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NUREG/CR-0089 NUREG/CR-0089ANL-78-25ANL-78-25
7 ''
LIGHT-WATER-REACTOR SAFETYRESEARCH PROGRAM:
QUARTERLY PROGRESS REPORT
October-December
UofC-AuA -USDOE
ARGONNE NATIONAL LABORATORY, ARGONNE.
Prepared for the Office of Nuclear Regulatory ResearchU. S. NUCLEAR REGULATORY COMMISSIONunder Interagency Agreement DOE 40-5QffIZggON OF THIS DUCUET i$ UNLIUTE
1977
ILLINOIS
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proved <lid r1 vitu d' il b 1th A K S Ir i t. till.
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U'. S. Nuclear Regulatory ClIui1lssiun
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Available fromNational Technical Inform Citi S Ii
Springfield, Virginia L 161
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TABLE OF CONTENTS
Page
ABSTRAC T . .. ... . .. . ... ..... . ... ... .. .. ... . .. .... .
I. LOSS-OF-COOLANT ACCIDENT RESEARCH: HEAT TRANSFERAND FLUID DYNAMICS . . . . . . . . . . . . . . . . . . . . . . . . . . . .
A. Transient Critical Heat Flux . . . . . . . . . . . . . . . . . . . . . ..
1. Status Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
a. Previous Work . . . . . . . . . . . . . . . . . . . . . . . . . . .b. Future Work . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
B. Reflood Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
References......... .....................................
II. TRANSIENT FUEL RESPONSE AND FISSION-PRODUCT RE-LEASE PROGRAM . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
A. Modeling of Fuel-Fission-product Behavior . . . . . . . . . . . . .
1. Modeling of Fission-gas Behavior during TransientHeating Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . .
a. GRASS--SST Calculation of Transient Gas Releaseduring DEH Test of Saxton Load Follower Fuel . . . . . .
b. Effect of Restraint on Fission-gas Release duringDEH Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
c. Additional Progress . . . . . . . . . . . . . . . . . . . . . . . .
B. Experimental Program. . . . . . . . . . . . . . . . . .
1. Determination of Radial Fission-gas ProfilesDEH-tested Fuel . . . . . . . . . . . . . . . . . . .
2. ANL-EG&G Crosscheck Program. . . . . . . . .a. Summary . . . . . . . . . . . . . . . . . . . . . .b. SEM Examination . . . . . . . . . . . . . . . ..
3. LOCA Simulations. . ........ ... 0..........0.
for
III. MECHANICAL PROPERTIES OF ZIRCALOY CONTAININGOXYGEN. .................................
A. Surnmay.... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
B. Impact Properties of Zircaloy-Oxygen Alloys. ...........
iii
v
1
1
116
6
11
12
13
13
13
1415
16
1619192028
... 0 31
31
31
TABLE OF CONTENTS
Page
C. Integral Tube-burst/Thermal- shock Properties of
Oxidized Zircaloy-4 Cladding . . . . . . . . . . . . . . . . . . . . . . 35
1. Failure of Oxidized Zircaloy-4 Cladding by Thermal
Shock . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 352. Measurements of the Leidenfrost Temperature during
Quench of Oxidized Zircaloy-4 Tubes by Bottom-flooding
with Water . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 363. Oxidation Characteristics of Inner and Outer Surfaces of
Zircaloy-4 Cladding in Steam . . . . . . . . . . . . . . . ..... 39
References........... ..................................... 44
iv
ABSTRACT
This progress report summarizes the Argonne National
Laboratory work performed during July, August, and Sep-tember 1977 on water-reactor-safety problems. The followingresearch and development areas are covered: (1) Loss-of-
coolant Accident Research: Heat Transfer and Fluid Dynamics;
(2) Transient Fuel Response and Fission-product Release Pro-
gram; and (3) Mechanical Properties of Zircaloy ContainingOxygen.
