that Pu23? is an inferior fuel in intermediate reactors, will
result in a poor neutron economy in comparison with that of U?33-fueled
reactors. However, the advantages of handling plutonium in a fluid fuel
system may make the plutonium-fueled molten-salt reactor more desirable
than other possible plutonium-burning systems. ‘

14-3.1 Initial states. Critical conceniration, mass, tnventory, and regen-
eration ratio. The results of calculations of a plutonium-fueled reactor
having a core diameter of 8 ft and no thorium in the fuel salt are described
below. The critical concentration was 0.013 mole % PuF3, which is an
order of magnitude smaller than the solubility limits in the fluoride salts
of interest. The critical mass was 13.7 kg and the critical inventory in a
600-Mw system (339 ft3 of external fuel volume) was only 31.2 kg.

The core was surrounded by the Li—-Be-Th fluoride blanket mixture
No. 2 (13% ThF}4). Slightly more than 19% of all neutrons were captured
in the thorium to give a regeneration ratio of 0.35. By employing smaller
cores and larger investments in Pu23®, however, it should be possible to
increase the regeneration ratio substantially.

* Neutron balance and miscellaneous details. Details of the neutron economy
of a reactor fueled with plutonium are given in Table 14-7. Parasitic cap-
tures in Pu23? are relatively high; n is 1.84, compared with a v of 2.9. The
neutron spectrum is relatively soft; almost 60% of all fissions are caused
by thermal neutrons and, as a result, absorptions in lithium are high.

14-3.2 Intermediate states. On the basis of the average value of a of
Pu?3? it is estimated that Pu?40 will accumulate in the system until it cap-
tures, at equilibrium, about half as many neutrons as Pu?3°. While these
captures are not wholly parasitic, inasmuch as the product, Pu®!!, is
fissionable, the added competition for neutrons will necessitate an increase
in the concentration of the Pu239. Likewise, the ingrowth of fission products
14-4] HETEROGENEOUS GRAPHITE-MODERATED REACTORS 657

will necessitate the addition of more Pu?39. Further, the rare earths among
the fission products may exert a common-ion influence on the plutonium
and reduce its solubility. On the credit side, however, is the U233 produced
in the blanket. If this is added to the core it may compensate for the in-
growth of Pu®*0 and reduce the Pu?3® requirement to below the solubility
limit, and it may be possible to operate indefinitely, as with the U235-
fueled reactors.

14—4. HETEROGENEOUS GRAPHITE-MODERATED REACTORS

The use of a moderator in a heterogeneous lattice with molten-salt
fuels is potentially advantageous. First, the approach to a thermal neutron
spectrum improves the neutron yield, %, attainable, especially with U235

TABLE 14-7
INITIAL-STATE NUCLEAR CHARACTERISTICS OF A
TypicAL MoLTEN FLUORIDE-SALT REACTOR
FueELED wiTH Pu239

 

 

 

 

Core diameter: 8 ft.
External fuel volume: 339 ft3.
Total power: 600 Mw (heat).
Load factor: 0.8.
Critical inventory: 31.2 kg of Pu239,
Critical concentration: 0.013 mole 9, Pu?39,
Neutrons absorbed per
neutrons absorbed in Pu239
Neutron absorbers
Pu?39 (fissions) 0.630
Pu23? (n—y) 0.372
Li% and Li7 in fuel salt 0.202
Be?® in fuel salt 0.022
F19 in fuel salt 0. 086
Core vessel 0.145
Th in blanket salt - 0.352
Li-Be-F in blanket salt 0.024
Reactor vessel 0.004
Leakage 0.003
Neutron yield, 5 1.84
Thermal fissions, %, 59
Regeneration ratio 0.352

 

 

 

 
058 NUCLEAR ASPECTS

TABLE 14-8
COMPARISON OF GRAPHITE-MODERATED M OLTEN-SALT
AND Li1QUID-M ETAL-FUELED ‘REACTORS

OF MOLTEN-SALT REACTORS

[cHAP. 14

 

 

 

 

 

 

LMFR MSFR-1 MSFR-2
Total power, Mw (heat) 580 600 600
Over-all radius, in. 75 75 72
Critical mass, kg of U233 9.9 9.6 27 .7
Critical inventory, kg of U233* 467 77.8 213
Regeneration ratio 1.107 | 0.83 1.07
Core
Radius, in. 33 33 34.8
Graphite, vol 9 45 45 45
Fuel fluid, vol 9 55 55 55
Fuel components, mole %,
Bi ~100
LiF 69 61
BeF 2 31 36.5
ThF, 2.5
Unmoderated blanket
Thickness, in. 6 6 13.2
Composition, mole %
Bi 90
Th 10 (Th) | 10 (ThF,) | 13 (ThFy)
LiF 70 71
BeF 2 20 16
U233 0.015 0.014
Moderated blanket
Thickness, in. 36 36 24
Composition, vol 9,
Graphite 66.6 66.6 100
Blanket fluid{ 33.4 33.4
Neutron absorption ratiof
Th in fuel fluid 0.566
U233 in fuel fluid 0.918 0.925 1.000
Other components of fuel fluid 0.081 0.324 0.106
Th in blanket fluid 1.110 0.825 0.490
U233 in blanket fluid 0.083 0.071
Other components of blanket fluid 0.040 0.092 0.038
Leakage 0.012 0.004 0.014
Neutron yield, n 2.24 2.24 2.21

 

 

 

*With bismuth, the external volume indicated in Ref. 10 was used. The molten-
salt systems are calculated for 339 ft3 external volumes.

1Same as unmoderated blanket fluid.

iNeutrons absorbed per neutron absorbed in U233,
14-4] HETEROGENEOUS GRAPHITE-MODERATED REACTORS 659

and Pu?3®. Second, in a heterogeneous system, the fuel is partially shielded
from neutrons of intermediate energy, and a further improvement in ef-
fective neutron yield, 7, results. Further, the optimum systems may prove
to have smaller volumes of fuel in the core than the corresponding fluorine-
moderated, homogeneous reactors and, consequently, higher concentrations
of fuel and thorium in the melt. This may substantially reduce parasitic
losses to components of the carrier salt. On the other hand, these higher
concentrations tend to increase the inventory in the circulating-fuel
system external to the core. The same considerations apply to fission prod-
ucts and to nonfissionable isotopes of uranium.

Possible moderators for molten-salt reactors include beryllium, BeO,
and graphite. The design and performance of the Aircraft Reactor Experi-
ment, a beryllium-oxide moderated, sodium-zirconium fluoride salt, one-
region, U2?3°fueled burner reactor has been reported (see Chapter 16).
Since beryllium and BeO and molten salts are not chemically compatible,
it was necessary to line the fuel circuit with Inconel. It is easily estimated
that the presence of Inconel, or any other prospective containment metal
in a heterogeneous thermal reactor would seriously impair the regeneration
ratio of a converter-breeder. Consequently, beryllium and BeO are elimi-
nated from consideration.

Preliminary evidence indicates that uranium-bearing molten salts may
be compatible with some grades of graphite and that the presence of the
graphite will not carburize metallic portions of the fuel circuit seriously [9].
It therefore becomes of interest to explore the capabilities of the graphite-
moderated systems. The principal independent variables of interest are
the core diameter, fuel channel diameter, lattice spacing, and thorium
concentration.

14—4.1 Initial states. Two cases of graphite-moderated molten-salt re-
actors have been calculated for the same geometry and graphite-to-fluid
volume ratio as those for the reference-design LMFR [10]. The results for
these two cases, together with those for the liquid bismuth case, are sum-
marized in Table 14-8. Only the initial states are considered, and a metallic
shell to separate core and blanket fluids has not been included. With no
thorium in the core fluid, the molten-salt-fueled reactor has a significantly
lower regeneration ratio than that of the liquid-metal-fueled reactor, with
only a slightly lower critical mass. Adding 2.5 mole 9% ThF4 to the core
fluid increases the initial regeneration ratio to about 1.07, with a critical
mass and a corresponding total fuel inventory that are acceptably low.
660 NUCLEAR ASPECTS OF MOLTEN-SALT REACTORS [cHAP. 14

REFERENCES

1. R. L. MackLiN, Neutron Activation Cross Sections with Sb-Be Neutrons,
Phys. Rev. 107, 504-508 (1957).

2. L. G. ALEXANDER et al., Operating Instructions for the Unwac Program
Ocusol-A, A Modification of the Eyewash Program, USAEC Report CF-57-6-4,
Oak Ridge National Laboratory, 1957.

