FFR_chap13.txt

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. Ivaluation 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 dynamie
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 containing 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 1s accompanied by the formation of
subsurfuce 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
molvbdenum, shows excellent compatibility with fluoride salts at tempera-
tures i 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 reaxctor construction. The
approximate compositions of the three alloys, Inconel, Hastelloy B, aund
INOR-S, 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.

-

395
 

 

 

 

 

 

 

 

 

596 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [cHAP. 13
TasLE 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)
Siticon 0.5 0.35 (max) 1.0 (max)
Sulfur 0.01 0.01 (max) (.03 (max)
Copper 0.5 0.35 (max)
Cobalt 0.2 (max) 2.5 (max)
Nickel 72 (min) Balance Balance
g
?‘ —E__t Hot-Leg

T

‘ [: - Samples

o

—i

C|um-‘ t‘ -t

shell }

Heaters — ;|

 

Fia. 13-1. Diagram of a standard thermal-convection loop, showing locations

 

 

 

    

—_— -

F: Cold-Leg

Samples

 

at which metallographic sections are taken after operation.

 
13~1] SURVEY OF SUITABLE MATERIALS 597

   

* s 7
W 5

i T

P // Cold Section

e
)
o7
& e‘f’//
- -1 K - Q‘ ,c"l"/ l

S -DTIIIIITES- I“/f
Heated Sections

LT T L Ty e

X

 

F1a. 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 commereial 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 strue-
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
508 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS — [cHAP. 13

graphite now heing 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°L", DMore tests are needed to finally establish the com-
patihility of graphite-salt-alloy systems.

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

13-2. CorrosioN or NICKEL-BAsE ArrLovs 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 arc heated, usually with clamshell heaters,
convection forces in the eontained 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.-I1D tubing, and somewhat higher
values can be obtained with smaller tubing.

13-2.2 Mechanism of corrosion. Most of the data on corrosion have
heen 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 1s depleted in chromium, chromium from the interior diffuses down
the concentration gradient toward the surface. Sinee diffusion occurs by
a vacancy process and in this particular situation is essentially monodirec-
tional, 1t 1s 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. Ixaminations 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 599

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:
Felks + Cr === CrFo + Fe. (13-1)
2. Oxide films on the metal surface:
2Fes03 + 3CrFs === 3Cr0O, 4 4Fel';. (13-2)
3. Clonstituents of the fuel:
Cr+ 2UF, == 2UF; 4 Crla. (13-3)

The ferrie 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-conveetion 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
cant 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 UTFy constituent of the salt [reaction (13-3)] to establish a quasi-
equilibrium amount of CrF» in the salt. At this point attack proceeds
linearlv with time and oceurs 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 UKy (Eq. 13-3). If nickel and iron are
considered inert diluents for chromium in Inconel, the process can be
simply deseribed. Under rapid circulation, a uniform concentration of UF4,
UF 3. and CrFs is maintained throughout the fluid; the concentrations must
satisfv the equilibrium constant

. g Yo, - Yiur, Nor. - N2ur,
K,=HK, Ky= ‘ C s 13-4
v AN Yor - Yiur,  Nor - N2ur, ( )

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

 

*q-’ ‘\ fi" . é' . . > A
i . u . L
- - -+
- . L ! . . " to .
Fi1a. 13-3. Hot-leg section from an Inconel thermal-convection loop which cir-

culated the fuel mixture NaF-ZrF;-UF, (50-46-4 mole %) 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 ecrystals in the
cold zone. In other melts, for example Nal-ZrF4+—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 UI'3
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

 

 

4

F1c. 13-4. Hot-leg section of Inconel thermal-convection loop which circulated
the fuel mixture NaF-ZrF4-UF4 (55.3-40.7-4 mole %) for 1000 hr at 1250°F.
200X

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 employved is small, and (2) the experimental and analytical difficul-
tie~ are considerable. Addition of more than the equilibrium quantity of
U1, may lead to deposition of some uranium metal in the equipment walls
through the reaction

4UF, == 3UF4 4 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 UFs and CrFy 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
<0 long as the same maximum metal-salt interface temperature is present
in both tvpes of loop. The results of many tests with both types of loop
are ~ummarized in Table 13-2 without distinguishing between the two
tvpes 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