V
1
I. LOSS-OF-COOLANT ACCIDENT RESEARCH:HEAT TRANSFER AND FLUID DYNAMICS
Responsible Section Managers:H. K. Fauske, R. E. Henry, and P. A. Lottes, RAS
A. Transient Critical Heat Flux (J. C. M. Leung and R. E. Henry, RAS)
1. Status Summary
a. Previous Work
Freon blowdown experiments)1-3 performed in a single heatedtube have successfully demonstrated the occurrence of very rapid critical heatflux (CHF), as observed in some water experiments. These tests also showthe strong influence of initial steady-state pressure on the early CHF zone.When the initial pressure is lowered while other system parameters remainthe same, only the subcooled boiling zone and its local wall temperature differfrom the corresponding high-pressure system. However, results indicatemore cooling in the low-pressure tests, with shorter region in the bottom zoneof the heated tube experiencing early CHF (<1 s) during blowdown with flowreversal. The results are summarized in Fig. I.1.
TEST No BREAK TOTAL PRESSURE TEST SECTION LENGTH , mDIAMETER POWER MPo
- BOTTOM TOP
0 05 1.0 1.5 2.0 2.5 2.74
1078 0.381 5.05 2.20 - ~~-- - ~-~ - - -JIII
108 & 0.381 5.03 2 7 7 ~ 1 ~~~2
112
141 0.450 5.13 221 - -~ -
142 0.450 5.20 2.76- --
113 0.508 4.95 2.23 -~ -~- - 1114 0.508 4.98 2.76
144 0.450 6.04 2.2146 0.4!'0 6.06 2 76 .
125 0.508 7 20 2.21
121 0.508 7 13 2.81
162 0.508 5.92 2 19 1163 0.508 6.03 2.76- 1
REGION IN EARLY CHF
U> j REGION IN DELAY CHF OR NO CHF
Fig. 1.1. Summary of Blowdown Heat-transfer Results in Test Sec-tion II (nonuniform heat flux). ANL Neg. No. 900-77-1674.
These observed phenomena were explained by the site-deactivation model related to the nucleation characteristics at the heater sur-face. 4 Under steady-state operating conditions, the Freon rig suffers severenucleation-site (cavity) deactivation on the surface of the heater. The sizesof surface cavities that remain unflooded can be approximated by the reentrantcavity model,
2
2Qrc p
= pPt,-P
where rc, a, Pt, and Pv are the critical cavity radius, the surface tension
corresponding to the local wall temperature, the local system pressure, andthe vapor pressure corresponding to
LIQUID LIQUID the wall temperature, respectively.P > This equation simply says that the
system pressure is only balanced byPE-ENTRANT the surface-tension force and the vapor
CAVITY pressure at the neck of the cavity (see
Fig. I.2)./ VAPOR
- - RSurface tension decreases(a) DEACTIVATED STAGE (D) ACTIVATED STAGE monotonically with increasing tem-
pe nature, as shown in Fig. 1.3 forFig. IALNgNo...0 7 Cavity. Freon -11, whereas vapor pressure in-
creases with temperature. The result-ing curve for the critical size cavity
as a function of wall-surface temperature exhibits a minimum, as shown in
Fig. 1.4. For those unflooded cavities to be reactivated, either the system
pressure has to drop or the local wall temperature has to increase in order
to satisfy the following equation for activation:
2arc pv -=
Figure 1.5 shows the results of this equation for various
unflooded cavity sizes in Freon-ll system. For example, an unflooded cavityof radius 90 A at a system pressure of 2.1 MPa (300 psia) would require a
wall temperature of 164"C or superheat of 15 C for reactivation. If the above
simple model is used to explain the observed blowdown heat-transfer results,
one would realize that the temperature -pressure history of the heater surface
plays an important role.
To investigate such a concern, we performed the following
definitive tests. Tests DB-145 and -146 were conducted with identical initial
conditions as shown in Table 1.1, and the blowdown transients were initiated
from a system pressure of 2.76 MPa. In DB-146, the wall temperature of the
test section was slowly raised by increasing the power input while maintaining
the system pressure and flow rate at the steady-state value. In essence, this
ensures the most severe site deactivation such that only cavities with radii of
less than 90 A (Fig. 1.4) remained unflooded in the lower portion of the heater,
i.e., the nonboiling zone.