3. J. H. ALexanper and N. D. GiveN, A Machine Multigroup Calculation.
The Eyewash Program for Univac, USAEC Report ORNL-1925, Oak Ridge
- National Laboratory, 1955.

4. J. T. RoertTs and L. G. ALEXANDER, Cross Sections for the Ocusol-A
Program, USAEC Report CF-57-6-5, Oak Ridge National Laboratory, 1957.

5. B. W. Kinvon, Oak Ridge National Laboratory, 1958, personal communi-
cation.

6. W. D. Powers, Oak Ridge National Laboratory, 1958, personal communi-
cation.

7. L. G. ALExaNDER and L. A. MaNN, First Estimate of the Gamma Heating
in the Core Vessel of a Molten Fluoride Converter, USAEC Report CF-57-12-77,
Oak Ridge National Laboratory, 1957. '

8. C. J. BarTon, Solubility and Stability of PuF3 tn Fused Alkali Fluoride-
Beryllium Fluoride, USAEC Report ORNL-2530, Oak Ridge National Labora-
tory, 1958.

9. F. Kertesz, Oak Ridge National Laboratory, 1958, personal communica-
tion.

10. BaBcock AND WiLcox Co., Liquid Metal Fuel Reactor, Technical Feasi-
bility Report, USAEC Report BAW-2(Del.), 1955.
CHAPTER 15

EQUIPMENT FOR MOLTEN-SALT REACTOR HEAT-TRANSFER
SYSTEMS*

The equipment required in the heat-transfer circuits of a molten-salt
reactor consists of the components needed to contain, circulate, cool, heat,
and control molten salts at temperatures up to 1300°F. Included in such
systems are pumps, heat exchangers, piping, expansion tanks, storage
vessels, valves, devices for sensing operating variables, and other auxiliary
equipment.

Pumps for the fuel and blanket salts differ from standard centrifugal
pumps for operation at high temperatures in that provisions must be made
to exclude oxidants and lubricants from the salts, to prevent uncontrolled
escape of salts and gases, and to minimize heating and irradiation of the
drive motors. Heat is transferred from both the fuel and the blanket salts
to sodium in shell-and-tube heat exchangers designed to maximize heat
transfer per unit volume and to minimize the contained volume of salt, es-
pecially the fuel salt.

Seamless piping is used, where possible, to minimize flaws. Thermal ex-
pansion is accommodated by prestressing the pipe and by using expansion
loops and joints. Heaters and thermal insulation are provided on all com-
ponents that contain salt or sodium for preheating and for maintaining
the circuits at temperatures above the freezing points of the liquids and
to minimize heat losses. Devices are provided for sensing flow rates, pres-
sures, temperatures, and liquid levels. The devices include venturi tubes,
pressure transmitters, thermocouples, electrical probes, and floats. Inert
gases are used over free-liquid surfaces to prevent oxidation and to apply
appropriate base pressures for suppressing cavitation or moving liquid or
gas from one vessel to another.

The deviations from standard practice required to adapt the various
components to the molten-salt system are discussed below. The schematic
diagram of a molten-salt heat-transfer system presented in Fig. 15-1 in-
dicates the relative positions of the various components. For nuclear op-
eration, an off-gas system is supplied, as described in Chapter 17. The
vapor condensation trap indicated in Fig. 15-1 is required only on systems
that contain ZrF4 or a comparably volatile fluoride as a component of the
molten salt.

*By H. W. Savage, W. F. Boudreau, E. J. Breeding, W. G. Cobb, W. B. Me-
Donald, H. J. Metz, and E. Storto.

661
662 MOLTEN-SALT REACTOR HEAT-TRANSFER EQUIPMENT [CcHAP. 15

Pump (Includes Expansion Volume)

 

 

 

 

 

 

  
     

 

 

 

 

 

 

 

 

Liquid Level Vapor Trap
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Fic. 15-1. A molten-salt heat-transfer system.

'15-1. Pumps FOR MOLTEN SALTS

Centrifugal pumps with radial or mixed-flow types of impeller have been
used successfully to circulate molten-salt fuels. The units built thus far
and those currently being developed have a vertical shaft which carries
the impeller at its lower end. The shaft passes through a free surface of
liquid to isolate the motor, the seals, and the upper bearings from direct
contact with the molten salt. Uncontrolled escape of fission gases or entry
of undesirable contaminants to the cover gas above the free-liquid surface
in the pump are prevented either by the use of mechanical shaft seals or
hermetic enclosure of the pump and, if necessary, the motor. Thermal and
radiation shields or barriers are provided to assure acceptable temperature
and radiation levels in the motor, seal, and bearing areas. Liquid cooling
of internal pump surfaces is provided to remove heat induced by gamma
and beta radiation.
15-1] PUMPS FOR MOLTEN SALTS 663

    
   
 
 

 

  

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Fie. 15-2. Sump-type centrifugal pump developed for the Aircraft Reactor
Experiment.

The principles used in the design of pumps for normal liquids are applic-
able to the hydraulic design of a molten-salt pump. Experiments have
shown that the cavitation performance of molten-salt pumps can be pre-
dicted from tests made with water at room temperature. In addition to
stresses induced by normal thermal effects, stresses due to radiation must
be taken into account in all phases of design.

The pump shown in Fig. 15-2 was developed for 2000-hr durability at
very low irradiation levels and was used in the Aircraft Reactor Experi-
ment for circulating molten salts and sodium at flow rates of 50 to 150 gpm,
at heads up to 250 ft, and at temperatures up to 1550°F. These pumps
have been virtually trouble-free in operation, and many units in addition
to those used in the Aircraft Reactor Experiment have been used in devel-
opmental tests of various components of molten-salt systems.

The bearings, seals, shaft, and impeller form a cartridge-type subassembly
that is removable from the pump tank after opening a single, gasketed
664 MOLTEN-SALT REACTOR HEAT-TRANSFER EQUIPMENT [cHAP. 15

joint above the liquid level. The volute, suction, and discharge connec-
tions form parts of the pump tank subassembly into which the removable
cartridge is inserted. The upper portion of the shaft and a toroidal area
in the lower part of the bearing housing are cooled by circulating oil. Heat
losses during operation are reduced by thermal insulation.

In all the units built thus far nickel-chrome alloys have been used in
the construction of all the high-temperature wetted parts of the pump to
minimize corrosion. The relatively low thermal conductivity and high
strength of such alloys permitted close spacing of the impeller and bearings
and high thermal gradients in the shaft.

Thrust loads are carried at the top of the shaft by a matched pair of pre-
loaded angular-contact ball bearings mounted face-to-face in order to pro-
vide the flexibility required to avoid binding and to accommodate thermal
distortions. Either single-row ball bearings or a journal bearing can be
used successfully for the lower bearing.

The upper lubricant-to-air and the lower lubricant-to-inert-gas seals
are similar, rotary, mechanical face-type seals consisting of a stationary
graphite member operating in contact with a hardened-steel rotating mem-
ber. The seals are oil-lubricated, and the leakage of oil to the process side
is approximately 1 to 5 cc/day. This oil is collected in a catch basin and
removed from the pump by gas-pressure sparging or by gravity.

The accumulation of some 200,000 hr of relatively trouble-free test op-
eration in the temperature range of 1200 to 1500°F with molten salts and
liquid metals as the circulated fluids has proved the adequacy of this basic
pump design with regard to the major problem of thermally induced dis-
tortions. Four different sizes and eight models of pumps have been used
to provide flows in the range of 5 to 1500 gpm. Several individual pumps
have operated for periods of 6000 to 8000 hr, consecutively, without main-
tenance.

15-1.1 Improvements desired for power reactor fuel pump. The basic
pump described above has bearings and seals that are oil-lubricated and
cooled, and in some of the pumps elastomers have been used as seals be-
tween parts. The pump of this type that was used in the ARE was de-
signed for a relatively low level of radiation and received an integrated
dose of less than 5 X 108 r. Under these conditions both the lubricants and
elastomers used proved to be entirely satisfactory.