TasLe 13-2

SuMMarY oF CorrosioN DaTa OBTaINgD 1IN THERMAL-CONVECTION AND
Forcep-CircuLaTioN Loop TEsTS oF INCONEL AND INOR-8
ExrosEp 10 Vartous CIRCULATING SaLT MIXTURES

 

 

 

 

s alt Ti ; Depth of subsurface
Constituents of UF, or ThF, Loop ‘t dxnmi? st m'li.(:) void formation at
base salts content material -empc:; ure, ()per}.llrmn, i hottest part of loop,
; .
Nal™-Zrly 1 mole €5 Ul Inconel 1250 1000 <0001
1 mole &, Ul‘y Inconel 1270 6300 000025
4 mole 7, Uly Inconel 1250 1000 ‘ (+.002
4 mole ¢, ULy Inconel 1500 1000 0.007-0.010
4 mole ¥, UF, INOR-8 1500 1000 0.002-0.003
0 Inconel 1500 1000 0.002-0.003
NaF-BeF. 1 mole %, Ul Inconel 1250 1000 0.001
0 [nconel 1500 a0o 0.004-0 010
3 mole 95, Uy Inconel 1500 00 ¢.008-0.014
1 mole %, UF, INOR-8 1250 6300 0.001
LiF-BeF; 1 mole 9 UFy Inconel 1250 1000 0.001-0 002
3 mole 9 UI', Ineonel 1500 00 0.012-6.020
1 mole €7 UF, INOR-8 1250 1000 1}
NaF LiF-BeF. 0 Inconet 1125 1000 0.002
{) Inconel 1500 500 0.003-0.005
3 mole ¢ UFy Tnconel 1500 500 0 008-0.013
NaF LiF-KF 0 Inconel 1125 1000 ‘ 0.001
2.5 mole 3 UF, Inconel 1500 500 0. 017
0 INOR-8 1250 1340 0
2.5 mole 7, UFy INOR 8 1500 1000 0.001-0.003
Lil 29 mole 7, ThF, Inconel 1250 1000 0-0.0015
NaF-Bel', 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 BeFs are somewhat more corrosive than those containing
ZrFy4, 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

 

»

&

Fic. 13-5. Hot-leg section of Inconel thermal-convection loop which circulated
the fuel mixture LiF-BeF2-UF4 (62-37-1 mole ) for 1000 hr at 1250°F. (250%)

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 Nak-
Z1F ULy 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 is extremely
dependent on the maximum system operating temperature. The results of
604 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS [crAP. 13

numerous tests indicate that the nickel-base alloys, such as Inconel and
INOR-8, are satisfactory containers for sodium at temperatures below
1300°F, and that above 1300°F the austenitic stainless steels are preferable.

13-3. FaBricaTion oF 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
is 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 0.019,. Carbon present in excess of this amount
precipitates as discrete particles of (N1,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)
wo— T T T T T T T T 1'%
200 |— 80

    
  
    
 

Ultimate Tensile Strength

150 60

0.2% Offset Yield Strength

Strength {psi)

Elongation

50 20

Per Cent Elongation (2-in. Gage Length}

 

 

gL T
0 10 20 30 40 50 60 70 80 90

Reduction in thickness (%)

 

F1g. 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.

Fxtensive 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 0.049, 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

 

 

STAINLESS STEE

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

WELD DEPOSIT —-

   

 

 

 

 

 

= 12 in.
o
_____ //
_____ ;o
i 1
i/l /2 in
L7 P / / \ 71
I

N
BACKING STRIP — " 4o-in-THICK PLATE

Fi1g. 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-Ib vacuum-melted ingots with nominal boron contents of up to 0.109,
were prepared, slotted, and welded as shown in Fig. 13-9. All ingots with
0.0297 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

i

Hydraulic Cylinder \

     
   
   
  
 

Specimen
Thermocouple

—— Bend

Quartz Red Specimen

 

Control
Thermocouple

 

 

 

Heating Element—/
Connections
(220v or 110v)