100TEMPERATURE, 'F
200 300
TEMPERATURE, F100 200 300 400
FREON-Il400- DEACTIVATION -
400
WEIGHT
-p --;
T
"
0 50 100 150 200
TEMPERATURE , 0CFig. 1.3. Surface Tension for Freon-11.
ANL Neg. No. 900-78-300.
TEMPERATURE , f260 270 280 290 300 310 320 330 340
2.5-
2.0-
1.5
1.0
0.5
130 140 I 50 16WALL TEMPERATURE, 'C
170
E- 3000
.a
iN
G
Q-
4C)
20O0
100
REENTRANT CAVITY
FLOODED CAVITY
P =2.21 MPo(320 psio)
P,
UNFLOODED CAVITY
50 100 150WALL TEMPERATURE , 'C
I
= 2.76 MPo(400 psia)
200
Fig. 1.4. Critical-size Cavity at Freon
Pressures of 2.21 and 2.76 MPa(320 and 400 psia). ANL Neg.
No. 900-78-297.
400
300
t0
200 J0a
Fig. 1.5
Activation Criteria for Various-sizeCavities. ANL Neg. No. 900-78-299.
100
0
3
I I
FREON-Il" DUPONT" DUNN REAY9.PARACHORm a MOLECULE AR
3020
0
b
20
2
I-
E 10N)
0
d.'W
CU)Wi
4.
W1-N.N)
- i1 I i I I I
FREON-Il c co 500A
-- 130ASATURATION LINE
90A
1 I I 1I
p _ LW-
Ie I
).-
of
4
TABLE I.1. Steady-state Test Conditions
FluidSystemF, Inlet Total
Break Diameter, Pressure, Temperature, C Velocity, Power,
Test No. cm MPa Inlet Outlet M/s kW
DB-145 0.45 2.76 115.5 150 1.25 6.05
DB-146 0.45 2.76 114.5 149 1.23 6.05
As for Test DB-145, the power and flow rate were slowly
brought up at the system pressure so that the region above the 70-cm location
was observed to be in subcooling boiling (i.e., with active nucleation sites).
After the desired power was achieved, the flow was then brought up to thesteady-state value. Hence the lower portion of the heated wall was allowed to
approach the equilibrium condition (see Fig. I.4) from a higher temperature.Therefore larger sites could remain unflooded as opposed to establishing
steady -state from ever -increasing wall temperature.
From Fig. 1.5, these larger cavities can be readily activated
upon the initial depressurization, and a boiling mode of heat transfer can be
expected at these surfaces. Indeed the blowdown heat-transfer results of these
two tests differ significantly, with DB-145 experiencing greater cooling in theheated region.
Figures 1.6 and 1.7 compare the measured wall-temperaturebehavior in Tests DB-145 and -146, respectively. Early rewvet was observed
0 E T E
TIME INTO SLOWDOWN (s)20 25
230
220 425
200
-8375
170
'-' iY 325
140 l
0030
202
120 25
225
-02
TE INTO SLOWDOWN (s)
Fig. 1.6. Wall Temperatures in DB-145. ANL
Neg. Nos. 900-78-489 and -487.Fig. 1.6. (Contd.)
41) 1
750
S325t
~300
275-
250 -
225-
zoo4
U- 1C45A- C57
POWER 11
LF(LI)
POWER TA-
19
130
120
20 25
Gj 450
m0G&W T
0-0 5-0 10-0 15-0
TIME INTO BLOWDOWN (s)
200 250
4500
4500
4Wid-
I-
J
-3
425-0
400-0-
375-0 -
3500 -
325-0-
300.0- 0
275'0
250'0 -
225-0 -
200-0-
LEGEND"/ On T(.43
.0. TCSP. WRd?