The fuel pump for a power reactor, however, must last for many years.
The radiation level anticipated at the surface of the fuel is 10° to 106 r/hr.
Beta- and gamma-emitting fission gases will permeate all available gas
space above the fuel, and the daughter fission products will be deposited
on all exposed surfaces. Under these conditions, the simple pump described
above would fail within a few thousand hours.
15-1] PUMPS FOR MOLTEN SALTS 665

Considerable improvement in the resistance of the pump and motor to
radiation can be achieved by relatively simple means. Lengthening the
shaft between the impeller and the lower motor bearing and inserting addi-
tional shielding material will reduce the radiation from the fuel to a low
level at the lower motor bearing and the motor. Hollow, metal O-rings or
another metal gasket arrangement can be used to replace the elastomer
seals. The sliding seal just below the lower motor bearing, which prevents
escape of fission-product gases or inleakage of the outside atmosphere,
must be lubricated to ensure continued operation. If oil lubrication is used,
radiation may quickly cause coking. Various phenyls, or mixtures of them,
are much less subject to formation of gums and cokes under radiation and
could be used as lubricant for the seal and for the lower motor bearing.
This bearing would be of the friction type, for radial and thrust loads.
These modifications would provide a fuel pump with an expected life of the
order of a year. With suitable provisions for remote maintenance and re-
pair, these simple and relatively sure improvements would probably
suffice for power reactor operation.

Three additional improvements, now being studied, should make pos-
sible a fuel pump that will operate trouble-free throughout a very long life.
The first of these is a pilot bearing for operation in the fuel salt. Such a
bearing, whether of hydrostatic or hydrodynamic design, would be com-
pletely unaffected by radiation and would permit use of a long shaft so
that the motor could be well shielded. A combined radial and thrust
bearing just below the motor rotor would be the only other bearing required.
The second improvement is a labyrinth type of gas seal to prevent escape
of fission gases up the shaft. There are no rubbing surfaces and hence no
need for lubricants, so there can be no radiation damage. The third inno-
vation is a hemispherical gas-cushioned bearing to act as a combined thrust
and radial bearing. It would have the advantage of requiring no auxiliary
lubrication supply, and it would combine well with the labyrinth type of
gas seal. It would, of course, be unaffected by radiation.

15-1.2 A proposed fuel pump. A pump design embodying these last
three features is shown in Fig. 15-3. It is designed for operation at a tem-
perature of 1200°F, a flow rate of 24,000 gpm, and a head of 70 ft of fluid.
The lower bearing is of the hydrostatic type and is lubricated by the
molten-salt fuel. The upper bearing, which is also of the hydrostatic type,
is cushioned by helium and serves also as a barrier against passage of gaseous
fission products into the motor. This bearing is hemispherical to permit
accommodation of thermally induced distortions in the over-all pump
structure.

The principal radiation shielding is that provided between the source
and the area of the motor windings. Layers of beryllium and boron for
666 MOLTEN-SALT REACTOR HEAT-TRANSFER EQUIPMENT [CHAP. 15

Shield Retainer

 
    

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Rotor
Laminations

   
  
  

Gas Circulation

Stator

   

Hydrostatic
Gas Bearing

   

Gas

  
  
  
  
  
 

Nuclear Radiation Shield

Gamma Shield
Support Ring

Neutron Shield

Purge Gas Flow

  
 
 

Maximum Fuel Level

/

 
 
  
 
 
  

Coolant
Blanket Salt Passage

By-Pass Flow

Fuel Expansion Tank

Minimum Fuel Level

Volute (Discharge to Rear)

  

\a

  
   
 

Hydrostatic Bearing Fuel Inlet

' 0 1 2 3

 

 

Scale—Feet

Frg. 15-3. Improved molten-salt fuel pump designed for power reactor use. Op-
erating temperature, 1200°F; flow rate, 24,000 gpm; head, 70 ft of fluid.
15-3] VALVES 667

neutron shielding and a heavy metal for gamma-radiation shielding are
proposed. The motor is totally enclosed, to eliminate the need for a shaft
seal. A coolant is circulated in the area outside the stator windings and
between the upper bearing and the shielding. Molten-salt fuel is circulated
over the surfaces of those parts of the pump which are in contact with the
gaseous fission products to remove heat generated in the metal.

15-2. HEaAT ExcHANGERS, ExpaNsioN TaANks, AND DRAIN TANKS

The heat exchangers, expansion tanks, and drain tanks must be especially
designed to fit the particular reactor system chosen. The design data of
items suitable for a specific reactor plant are described in Chapter 17.
The special problems encountered are the need for preheating all salt-
and sodium-containing components, for cooling the exposed metal surfaces
in the expansion tank, and for removing afterheat from the drain tanks.
It has been found that the molten salts behave as normal fluids during
pumping and flow and that the heat-transfer coefficients can be predicted
from the physical properties of the salts.

15-3. VALVES

The problems associated with valves for molten-salt fuels are the con-
sistent alignment of parts during transitions from room temperature to
1200°F, the selection of materials for mating surfaces which will not
fusion-bond in the salt and cause the valve to stick in the closed position,
and the provision of a gastight seal. Bellows-sealed, mechanically operated,
poppet valves of the type shown in Fig. 15-4 have given reliable service in
test systems. |

A number of corrosion and fusion-bond resistant materials for high-
temperature use were found through extensive screening tests. Molyb-
denum against tungsten or copper and several titanium or tungsten carbide-
nickel cermets mating with each other proved to be satisfactory. Valves
with very accurately machined cermet seats and poppets have operated
satisfactorily in 2-in. molten-salt lines at 1300°F with leakage rates of less
than 2 cc/hr. Consistent positioning of the poppet and seat to assure

leaktightness is achieved by minimizing transmission of valve-body dis-
~ tortions to the valve stem and poppet.

If rapid valve operation is not required, a simple “freeze’”’ valve may be
used to ensure a leaktight seal. The freeze valve consists of a section of
pipe, usually flattened, that is fitted with a device to cool and freeze a salt
plug and another means of subsequently heating and melting the plug.
668 MOLTEN-SALT REACTOR HEAT-TRANSFER EQUIPMENT [cHAP. 15

Bellows
Protector

 

 

 

 

Bellows

 

 

 

 

 

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SIS, /¢ L L RN Y

 

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F1gc. 15-4. Bellows-sealed, mechanically operated poppet valve for molten-salt
service. -

15-4. SysteEm HEATING

Molten-salt systems must be heated to prevent thermal shock during
filling and to prevent freezing of the salt when the reactor is not operating
to produce power. Straight pipe sections are normally heated by an elec-
tric tube-furnace type of heater formed of exposed Nichrome V wire in a
ceramic shell (clamshell heaters). A similar type of heater with the Ni-
chrome V wire installed in flat ceramic blocks can be used to heat flat
surfaces or large components, such as dump tanks, ete. In general, these
heaters are satisfactory for continuous operation at 1800°F. Pipe bends,
irregular shapes, and small components, such as valves and pressure-
measuring devices, are usually heated with tubular heaters (e.g., General
Electric Company *‘Calrods”) which can be shaped to fit the component
or pipe bend. In general, this type of heater should be limited to service at
15-5] JOINTS 669

1500°F. Care must be exercised in the installation of tubular heaters to
avold failure due to a hot spot caused by insulation in direct contact with
the heater. This type of failure can be avoided by installing a thin sheet of
metal (shim stock) between the heater and the insulation.

Direct resistance heating in which an electric current is passed directly
through a section of the molten-salt piping has also been used successfully.
Operating temperatures of this type of heater are limited only by the
corrosion and strength limitations of the metal as the temperature is
increased. Experience has indicated that heating of pipe bends by this
method is usually not uniform and can be accompanied by hot spots caused
by nonuniformity of liquid flow in the bend.

15-5. JoINTS

Failures of some system components may be expected during the de-
sired operating life, say 20 years, of a molten-salt power-producing reactor;
consequently, provisions must be made for servicing or removing and re-
placing such components. Remotely controlled manipulations will be
required because there will be a high level of radiation within the primary
shield. Repair work on or preparations for disposal of components that
fail will be carried out in separate hot-cell facilities.

The components of the system are interconnected by piping, and flanged
connections or welded joints may be used. In breaking connections between
a component and the piping, the cleanliness of the system must be pre-
served, and in remaking a connection, proper alignment of parts must be
re-established. The reassembled system must conform to the original
leaktightness specifications. Special tools and handling equipment will
be needed to separate components from the piping and to transport parts
within the highly radioactive regions of the system. While an all-welded
system provides the highest structural integrity, remote cutting of welds,
remote welding, and inspection of such welds are difficult operations.
Special tools are being developed for these tasks, but they are not yet
generally available. Flanged connections, which are attractive from the
point of view of tooling, present problems of permanence of their leak-
tightness.