Quartz Mirror

1012345

Scale in Inches

Fra. 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-UF4 in
tests conducted at 1500°F for 100 hr. Further, two precious metal-base
brazing alloys, 829, Au-18%, Ni and 809, Au-209, Cu, were unattacked
in the LiF-KF-NaF-UF4 salt after 2000 hr at 1200°F. These two precious
TaBLE 13-3

Resurrs oF Sipe-BenND TEsts oF As-WELDED
INOR-8 AND INCONEL SAMPLES

 

 

 

 

 

 

Filler metal
Test ,
¢ INOR-8 (Heat M—5) INOR-8 (Heat SP-19) Inconel
emperature,
O
F
Bend angle,* Elongationt in Bend angle, Elongation in Bend angle, Elongation in

deg 1/4in., 9%, deg 1/4 in., % deg 1/4 in., 9
Room >90 >40 > 90 >40 >90 > 40
1100 > 90 > 40 > 90 > 40 >90 >40
1200 > 90 >40 > 90 >40 >90 >40
1300 >90 > 40 30 15 >90 >40

> 90 >40 15 8 >90 >40
1500 >90 > 40 15 8 15 8

> 90 >40 15 8

 

 

 

 

 

 

 

 

*Bend angle recorded is that at which first crack appeared.

tElongation recorded is that at outer fiber at time first crack appeared.

[e-e1

8-HONI J0 NOILVOIHEVA

609
610 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [cHAP. 13

TaBLE 13-4

Nickrer-Base Brazing ALLoys ror USE IN
HeaT 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

 

 

 

 

 

 

metal alloys were also tested in the LiF-BeFo—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. Iiddy-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 2 high-intensity radiation field, such as that which will be present if
maintenance work 1s 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 5%,
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-8 611

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 H 5
Welds 14 14

 

 

 

 

 

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

13-4, MEcHANICAL AND THERMAL PropreErTIES OoF 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 Iig. 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°I° 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 is 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
1.0, total strain, as a function of stress, in the data shown in Iig. 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-
Jation between the creep curves is obtained at the lower stress values.
Some seatter in time to rupture occurs at 25,000 psi, a stress which corre-
sponds to the 0.2, offset yield strength at this temperature. Such scatte
is to be expected at this high stress level.
612 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS [cHAP. 13

 

33,000 | ] | A
30,000 — / / ]
25,000 / —

20,000 — —

Stress (psi)

15,000 |— / ]

10,000 — / —

5,000 — / |

 

 

 

0 ! | L L

0.0006 00012 0.0018 0.0024 0.0030 0.0036 0.0042
Elongation {in./in.)

Fra. 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.

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.

 

 

 

. Tensile strength at
Material 1300°F, psi
18-8 stainless steel 40,000
Cr-Mo steel (59 Cr) 20,000
Hastelioy B 70,000
Hastelloy C 100,000
Inconel 60,000
INOR-8 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

 

     
 
   
 

 

 

 

 

(x10%)
110 2
I I l l | I 1 l -
Tensile Strength
100 -4 8
@0 g
80 —g
5 70 —Heo .
a Elengation, % ~ =
@ c
g 3
@ 50 —3 A
50 — ]
40 — 9
Yield Strength
30 — —g
20 | | | { l | [ 1 &
0 2 4 6 8 0 12 14 16 18(x102j

Temperature, °F

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

{x108)
34

 

32 —

30 —

28 +—

Young's Medulus {psi)

 

22

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

 

 

 

Temperature (°F)

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

(x'|03)

28 T T TTT1T0 T T 1T/ T TTTT77

0.2% Constant Strain
24 |— W"'

 

20 —

0.1% Constant Strain  —

=
|

0.05% Constant Strain

Stress (psi)

|
l

 

 

b L Y [ L LIl
0.1 0.2 0.5 1.0 2 5 10 20 50 100
Time (hr}

 

F16. 13-14. Relaxation of INOR-8 at 1300°F at various constant strains.