POWER TRIP
230-0
220-0
210-0
200-0
- 190-0
1800
170-0
1500
140-0
130-0
120-0
-10-0
100-0
LEGEND0= TCSJ/0. Tc9t." TC754.TCg3
- -k
POWER TRIP
-0 0-0 5-0 10-0 150 20-0 250
TIME INTO SLOWDOWN (s)
230-0
220-0
210-0
200-0
- 1900
- 8C-0
170-0
5 10-0
1500
140-0
130-0
120-0
110-0
100-0
5
425-0
400-0
375-0
3500-
325-03-
s000-0
275-0
250-0
225-0
4
a-
Lii
/-
-JJ
Fig. 1.7
. ,all lcmpcratures in DB-14(;. AN ,
Neg. Nos. '00-78-469 and -480.
Fig. 1.7. (Contd.)
'e~.. uu
6
for TC51 [i.e., 51 in. (130 cm) in elevation], and efficient cooling was indicated
at TC57 and TC63 in DB-145; sustained early CHF was measured at these lo-cations in DB-146.
These two tests have since been repeated with reproducible
results. However, in the corresponding test of DB-146, TC51 experiencedcontinuous cooling as opposed to early rewet in DB-146.
These tests have clearly demonstrated the importance ofpressure -temperature history of the surface and the validity of the site-
deactivation model in explainingTEMPERATURE (C) the observed phenomena. The rele -
50 100 '50 200 250 300 350 vance of such a concern in PWR
WATER -6 systems can be shown in Fig. 1.8,'401 which indicates that at current PWR
5 pressure (15.6 MPa), the most120
severe deactivation of surface
4 cavities occurs at around 288 C.
However, this condition is ensuredP =12.1MPo
801- (750psia) _ since the startup of the PWR sys-
V--65A1 -- 0 tems is in the hot standby conditions60- (i.e., 285 C inlet temperature and
P=15.6 MPao(2260psia) 2 zero core power). If the operating
40-- -401 pressure is lowered to 12.1 MPa,PWR HOT STANDBY -I the deactivation of surface cavities
20 - TEMPERATURE is still severe. But the most im-0-0 portant change comes from the00 200 300 400 500 600 700 lower saturation temperature,
TEMPERATURE (F)resulting in a greater extended
Fig. 1.8 region in subcooled (local) boiling.These active nucleation sites have
(ritical-size Cavity Remaining l)nflooded been demonstrated to promoteat Pressures of 12.1 and 15.6 MPa (1750 greater coolability or heat transportand 22(0 psia). ANL Neg. No. 900-78-:328. during the postulated LOCA.
b. Future Work
A uniformly heated test section has been installed in the pres -
ent Freon rig, and modification of the rig to include double-ended break simu-lation is completed. Steady-state thermal-hydraulic data (including CHF
measurements) will be gathered in this new test section, followed by moreblowdown investigation.
B. Reflood Tests (Y. S. Cha, CT; R. E. Henry and P. A. Lottes, RAS)
The discharge end of the reflood-test facility was modified in order
to measure the amount of liquid being carried out of the test section during
7
TO ATMOSPHERE
TURBINE FLOWMETER
266710.16
COPPER TUBING25.4 (317 ID)
UPPER PLENUM7.62
15.247.62 TYGON TUBING
-IRECIRCULATION
TANK
the test. Figure 1.9 is a schematic
of the modified portion of the reflood-
test facility. Two passages were pro-vided to drain the liquid back into the
recirculation tank, where the liquidlevel is being measured by a pressuretransducer. The upper plenum, thecopper tubing s, and the turbine meter
were wrapped with heating wires and
were thermally insulated from thesurroundings. As shown in Fig. 1.9,the two-phase mixture leaving the
test section will experience a suddenenlargement in flow area in the upperplenum and two sharp turns throughthe copper tubing s before discharginginto the atmosphere.S1J N , ..