Three types of flanged joints are being tested that show promise. One
is a freeze-flange joint that consists of a conventional flanged-ring joint
with a cooled annulus between the ring and the process fluid. The salt that
enters the annulus freezes and provides the primary seal. The ring provides
a backup seal against salt and gas leakage. The annulus between the ring
and frozen material can be monitored for fission product or other gas leak-
age. The design of this joint is illustrated in Fig. 15-5.
670 MOLTEN-SALT REACTOR HEAT-TRANSFER EQUIPMENT [CHAP. 15

—.H.<-Gap (~1/16in.)

-Soft-Iron or Copper Seal Ring

          

S

 

 

 

 

 

 

\“§‘
\ 23— Air Channel for Cooling
A N
Frozen -Salt Seal
AN
1
BN
BN
\ .
JW% \\ ‘\X&/Weld of Flange to Loop Tubing
Salt Flow
; Ring Insert to Provide Labyrinth
p for Salt Leakage
; ~—~Frozen-Salt Seal

 

Narrow Section to Reduce Heat
Transfer from the Molten Salt
in the Loop Tubing

 

\
X

SO IO 35

 

 

 

 

1 % 0 VY2 1
= e T ]
Inches

Air Inlet to Cooling Channel
Between Flange Faces Indicate Region of Frozen-Salt

Seal; Indicate Region of Transition from Liquid
to Solid Salt

Fig. 15-5. Freeze-flange joint for 1/2-in.-OD tubing.

 

 

 

 

 

 

 

+— Seal Material Insert

{ Shown Before Being
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AV /\V
77 7 72 W\
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Weld of / Inches
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Fi1a. 15-6. Cast-metal-sealed flanged joint.
15-6] INSTRUMENTS 671

  

Weld of Flange to
Loop Tubing

Copper Ring —
nches

Figc. 15-7. Indented-seal flange.

A cast-metal-sealed flanged joint is also being tested for use in vertical
runs of pipe. As shown in Fig. 15-6, this joint includes a seal which is cast
in place in an annulus provided to contain it. When the connection is to
be made or-broken the seal is melted. Mechanical strength is supplied by
clamps or bolts.

A flanged joint containing a gasket (Fig. 15-7) is the third type of joint
being considered. In this joint the flange faces have sharp, circular, mating
ridges. The opposing ridges compress a soft metal gasket to form the seal
between the flanges.

15—-6. INSTRUMENTS

Sensing devices are required in molten-salt systems for the measurement
of flow rates, pressures, temperatures, and liquid levels. Devices for these
services are evaluated according to the following criteria: (1) they must
be of leaktight, preferably all-welded, construction, (2) they must be cap-
able of operating at the maximum temperature of the fluid system, (3) their
accuracies must be relatively unaffected by changes in the system tem-
perature, (4) they should provide lifetimes at least as great as the lifetime
of the reactor, (5) each must be constructed so that, if the sensing element
fails, only the measurement supplied by it is lost. The fluid system to
which the instrument is attached must not be jeopardized by failure of the
sensing element.

15-6.1 Flow measurements. Flow rates are measured in molten-salt
systems with orifice or venturi elements. The pressures developed across the
sensing element are measured by comparing the outputs of two pressure-
measuring devices. Magnetic flowmeters are not at present sufficiently
sensitive for molten-salt service because of the poor electrical conductivity
of the salts.

 
672 MOLTEN-SALT REACTOR HEAT-TRANSFER EQUIPMENT [cHAP. 15

15-6.2 Pressure measurements. Measurements of system pressures re-
quire that transducers operate at a safe margin above the melting point of
the salt, and thus the minimum transducer operating temperature is usually
about 1200°F. The pressure transducers that are available are of two types:
(1) a pneumatic force-balanced unit and (2) a displacement unit in which
the pressure is sensed by displacement of a Bourdon tube or diaphragm.
The pneumatic force-balanced unit has the disadvantages that loss of the
instrument gas supply (usually air) can result in loss of the measurement,
and that failure of the bellows or diaphragm would open the process system
to the air supply or to the atmosphere. The displacement unit, on the other
hand, makes use of an isolating fluid to transfer the sensed pressure hydro-
statically to an isolated low-temperature output element. Thus, in the
event of a failure of the primary diaphragm, the process fluid would merely
mix with the isolating fluid and the closure of the system would be
unaffected.

15-6.3 Temperature measurements. Temperatures in the range of 800
to 1300°F are commonly measured with Chromel-Alumel or platinum-
platinum-rhodium thermocouples. The accuracy and life of a thermocouple
in the temperature range of interest are functions of the wire size and, in
general, the largest possible thermocouple should be used. Either beaded
thermocouples or the newer, magnesium oxide-insulated thermocouples
may be used.

15-6.4 Liquid-level measurements. Instruments are available for both
on-off and continuous level measurements. On-off measurements are made
with modified automotive-type spark plugs in which a long rod is used in
place of the normal center conductor of the spark plug. To obtain a con-
tinuous level measurement, the fluid head is measured with a differential
pressure instrument. The pressure required to bubble a gas into the fluid
1s compared with the pressure above the liquid to obtain the fluid head.
Resistance probe and float types of level indicators are available for use in
liquid-metal systems.

15-6.5 Nuclear sensors. Nuclear sensors for molten-salt reactors are
similar to those of other reactors and are not required to withstand high
temperatures. Existing and well-tested fission, ionization, and boron tri-
fluoride thermal-neutron detection chambers are available for installation
at all points essential to reactor operation. Their disadvantages of limited
life can be countered only by duplication or replacement, and provisions
can be made for this. It should be pointed out that the relatively large,
negative temperature coefficients of reactivity provided by most circulating-
fuel reactors make these instruments unessential to the routine operation
of the reactor.
CHAPTER 16
AIRCRAFT REACTOR EXPERIMENT*

The feasibility of the operation of a molten-salt-fueled reactor at a truly
high temperature was demonstrated in 1954 in experiments with a reactor
constructed at ORNL. The temperature of the fuel exiting from the core
of this reactor was about 1500°F, and the temperature of the fuel at the
inlet to the core was about 1200°F. The reactor was constructed before
the mechanism and control of corrosion by molten salts had been fully
explored, and therefore the experimental operation of the reactor was of
short duration. Since the work was supported by the Aircraft Reactors
Branch of the Atomic Energy Commission, the reactor was called the Air-
craft Reactor Experiment (ARE).} |

The ARE was a thermal reactor in which moderation was accomplished
by BeO blocks through which the fluoride fuel was circulated in Inconel
tubes arranged in a symmetrical, heterogeneous matrix. The Inconel ves-
sel containing the core was essentially a right cylinder, approximately
52 in. OD and 44 in. in height, with 2-in.-thick walls. The fuel passages
consisted of 1-in.-diameter Inconel tubes arranged in six parallel circuits,
and each circuit, by the use of reverse bends at top and bottom of the
core, made eleven passes through the core. The fuel passages did not
traverse the peripheral BeO blocks which served as a reflector around the
core of the reactor. A top view of the BeO blocks and the Inconel tubes
is shown in Fig. 16-1. The moderator and reflector blocks were cooled
by circulating liquid sodium from the bottom to the top of the pressure
vessel. The sodium permeated all interstices of the BeO and flowed rap-
idly through 1/2-in. vertical holes in the reflector sections of the BeO. An
elevation drawing of the reactor which illustrates these features is presented
in Fig. 16-2, and a photograph of the reactor vessel that was taken before
assembly of the thermal shield is shown in Fig. 16-3.

Since the purpose of the operation of this experimental reactor was to
study the behavior of the circulating-fluoride-fuel system and to identify
the problems associated therewith, the power output of the reactor was
not utilized but, rather, was simply dumped as heat. The heat-removal
system is shown schematically in Fig. 16-4. The fuel was circulated
through a finned-tube radiator type of heat exchanger. This radiator was
located within a sheet-metal housing of a toroidal shape. In another part
of the toroidal housing there was a second finned-tube radiator through

*By E. S. Bettis and W. K. Ergen.
R. C. Briant et al., Nuclear Science and Engineering, Vol. 2, No. 6, 795-853 (1957).

673
674 AIRCRAFT REACTOR EXPERIMENT [cHAP. 16

 

Fie. 16-1. Top view of the reactor core of the ARE. Hexagonal beryllium
oxide blocks serve as the moderator. Inconel tubes pass through the moderator
blocks to carry the molten-salt fuel.

which plant water flowed. A large centrifugal blower circulated the coolant
gas (helium) in the toroidal loop so that heat was picked up from the fuel
radiator and dumped into the water radiator.