(x103)
28 1 T T T TTT] T 7T T T T T T11

 

24 —  0.1% Constant Strain —

20

—_
o

    
 

0.05% Censtant Strain

Stress (psi)

I

—_— N

* Discontinued Test

 

 

0 | L bl | LI | |
1 2 5 10 20 50 100 200 500 1000
Time (hr}

 

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

 

   
  

 

  

 

 

 

 

    
    
    

 

 

  

 

103 m
rrrimm brrrnnt T TTTEAT T TTTIhn T TTTT13
5 /1100"F _
2 — —_-\\‘:
104 L / =~ —]
= . =
- — 1300°F ~ —]
= o T~
£ 21 ~o
. 5 Extrapolated \\:\_
E 3
| 1 yr 10 yr ]
i Ll
102 e e b L bl
] 10 100 1000 10,000 100,000
Time (hr) to 1% Strain
Fig. 13-16. Creep data for INOR-S8.
100 = T T T T 11171 T T T ]
50— -
- Results for Three Samples —]
Stressed at 25,000 psi
20— —
_E; 10— 1
2 — _]
0.5 :
- Results for Two Samples N
- Stressed at 20,000 psi 1
0.2 | —
o R Ll
10 20 50 100 200 500 1000 2000

Time thr}

Fia. 13-17. Creep-rupture data for INOR-8.
616 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [cHAP, 13

 

i T T TTTI T T TTTTT T TTTTTH

Stress {psi)

 

 

 

 

102 oLl L LTl L 1 1 IIt]

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

Fic. 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 is 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
gystem, but no brittle phase exists if the alloy contuins less than 209, Mo,
8% 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 (N1, Mo)sC. Studies of the effect
of the carbides on ereep 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.
13-4] MECHANICAL AND THERMAL PROPERTIES OF

(x10%)
10

INOR-8 617

 

00
90
80
70

60

Tensile Strength (psil

50 —

40 +—

 

Annealed 1 hr at 2100°F
e == Aged 500 hr at Test Temperature

 

0.2% Yield Point

 

30 [

D S e G = ]

 

20

 

60

 

50

40

Elongation (%}
L
S
[

 

   
  
 

| |

Elongation
—— e —

|

 

 

1100 1200

1300

Test Temperature (°F)

1400

Figc. 13-19. Effect of aging on high-temperature tensile properties of INOR-S8.

TaBLE 13-5

Resvrrs or RooM-TEMPERATURE EMBRITILEMENT TESTS orF INOR-8

 

 

 

 

 

 

Heat treatment Ultimate tensile | Yield point at- Elongation,
‘ - strength, psi 0.29, offset, psi o
Annealed* 114,400 44 700 50
Annealed and aged 500 hr at
1000°F 112,000 42 500 53
Anuealed 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

 

 

 

*().045-in. sheet, annealed 1 hr at 2100°F and tested at a strain rate of 0.05 in/min.
618 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 For INOR-8
AND INCONEL AT SEVERAL TEMPERATURES

 

 

 

 

 

 

 

 

 

Thermal conductivity, Btu/ (ft?)(sec) (°F/ft)
Temperature,
°F
INOR-8 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

TasLE 13-7

CorrriciENT oF LiNnear Exransion or INOR-8
FOR SEVERAL TEMPERATURE RANGES

 

 

 

o Coeflicient of linear expansion,
Temperature range, °F in/(in) C°F)
X106
70-400 5.76
70-600 6.23
70-800 6.58
70-1000 6.89
70-1200 7.34
70-1400 7.61
70-1600 8.10
70-1800 8.32

 

 

 

 
13-5] OXIDATION RESISTANCE 619

13-5. OXIDATION RESISTANCE

The oxidation resistance of nickel-molybdenum alloys depends on the
service temperature, the temperature eycle, 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
senle formed by the oxidation 18 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 cyeles 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
6°¢ Cr.  Alloys containing more than 69, Cr are insensitive to thermal
cyeling and the molybdenum content because the oxide scale is pre-
dominantly stable CryO3. An abrupt decrease, by a factor of about 40, in
the oxidation rate at 1800°F is observed when the chromium content is
inereased from 5.9 to 6.297.