TEST SECTION Three forced-oscillation(1.27 1.0.) 10.22 06.68 tests (Runs No. 37-39) and one forced-
flow test (Run No. 40) were per-formed. The operating conditions
-TOfor these tests are shown in Ta-
SUPPLY17.78 ble 1.2, where T1 8 ,1 is the initialTANK temperature measured by thermo-
PRESSURE couple No. 18, f is thefrequencyALL DIMENSIONS IN cm TRANSDUCERon
of forced oscillations, Vi is the
Fig. 1.9. Schematic of Modified Por- average inlet velocity to the test
tion of Reflood Test Facility section, and Of is the opening posi-tion of the micrometer valve be-
tween the test section and the water-supply tank for forward flow (themicrometer valve is fully closed at zero position).
TABLE 1.2. Operating Conditions for Forced-oscillation and
Forced-flow Reflood Tests
T_,i Vi'pPowe r, f
Run No. OF C f, cm/s kW turns
37 1010 543 0.298 6.43 0 3.00
38 1020 549 0.943 6.39 0 3.75
39 1020 549 3.000 6.26 0 3.00
40 1020 549 0 6.55 0 2.50
8
For all the tests, the inlet water temperature was 65 C (150F), the
water -supply -tank pressure for forward flow was 0.239 MPa (20 psig), the
water -receiving -tank pressure for reverse flow was 1 atm (0.1 MPa), theopening position for the micrometer valve between the test section and the
receiving tank was 15 turns, and the temperatures at the discharge end of the
test section (including the upper plenum, the copper tubings, and the turbine
meter) were maintained at about 121 C (250 F).
The average inlet velocity to the test section (Vi) was calculated as
follows:
.-- (datf - o r(/~Vi- = ~
where otf is the charge in liquid level in the water-supply tank for forward
flow, ttr is the change in liquid level in the water-receiving tank for reverse
flow, D is the inside diameter of the supply tank and the receiving tank (which
happen to be the same) and equals 10.22 cm (4.026 in.), d is the inside diame-
ter of the test section (1.27 cm or 0.5 in.), and At is the time interval from the
start of the test to the end of the test. Table 1.3 shows the measured values
of these parameters. Also shown in Table 1.3 is the change in liquid level in
the recirculation tank (Ate), which serves as an indication of the amount of
liquid entrainment during the test. The inside diameter of the recirculation
tank is also 10.22 cm.
TABLE 1.3. Experimentally Determined Values of
tlf, Ltr, and M-e
f , UA r t, e
Run No. cm cm s cm
37 30.95 5.87 247.3 12.36
38 31.43 6.19 250.7 13.18
39 29.84 5.55 246.1 11.83
40 26.50 0 257.0 12.89
Figure I.10 shows plots of quench time versus axial distance for these
tests. These plots indicate that the quench front moved faster for tests with-
out forced-oscillation under similar conditions. The speed of the quench front
for tests with f = 0.943 Hz was slower than for other tests. These results aresimilar to those from tests reported previously that had much higher inlet
average velocity to the test section (~20 cm/s) than the present tests. The
relatively slow quench process associated with the forced-oscillation test with
f = 0.943 Hz was probably caused by the amplification of the natural oscilla-tion of the system.
AXIAL DISTANCE, ft0 2 4 6
2AXIAL DISTANCE, m
eAdditional evidence supportingthis important observation can be ob-
tained by examining the total liquid-
entrainment data shown in the last0 column of Table 1.3. Run No. 38 (f =0
2 * 0.943 Hz) had more entrainment than
2 either Run No. 37 (f = 0.298 Hz) orRun No. 39 (f = 3.000 Hz). This indi-cates that the test with f = 0.943 Hzwas more violent than the other twoforced-oscillation tests, since moreliquid was being carried out of the testsection. Thus, less liquid was avail-
able to cool the test section for approx-imately the same average inlet velocity
to the test section. Consequently, thequench process was slowed down for the
test with f = 0.943 Hz as compared to
3 the other two tests.