An identical arrangement of radiators and blower was used for cooling
the sodium used as the moderator-reflector coolant. In the interest of
safety (for removal of afterheat in the event of a pump failure), the so-
dium circuit was installed in duplicate so that an entire sodium cooling
system was availlable as a spare. These two sodium loops were operated
alternately during the experiment in an effort to keep a check on the op-
erability of each loop. Had one loop failed to operate, the experiment
would have been terminated for lack of a spare cooling system.

The control system of the reactor was based on conventional practice.
The three safety shim rods were actuated by electrically driven lead screws
which moved electromagnets in a vertical plane. When these magnets were
driven to their lowest extremity, an armature was engaged to which the
shim (poison) rods were attached. Loss of current in the electromagnets
would allow the rods to fall under the action of gravity into thimbles in
the central region of the core. The regulating rod was a simple stainless-
steel pipe which was rigidly attached to a rack driven by a reversible elec-
CHAP. 16]

Reflector Coolant

Fic. 16-2. Elevative section of the Aircraft Reactor Experiment.

Safety Rod Assembly

Tubes

Fuel Tubes

AIRCRAFT REACTOR EXPERIMENT

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Scale in Inches

BeO Moderator
and Reflector

Thermal Shield

675

tric motor through a pinion. Fission chambers located in the reflector, as
well as ionization chambers located outside the pressure shell of the reactor,

furnished the neutron and gamma-ray signals for the control system.

The shim and control rods which entered the hot reactor core had to be
cooled to prevent overheating from neutron capture and gamma-ray ab-
sorption. This cooling was effected by circulating helium in a closed loop
that included a water-cooled radiator, as in the case of the fuel and sodium
circuits. This helium circuit was integral with a helium-filled monitoring
annulus which surrounded all fuel and sodium piping in the system. This
annulus was formed by putting a continuous stainless-steel sleeve around
all hot piping, and the helium circulated in the annulus performed two
functions: (1) it kept the hot lines at essentially an even temperature dur-
ing the warmup period when the system was heated by means of electrical
heating units placed on the outer surface of the annulus, and (2) the helium
was monitored to ensure that the fuel and sodium piping was leaktight.

Large, heated reservoir tanks were connected to the system through
isolation valves so that the sodium and the fused fluoride mixture could
676 AIRCRAFT REACTOR EXPERIMENT [cHAP. 16

 

Fig. 16-3. View of the ARE Vessel before addition of the thermal shield. The
external strip heaters with their electrical leads are shown in place.
CHAP. 16] AIRCRAFT REACTOR EXPERIMENT 677

 

Rod
Actuator

 

 

 

Helium Pump Pump Helium
Blower : Blower

  

  
  

Absorber Rod

   

  

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———

Heat Exchanger

 

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Heat Exchanger

Heat Exchanger

 

b
o
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o
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Reflector Coolant

Fic. 16-4. Schematic diagram of the heat-removal system for the ARE.

be pressurized from the tanks into the system and could be drained back
into the tanks after the experiment was over. Dry helium was used for
operating penumatic instruments and for pressurizing the liquids into the
system from the tanks.

Pumps for both the sodium and the molten-fluoride mixture consisted
of sump-type centrifugal pumps with overhanging shafts. The pumps were
mounted vertically, and a gas space was provided between the liquid level
and the upper bearings of the pump. The pumps were located so that the
free-liquid surface in the sump tank was the high point in both the fuel
and the sodium circuits. The sump tank of the pump also served as an
expansion tank for the liquid. The isometric drawing of the fuel system
presented in Fig. 16-5 indicates the relative levels of the components.

Both of the liquid systems, fuel and sodium, were fabricated entirely of
Inconel, and all closures were made by inert-gas-shielded electric-arc
(Heliarc) welding. The welding procedure was adopted after extensive
experimental research and developmental work, and meticulous care was
exercised in all welding operations. The entire reactor system, that is, the
reactor vessel, heat exchangers, pumps, dump tanks, piping, and auxiliary
equipment (with the exception of control rod drives), was located in con-
crete pits below ground level. After the reactor was brought to criticality
by manual fuel injection, concrete blocks were placed on top of the pits
to complete the shielding of the system as required during power operation.

Fuel was added as a molten mixture of NaF and UF4 (enriched in U23%)
after the sodium system had been heated and filled with sodium and the
fuel system had been heated and filled with fuel carrier—a molten mixture
of NaF and ZrF.. The fuel additions were made into the sump of the fuel
pump through the use of a temporary enrichment system that was capable
of injecting (by manual operation) a few hundred grams of fuel mixture
678 AIRCRAFT REACTOR EXPERIMENT [cHAP. 16

Standby Fuel Pump A2

     
  
  
  

7
7
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Disk Valve R

Heat
Exchanger &y
No. 2 b

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Hot Fuel
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Carrier
Fill Line 4
Reserve Tank No. 1

Fill Tank No. 2

F16. 16-5. Layout of the fuel system components for the ARE.

at a time. This method of fuel addition was laborious and time-consuming,
but it effectively and safely enriched the reactor to a critical concentration.

The reactor was taken to criticality essentially without incident. The
total amount of U235 added to the system to make the reactor critical was
approximately 61 kg, but small amounts of fuel were withdrawn from the
system for sampling and in trimming the pump level. The uranium con-
centration at criticality was 384 g/liter of fluoride mixture. The calculated
volume of the core was 38.8 liters at 1300°F, and thus the clean critical
mass of the reactor was 14.9 kg of U235,

It was demonstrated that the reactor had an over-all temperature co-
efficient of reactivity of —6 X 10~5 (Ak/k)/°F. As was anticipated, the
fast negative temperature coefficient of reactivity (associated with the fuel
expansion coefficient) served to stabilize the reactor power level. From a
power lever of 200 kw upward, the temperature coefficient controlled the
system so precisely that the reactor responded to load demands in a
thoroughly reliable manner.

The response of the reactor was demonstrated in a number of experi-
ments, one of which is described in Fig. 16-6. The abscissa, to be read from
right to left, is the time in minutes, and the print-outs from recorders
giving the reactor inlet and outlet temperatures are the ordinate. Initially,
in this experiment, the reactor was operating at low power. Then the heat
CHAP. 16] AIRCRAFT REACTOR EXPERIMENT 679

Reactor Inlet and Outlet Tube Temperatures

0045 hundreds o‘f" degrees F

 

0040 T
11111 teed Shim Rods Ins
0035 , Fuel and Na Blowers
‘ Low Temperature
3 0030 Interlock
o Reduced Blower
~ 0025 | Turlned o? Fuel
o ) i1]'| Helium B
— Cooled & 1'{ Helium Blower
> 0020 'Broughf (FU" Speed)
2 Inserted Regulating Rod
0015 Power Reactor Subcritical
Extraction Turned Off Fuel Helium
0010 Blower
0005
2400
> 2355 l1l{Turned on System
o Helium Blower
w 2350 (Full Speed)
g 2345 ‘ Inserted Shim Rods
= A Fuel System
2340 : Helium Blower Speed
to Zero
2335 |
B
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°:~ 2325
>
2 2320
System
2315 Blower on °
_ Increase Fuel
2310 s | System Helium Blower

» Speed to Maximum
2305 1 bl IR
| 1 "LReactor Critical:
2300 .. 1 Operating
| e

2255

Fic. 16-6. Chart of inlet and outlet temperatures for the ARE as influenced by
various experimental procedures.

extraction from the fuel was slowly increased and there was, first, a re-
sultant decrease in the temperature of the fuel which reached the reactor
inlet from the heat exchanger. This increased the reactivity and the re-
actor power, as indicated by the temperature rise at the reactor outlet.
The spread of inlet and outlet temperatures corresponds to a power level
of 2.5 Mw. When the heat extraction was reduced, the inlet temperature
680 AIRCRAFT REACTOR EXPERIMENT [cHAP. 16

rose and the outlet temperature fell until the two temperatures became
nearly coincident. As may be seen, the control rods did not determine the
power output; they only influenced the average temperature. Insertion
of the shim rods decreased the temperature. Another rapid increase in the
power demand on the fuel system again spread apart the inlet and outlet
temperature recordings, and full insertion and full withdrawal of the
regulating rod depressed and then raised both temperatures simultaneously.
Next, the power extraction was stopped and the regulating rod was in-
serted to make the reactor subecritical.

The third spread of the temperatures in Iig. 166 was a result of a
demonstration which showed that the reactor could be brought to criticality,
without use of the rods, by the power demand alone. Power extraction
from the sodium system cooled the reactor to make it critical, and power
extraction from the fuel again caused the spread of inlet and outlet tem-
peratures.