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°I" could be tolerated. IFor 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 1, would ocecur 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

OxmaTioN RATE oF INOR-8 AT Vartious TEMPERATURES®

 

Weight gain, mg/cm?2
Test temperature, S S T8 Shape of

°R rate curve
In 100 hr In 1000 hr

 

 

 

1200 0.00 0.00 Cubic or logarithmic
1600 0.25 0.67t Cubic

1800 0.48 1.5% Parabolic

1900 (.52 2.07 Parabolic
2000 2.70 28.27 Linear

 

 

 

 

 

 

*3.7 mg/em? = 0.001 in. of oxidation.
TIixtrapolated from data obtained after 170 hr at temperature.
620 MOLTEN-SALT REACTOR CONSTRUCTION MATERIALS  [cHAP. 13

 

 

- P

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

13-6. FaBricaTioN oF A DupLEx TusiNe HEAT IEXCHANGER

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 Iig. 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 TFig. 13-21.

The problem of welding INOR-8-stainless steel duplex tubing is being
studied. Ixperiments have indicated that proper sclection of alloy ratios
and weld design will assure welds that will be satistactory 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 inereased in depth along
the grain boundaries with increases in the temperature. The interface of a
duplex sheet held at 1300°} for 500 hr is shown in IMig. 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

 

F1e. 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-
sistunce 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 [cHAP. 13

- . u e . : v v e
- v o DR . P ., . . . T o .
s s s e Unstressed - & e M R
i e ~ - § et . . . R _z . ,: . ‘s .“ " B - -
- - +
. . wiljsinapeiiiioiiis i
SR - : ’ . R INOR-
'0 - - - . * -
. v . o
- - - * ‘
. ! -

AN
S L 3 T
S / a8 -w—-ifl-— Interfcte ce

 

‘ N AN ‘

) . \ : A T o ‘ . Type 316
SN . B o . .

/% : T : e Stainless

N ; LN e N R _ © . Steel

 

. . ‘ - L T . Type 316
C T T g e L . {7 — Ctrinless

T TR o . v
. ";;\ . : e . A . L : % Steel
e o £ « . N " o . o . .
L —" . - i e ~. e
. e e . f . . . {r')

g S *rq,u e @E i e nt Ny ?g«f:g::r;‘ks Rt n-m Interface ~ 02"

 

) (*‘ i . x~
w . . . “ e
- - - + l\‘ - . - . = "
. . - - )
. L ep . . e E e s " wippnegmeii [NOQR -8
- A < e - 2 T
e . AR Stressed oo S
" - ;‘\ "' . * s g - : - ' - ‘. '\:‘ " . :-.' - - 't -
Ve . " = - . s Ey L.

Fra. 13-22. Unstressed and stressed specimens of INOR-8 clad with type-316
stainless steel after 500 hr at 1300°F. Etchant: electrolytic HoS04 (29, solution).
(100%)

 

Fra. 13-23. CCN graphite (A) before and (B) after exposure for 1000 hr to Nal'-
ZrF4~-UF4 (50-46-4 mole 97) at 1300°T" 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. AvarLasiLity oF INOR-8

Two production heats of INOR-8 of 10,000 Ib 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. CompraTIiBILITY OF GRAPHITE WITH MOLTEN SALTS AND
Nicker-Base ArLoys

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.

Aany 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
Nal-7ZrI'y-UF, gave no indication of the decomposition of the fluoride
and no gas evolution so long as the graphite was free from a silicon 1m-
purity.

Longer-time tests of graphite immersed in fluoride salts have shown
greater indications of penctration of the graphite by salts, and 1t must be
assumed that the salt will eventually penctrate the available pores in the
araphite. The “"impermeable” grades of graphite available experimentally
show greatly reduced penetration, and a sample of high-density, bonded,
naturial 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 19 carbon may be picked up in 100 hr. The carbon
pickup appears to be sensitive to temperature, however, inasmuch as
only 0.0259; 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

 
 

“Film of Cr3C2: "

   
 

P
LW
e

 

Fic. 13-24. Cross sections of samples shown in Fig. 13-23. (A) Before exposure;
(B) after exposure. Note the thin film of CrzCs 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 ecarbide, CrzCs. 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°I"). 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-ZrF4—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

TasLE 13-9

ErrecT oF TEMPERATURE 0N CARBURIZATION OF
INncoNeL AnD INOR-8 1v 100 ur

 

 

 

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.