The remarkable stability of the system made it unexpectedly possible
to demonstrate that no more than 59, of the Xel3% was retained in the
molten fuel. It had been computed that the xenon poisoning after 27 hr
of operation at full power would amount to 2 X 1073 in Ak/k if all the
xenon formed stayed in the fuel until it decayed. This level of poisoning
was less than would be expected from the usual equations, partly because
the fuel spent only one-fourth of the time in the core and was thus effec-
tively only subjected to one-fourth of the flux, and partly because many
of the neutrons had energies above the large Xe!3% absorption resonance.
As little as 59 of this computed poisoning would have been detectable,
but none was found.

There was a small leakage from the gas volume above the liquid surface
of the fuel pumps which made operation at a high power level somewhat
awkward, but danger to operating personnel was circumvented by operat-
ing with the reactor pit at a subatmospheric pressure and remotely ex-
hausting the pit gases to the atmosphere at a location where they were
adequately dispersed.

The entire program of experiments that had been planned for the reactor
was completed satisfactorily. The reactor was shut down after a total
power production of 96 Mwh, and it was later dismantled. The fuel and
sodium systems had been in operation for a total of 462 and 635 hr, re-
spectively, including 221 hr of nuclear operation, with the final 74 hr of
operation in the megawatt range.
CHAPTER 17
CONCEPTUAL DESIGN OF A POWER REACTOR*

The design of a homogeneous molten-salt reactor of the type discussed
in the preceding chapters is described below. The choice of the power
level for this design is arbitrary, since the 8-ft-diameter reactor core, chosen
from nuclear considerations, is capable of operating at power levels up to
1900 Mw (thermal) without excessive power densities in the core. An
electrical generator of 275-Mw capacity was chosen, since this is in the
size range that a number of power companies have used in recent years.
It is estimated that about 6% of the power would be used in the station,
and thus the net power to the system would be about 260 Mw.

Two sodium circuits in series were chosen as the heat-transfer system
between the fuel salt and the steam. Delayed neutrons from the circulating
fuel will activate the primary heat exchangers and the sodium passing
through them. A secondary heat-exchanger system in which the heat will
transfer from the radioactive sodium to nonradioactive sodium will serve
to prevent radioactivity at the steam generators, superheaters, and re-
heaters. The fuel flow from the core is distributed among four primary
heat exchangers which serve as the first elements of the four parallel paths
for heat transfer to the steam. A single primary heat exchanger and path
1s provided for the blanket circuit.

Plan and elevation views of the reactor plant are shown in Figs. 17-1
and 17-2, and an isometric drawing showing the piping of the heat-transfer
systems is shown in Fig. 17-3. The reactor and the primary heat exchangers
are contained in a large rectangular reactor cell, sealed to contain any
leakage of fission-product gases. All operations in the cell must be carried
out remotely after the reactor has operated at power. The principal char-
acteristics of the plant are listed in Table 17-1. |

17-1. FUEL AND BLANKET SYSTEMS

17-1.1 Reactor vessel. The reactor vessel and the fuel and blanket
pumps are a closely coupled assembly (Fig. 17-4) which is suspended.
from a flange on the fuel pump barrel. The vessel itself has two regions—
one for the fuel and one for the blanket salt. The fuel region consists of
the reactor core surmounted by an expansion chamber, which contains
the single fuel pump. The blanket region completely surrounds the fuel
region, and the blanket salt cools the walls of the expansion chamber gas
space and shields the pump motor. The floor of the expansion chamber is

*By L. G. Alexander, B. W. Kinyon, M. E. Lackey, H. G. MacPherson, L. A.
Mann, J. T. Roberts, F. C. VonderLage, G. D. Whitman, and J. Zasler.

681
682 CONCEPTUAL DESIGN OF A POWER REACTOR [cHAP. 17

Primary Sodium

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Pump (1 of 5)
Reactor .
Intermediate HeITt Primary Sodium To Secondary
Exchanger Ce Sodium Heat Exchanger (1 of 5)
Blanket Pump Primary Secondary
Air Lock Shield Secondary Sodium Circvuits
Shield
Fuel Pump Boiler
Hot Maintenance Area\ (1 of 5)
A T XA p
3
Maintenance :,
T .4_:5\:?0 . T Turbo- ¢
Heater | y Generator
' Removal '
Fuel Drain-g .
| Tank b el Fd Rl | [ ree—= | spe—Oflcl
L
L +
’ ; | (lr‘. R h
Chemical " eheater ¢
,  Processing : (1 of 4)
Control /s
S 2 Blanket Superheater
Fuel Chemicol EnriCher B|Onkef H.eCIf. (] Of 5)
Processing Transfer Circuit
/ F.Ud Expansion Secondary Sodium
Blanket Drain Enricher Pump (1 of 5)
Blanket  BlanketTo-Sodium
Chemical  Heat Exchanger .
Processing Fuel-To-Sodium

Heat Exchanger (1 of 4)

Fia. 17-1. Plan view of molten salt power reactor plant.

a flat disk, 3/8 in. thick, which serves as a diaphragm to absorb differential
thermal expansion between the core and the outer shells.

17-1.2 Fuel pump. The fuel pump is of the type illustrated in Chap-
ter 15 (Fig. 15-3) and is designed to have a capacity of 24,000 gpm. It is
driven by a 1000-hp motor with a shaft speed of 700 rpm. This pump
incorporates three major advanced features that are being developed, but
which are not present in any molten-salt pump operated to date. These
are a hydrostatic lower bearing to be operated in the molten salt, a laby-
rinth type of gas seal to prevent escape of fission-product gases up the
shaft, and a hemispherical gas-cushioned upper bearing to act as a com-
bined thrust and radial bearing. These advanced features are intended
to provide a pump with greater resistance to radiation damage and less
complex auxiliary equipment than necessary for pumps presently used
for molten salts.

17-1.3 System for removal of fission-product gases. About 3.5%, of
the fuel passing through the fuel pump is diverted from the main stream,
17-1] FUEL AND BLANKET SYSTEMS 683

Primary Sodium
To Secondary Sodium
Heat Exchanger
(1 of 5)
Fuel To Sodium
Heat Exchanger
(1 of 4)

  
       

 
   

Jet Pump Superheater

  

  
 
 
 
 

   
  
  

 
    
 

 
   
 

Fuel crmary (1 of 5) (1 of 5)
Manipulator Scm '
Pump Removable Boiler (1 of 5)

Air Lock (1 of 5)

Hot Maintenance Area

Concrete Slabs

       
 
  

Turbine-Generator

O

     
    
   

Maintenance Area

 
 
   
 
    
 
 
 
 

Heater
Removal

Steam
Header

Fuel
Tank

   

    
    
   

Blanket To Sodium

   

Blanket Pump

 

  

Heat Exchanger S:Z?fr:dP%Zp (1 of 3)
(1 of 5) Boiler Feed
Water Pump
Primary Sodium Secondary Sodium (1 of 5)
Drain Tank Drain Tank
(1 of 5) (1 of 5)

Fia. 17-2. Elevation view of molten salt power reactor plant.

 

 

 

 

 

 

F1c. 17-3. Isometric view of molten salt power reactor plant.
Section A-A

Siphon Drain

 

 

 

Fuel Pump
Motor

 

 

 

 

 

 

Fuel Line To

 

Heat Exchanger <

 

 

0123 435

Scale—Feet

Fuel Return

i

F1g. 17-4. Reactor vessel and pump assembly.

 

 

 

/ /

Molten Salt

NN

 

~

     
 

 

 

Blanket
Pump
Motor

 

 

Blanket Expansion
Tank

 

 

 

 

Wil

 

 

 

 

Blanket
Return

. Fuel Expansion

Tank

Breeding
Blanket
17-1] FUEL AND BLANKET SYSTEMS 685

TABLE 17-1

REACTOR PLANT CHARACTERISTICS

Fuel
Fuel carrier

Neutron energy
Moderator
Primary coolant

Power
Electric (net)
Heat
Regeneration ratio
Clean
Average (20 yr)
Blanket salt

Refueling cycle at full power
Shielding
Control
Plant efficiency
Exit fuel temperature
Steam
Temperature
Pressure
Second loop fluid
Third loop fluid
Structural materials -
Fuel circuit
Secondary loop
Tertiary loop
Steam boiler
Steam superheater
Active-core dimensions
Fuel equivalent diameter
Blanket thickness
Temperature coefficient, (Ak/k)/°F
Specific power
Power density
Fuel inventory
Initial (clean)
Average (20 yr)
Clean critical mass
Burnup

> 909, U235F,
62 mole 9, LiF, 37 mole 9, BeFs,
1 mole 9, ThF,
Intermediate
LiF Ber
Circulating fuel solution,
23,800 gpm

260 Mw
640 Mw

0.63

0.50

71 mole 9, LiF, 16 mole 9, BeF,
13 mole 9, ThFy

Semicontinuous

Concrete room walls, 9 ft thick

Temperature and fuel concentration

44.39,

1210°F at approximately 83 psia

1000°F, with 1000°F reheat
1800 psia,

Sodium

Sodium

INOR-8

Type-316 stainless steel
5%, Cr, 19, Si steel
2.59%, Cr, 19, Mo steel
5% Cr, 19, Si steel

8 ft

2 ft

—(3.84+0.04) X 10°5
1000 kw/kg

80 kw/liter

604 kg of U235
1000 kg of U235
267 kg of U235
Unlimited
686

CONCEPTUAL DESIGN OF A POWER REACTOR

[cHAP. 17

 

 

 

  

 

 

 

 

 

 

 
 

 

 

 

 

 

 

 

 

 

 

 

 

Y
~—~Blanket
Pump A
Blanket Expansion SCFM
) Tank — 0.1 CFS
Fuel Expansion 1210% Blanket Bypass
Tank 1.8 CFS 0.1CFS
1250°F Fuel Bypass i anke Y
< - A
235 F13 785 Ft3 61 Ft3 16 F3
v 235 Ft3 785 Ft3 61 F13 16 Ft3
738%F 212°F 62°F —40°F
Av Av ~Av Av
Y
y i
Coolant Coolant Coolant Coolant

Fic. 17-5. Schematic flow diagram for continuous removal of fission-product gases.

mixed with helium from the pump-shaft labyrinth seal, and sprayed into
the reactor expansion tank. The mixing and spraying provides a large
fuel-to-purge-gas interface, which promotes the. establishment of low
equilibrium fission gas concentrations in the fuel. The expansion tank
provides a liquid surface area of approximately 26 ft? for removal of the
entrained purge and fission gas mixture. The gas removal is effected by
the balance between the difference in the density of the fuel and the gas
bubbles and the drag of the opposing fuel velocity. The downward surface
velocity in the expansion tank is less than 1 in/sec, which should allow all
bubbles larger than 0.008 in. in radius to come to the surface and escape.
In the Aircraft Reactor Experiment at least 97% of the fission-product
gases were continuously purged by similar techniques.

With a fuel purge gas rate of 5cfm, approximately 350 kw of beta
heating from the decay of the fission-product gases and their daughters
is deposited in the fuel and on metal surfaces of the fuel expansion tank.
This heat is partly removed by the bypass fuel circuits and the balance is
transferred through the expansion tank walls to the blanket salt.

The mixture of fission-product gases, decay products, and purge helium
leaves the expansion tank through the off-gas line, which is located in the
top of the tank, and joins with a similar stream from the blanket expansion
tank (see Fig. 17-5). The combined flow is delayed approximately 50 min
in a cooled volume to allow a large fraction of the shorter-lived fission
products to decay before entering the cooled activated-carbon beds. The
17-2] HEAT-TRANSFER CIRCUITS AND TURBINE GENERATOR 687

capacity of the carbon beds will hold krypton from passing through for
approximately 6 days, and xenon for much longer times.

The purge gases, essentially free from activity, leave the carbon beds to
join the gases from the gas-lubricated bearings of the pumps. The gases
are then compressed and returned to the reactor to repeat the cycle. Ap-
proximately every four days the gas stream is diverted from one set of
carbon beds to the other. The inactive bed is then regenerated by warming
it to expel the Kr3% and other long-lived fission products. It will probably
be economical to recover some of these gases; others may be expelled to
the stack.

17-2. HeAT-TRANSFER CirRcuIiTs AND TURBINE (GENERATOR

The primary heat exchangers are designed to have the fuel on the shell
side and sodium inside the tubes. This arrangement makes full use of the
superior properties of sodium as a heat-transfer fluid and appears to yield
the lowest fuel volume.

The heat exchangers, which are of semicircular construction, as shown
in Fig. 17-3, provide convenient piping to the top and bottom of the
reactor. The thermal characteristics of the primary heat exchanger, to-
gether with the characteristics of other heat exchangers of the reactor
system, are listed in Table 17-2.

The sodium in the intermediate heat-transfer system (see Fig. 17-6) 1s
heated by the fuel in the primary heat exchanger and is pumped out of the
reactor cell and through the reactor cell shield to adjacent cells, which con-
tain the secondary sodium-to-sodium heat exchangers and the pump. No
control of intermediate sodium flow is required, so there are no valves and
a constant speed centrifugal pump is used. To permit the sodium to be at
a lower pressure than the fuel in the primary heat exchanger, the pump for
the intermediate sodium is in the higher temperature side of the circuit.
The secondary heat exchangers are of the U-tube in U-shell, counterflow
design, with the intermediate sodium in the tubes and the final sodium on
the shell side.

The final sodium circuit, except for the secondary exchanger, is outside
the shielded area and thus available for adjustment and maintenance at all
times. The principal problems in this circuit are concerned with the ad-
justment of sodium temperature. Excessive thermal strains are prevented
in the steam generator by limiting the temperature of the sodium entering
it, and in the intermediate heat exchanger by the regulation of sodium flows
so that too cold sodium is never returned to it. The hot sodium from the
secondary exchanger is split into three streams with regulating valves for
control of the relative flows. One stream bypasses the steam system and
goes directly to a blender; the flow in it is, of course, greatest at low power
CONCEPTUAL DESIGN OF A POWER REACTOR [cHAP. 17

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690 CONCEPTUAL DESIGN OF A POWER REACTOR [cHAP. 17

   

Low- and

   
  
  
  

 
 

Boiler

 

Turbines

Turbine

  
 
 
 
 
 
 

urbine Stop

Relief
Water

Station
Pump

   

 

Sodium

Superheater In Series

 
  
   
 
 

Steamchest

Pump

Feedwater
in Series

Fig. 17-6. Schematic diagram of heat-transfer system.

levels. The other two streams go to the superheater and the reheater, and
are then combined with the bypass flow in the blender. On leaving the
blender, the sodium stream is split again by a three-way valve into two
streams; one enters a second bypass and goes directly to the main pump
and the other enters a jet pump that keeps a large sodium flow recirculating
through the boiler, which is of the Lewis type. The three-way valve is
adjusted, at design point, so that about two-thirds of the flow goes to the
jet pump and one-third bypasses the boiler. At low power levels the valve
would be adjusted to give very low flows to the boiler.

The centrifugal pump in this circuit has two speeds, full speed and one-
fourth of full speed. The low-speed operation provides for better regulation
of the sodium flow at very low power levels.

The turbine selected uses 1800-psia steam at 1000°F with reheat to
1000°F and is rated at 275 Mw. It is a 3600-rpm single-shaft machine with
three exhaust ends. The turbine heat rate is estimated to be 7700 Btu/ kwh,
or 44.39%, cycle efficiency, while 7860 and 8360 Btu/kwh are the generator
and station heat rates, respectively. With 69, of generator output used
for station auxiliaries, 260 Mw is supplied to the bus bar.

17-3. REMOTE MAINTENANCE PROVISIONS

Remotely controlled mechanized tools and viewing devices are provided
in the reactor cell for making minor repairs and for removing and re-
17-4] MOLTEN-SALT TRANSFER EQUIPMENT 691

placing any component in the cell. The tools will be able to handle any
pump, heat exchanger, pipe, heater for pipe and equipment, instrument, and
even the reactor vessel, and, correspondingly, the components will be de-
signed and located for accessibility and separation.

The removal and replacement of components requires a reliable method
of making and breaking joints in the pipe. Cutting and welding of pipe
sections can be used, but in the low-pressure molten-salt system 1t 1s be-
lieved that a flanged-pipe joint (see Section 15-5) may be satisfactory.

All equipment and pipe joints in the reactor cell are laid out so that they
are accessible from above. Directly above the equipment is a traveling
bridge on which can be mounted one or more remotely operated manipu-
lators. At the top of the cell is another traveling bridge for a remotely
operated crane. At one end of the cell is an air lock that connects with the
maintenance area. The erane can move from the bridge in the cell to a
monorail in the air lock.

Closed-circuit television equipment is provided for viewing the mainte-
nance operation in the cell. A number of cameras are mounted to show the
operation from different angles, and a periscope gives a direct view of the
entire cell. |

17—4. MoOLTEN-SALT TRANSFER L QUIPMENT

The fuel-transfer systems are shown schematically in Fig. 17-7. Salt
freeze valves (see Section 15-3) are used to isolate the individual com-
ponents in the fuel-transfer lines and to isolate the chemical plant from
the components in the reactor cell. With the exception of the reactor
draining operation, which is described below, the liquid is transferred frqm
one vessel to another by a differential gas pressure. By this means, fuel
may be added to, or withdrawn from, the reactor during power operation.

The fuel added to the reactor will have a high concentration of UF4 with
respect to the process fuel, so that additions to overcome burnup will re-
quire transfer of only a small volume; similarly, thorium-bearing molten
salt may be added at any time to the fuel system. The thorium, in addi-
tion to being a design constituent of the fuel salt, may be added in amounts
required to serve as a nuclear poison.

For the main fuel drain circuit, bellows-sealed, mechamcally operated,
poppet valves (see Section 15-3) will be placed in series with the freeze
valves to establish a stagnant liquid suitable for freezing. Normally these
mechanical valves will be left open. By melting the plug in the freeze line
and opening gas-equalization valves, the liquid in the reactor will flow by
gravity to the drain tank, and the gas in the drain tank will be transferred
to the reactor system. Thus gas will not have to be added to, or vented
from, the primary system.
[cHAP. 17

CONCEPTUAL DESIGN OF A POWER REACTOR

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17-6] CHEMICAL REPROCESSING METHOD 693

17-5. FueL DRraIN TANK

For the drain vessel design calculations, it was assumed that at 1200°F
the fuel system volume would be 600 ft3. The design capacity of the
drain vessel was therefore set at 750 ft3 in order to allow for temperature
excursions and for a residual inventory. An array of 12-in.-diameter pipes
was selected as the primary containment vessel of the drain system in
order to obtain a large surface area-to-volume ratio for heat-transfer effi-
ciency and to provide a large amount of nuclear poison material. Forty-
eight 20-ft lengths of pipe are arranged in six vertical banks connected on
alternate ends with mitered joints (see Fig. 17-8). The six banks of pipe
are connected at the bottom with a common drain line that connects with
the fuel system. The drain system is preheated and maintained at the de-
sired temperature with electric heaters installed in small-diameter pipes
located axially inside main pipes. These bayonet-type heaters can be re-
moved or installed from one face of the pipe array to facilitate maintenance.
The entire system is installed in an insulated room or furnace to minimize
heat losses.

The removal of the fuel afterheat is accomplished by filling boiler tubes
installed between 12-in.-diameter fuel-containing pipes with water from
headers that are normally filled. The boiler tubes will normally be dry and
at the ambient temperature of about 950°F. Cooling will be accomplished
by slowly flooding or “‘quenching” the tubes, which then furnish a heat
sink for radiant heat transfer between the fuel-containing pipes and the
low-pressure, low-temperature boiler tubes. For the peak afterheat load,
a flow of about 150 gpm of water is required to supply the boiler tubes.

The design criteria for this fill-and-drain system are that it always be in
a standby condition immediately available for drainage of the fuel, that 1t
be capable of adequately handling the fuel afterheat, and that it provide
double containment for the fuel. The heat-removal scheme is essentially
self-regulating in that the amount of heat removed is determined by the
radiant exchange between the vessels and water wall. Both the water and
the fuel systems are at low pressure, and a double failure would be required
for the two fluids to be mixed. The drain system cell may be easily enclosed
and sealed from the atmosphere because there are no large gas-cooling
ducts or other major external systems connected to it. |

17-6. CHEMICAL REPROCESSING METHOD

In this plant it is assumed that the core and blanket salts will be re-
processed by the fluoride volatility process [1] to remove UFs. The UFg
will be reduced by a fluorine-hydrogen flame process [2] and returned to
‘the reactor core. The blanket salt with its U233 removed will be returned
to the blanket, since the buildup of fission products in the blanket salt will
694 CONCEPTUAL DESIGN OF A POWER REACTOR . [cHAP. 17

 

 

 

 

 

 

 

 

 

 

 

 

 

Fia. 17-8. Drain and storage tank for fuel salt of molten salt power reactor.

not be an appreciable poison for many years. For purposes of the cost
study, it is assumed that the spent fuel salt with its contained fission prod-
ucts will be discarded after each processing, and new fuel salt will be pro-
vided. Since several means of recovering the fuel salt appear to be feasible,
this is a conservative assumption. The chemical processing is assumed to
be an integral part of this reactor plant.

17-7. CosT ESTIMATES

All the costs were calculated for a 260-Mw net-electrical-output plant
operated at a load factor of 0.80. An annual charge of 149 was made for
all capital items, other than the uranium inventory, for which a 4% annual
charge was made. It is assumed that the reactor plant components will
have been engineered and developed with separate funds not included in
this estimate.
17-7] COST ESTIMATES

695

Table 17-3 gives the fuel cycle costs separately from the other reactor
costs, so that a comparison can be made with comparable costs for solid-

fuel-element reactors.
TABLE 17-3

FueL .CycLeE Costs

Operating costs (fuel cycle only)

U235 consumed

Fuel salt makeup (or recovery cost)

Chemical plant operation (listed as part of Operation
and Maintenance in Table 17-5)

Caprtal charges (fuel cycle only)

U233 and U235 inventory, 1000 kg at 49,

Chemical plant and equipment $3,000,000 at 149,
(included in capital costs in Table 174 and fixed
cost in Table 17-5)

Total fuel cycle cost

$/year

2,260,000
760,000
500,000

3,520,000

680,000
420,000

1,100,000

4,620,000

mills/kwh

1.24
0.42
0.28

Do
ot
N

|

An estimate of the capital costs for the reactor plant is given in Table
17—4. This breakdown leads to the estimated power costs in Table 17-5.

TABLE 174

CaritaL CosTts

Land and land rights

Structures and improvements

Reactor system (including chemical plant)

Steam system

Turbine-generator plant

Accessory electrical equipment

Miscellaneous power plant equipment
Direct costs subtotal
Contingency subtotal

General expense

Design costs

Total cost

$500,000
7,500,000
20,039,000
3,750,000
11,750,000
4 600,000
1,250,000
49,389,000
10,216,000
7,500,000
2,450,000

$69,555,000
696 CONCEPTUAL DESIGN OF A POWER REACTOR [cHAP. 17

Note that in Table 17-5 the chemical plant capital and operating costs
listed in Table 17-3 are part of the fixed cost and operation and mainte-
nance, respectively, so that the amount listed for fuel charges is only 2.03

TABLE 17-5

PowEeEr CosTs

Item Annual charge malls/kwh
Fixed cost $9,738,000 5.34
Operation and maintenance 2,700,000 1.48
Fuel charges 3,700,000 2.03
Total annual charge $16,138,000

Total power cost 8.85

mills/kwh. This breakdown is in keeping with the philosophy of regarding
the chemical reprocessing as an integral part of the reactor plant. The
difference in cost between having the reactor on standby and having it on
the line is less than 2 mills/kwh.

REFERENCES

1. G. I. CatuErs, Uranium Recovery for Spent Fuel by Dissolution in Fused
Salt and Fluorination, Nuclear Sct. and Eng. 2, 767 (1957).

2. S. H. SmiLey and D. C. BraTter, Conversion of Uranium Hexafluoride to
Uranium Tetrafluoride, Progress in Nuclear Energy, Series III, Process Chem-
istry, Vol. II. New York: Pergamon Press, in preparation.
697

BiBLioGRAPHY FOR PART II

Conceptual design studies of molten-salt power reactors:

BuLMER, J. J. et al.,, Fused Salt Fast Breeder, USALEC Report CF-56-8-
204(Del.), Oak Ridge National Laboratory, 1956.

Davipson, J. K. and W. L. RosB, A Molten-salt Thorium Converter for Power
Production, USAEC Report KAPL-M-JKD-10, Knolls Atomic Power Labora-
tory, 1956.

Davies, R. W. et al., 600 Mw Fused Salt Homogeneous Reactor Power Plant,
USAEC Report CF-56-8-208(Del.), Oak Ridge National Laboratory, 1956.

WEHMEYER, D. B. et al., Study of a Fused Salt Breeder Reactor for Power
Production, USAEC Report CF-53-10-25, Oak Ridge National Laboratory, 1993.