Fig. 1.10 The more violent oscillationfor the test with f = 0.943 Hz is also
Quench Time vs AxialDistance for Runs 37-40 supported by the information provided
in the last column of Table 1.2. Toachieve approximately the same average inlet flow to the test section, themicrometer valve between the test section and the water-supply tank had to beopened wider for Run No. 38 than for Runs No. 37 and 39. This indicates that
the flow resistance in the test section was larger for Run No. 38 (due to moreviolent oscillation) than for Runs No. 37 and 39.
Typical forward- and reverse-flow data measured by the orifice plates
are shown in Figs. 1.11-1.13. The irregular shapes of these profiles make itdifficult to obtain an accurate average inlet velocity to the test section. Thus,
the method described in the foregoing paragraphs was used to determine the
average inlet flow to the test section.
Typical entrainment-versus-time curves are shown in Fig. 1.14. It can
be seen that the rate of entrainment remained fairly constant during most ofthe test, and it decreased slightly during the latter part of the test.
250
200 --
9
I-
IzWDO
150 -
100 -
RUN NO.O 37
O 38
O 39
* 40
82
50 H-
0- I I
)i
NO FORWARD FLOW
FORWARD FLOW
REVERSE
FLOW
NO REVERSE FLOW
_ 3,s _,
FORWARD FLOWNO FORWARD FLOW
REVERSE FLOW NO REVERSE FLOW
is
Fig. I.11. Typical Plot of Flow vs Time for Run 37
FORWARD FLOW NO FORWARD FLOW
REVERSE FLOWNO REVERSE FLOW
5s
Fig. 1.13. Typical Plot of Flow vs Time for Run 39
Fig. 1.12. Typical Plot of Flow vs Time for Run :38
TIME
Fig. 1.14. Change in Liquid Level in Recircula-tion Tank vs Time for Runs 38 and 39
10
1.2
> 1.0
U
z S0.8
U)0.6a-
o0 0.4
F-
a.r-o 0.2
RUN NO. 39
60 $
RUN NO. 38.
11
References
1. J. C. M. Leung and R. E. Henry, "Transient Critical Heat Flux," Light-
water-reactor Safety Research Program: Quarterly Progress Report,Oct-Dec 1976, ANL-77-10 (Mar 1977).
2. Ibid. Jan-Mar 1977, ANL-77-34 (June 1977).
3. Ibid. April-June 1977, ANL -77 -59 (1977).
4. R. E. Henry and J. C. M. Leung, "A Mechanism for Transient Critical
Heat Flux," Proc. Topical Meet. Thermal Reactor Safety, Vol. 2, p. 692
(1977).
12
II. TRANSIENT FUEL RESPONSE AND FISSION-PRODUCTRELEASE PROGRAM
Responsible Group Leaders:
L. A. Neimark and M. C. Billone, MSD
Coordinated by:
L. R. Kelman, MSD
A physically realistic description of fuel swelling and fission-product
release is needed to aid in predicting the behavior of fuel rods and fission prod-
ucts under certain hypothetical light-water-reactor (LWR) accident conditions.
To satisfy the near-term need, a comprehensive computer-based model, the
Steady State and Transient Gas-release and Swelling Subroutine (GRASS-SST),is being developed at Argonne National Laboratory (ANL). This model is being
incorporated into the Fuel-rod Analysis Program (FRAP) code being developed
by EG&G Idaho, Inc., at the Idaho National Engineering Laboratory (INEL).
Also being developed, but at a lower priority, is a model to predict the behaviorof volatile fission products in, and release from, LWR fuel under hypothetical
accident conditions. The volatile fission-product results will also serve as
input to NRC-sponsored programs that are developing a radiological source
term for hypothetical accidents.
The analytical effort is supported by a data base and correlations
developed from characterization of irradiated LWR fuel and from out-of-
reactor transient heating tests of irradiated commercial and experimental
LWR fuel under a range of thermal conditions.
Emphasis in the early stages of the program has been on thermal
conditions in pressurized-water reactor (PWR) fuel that are applicable to
anticipated hypothetical power -cooling -mismatch (PCM) accidents. Recent
efforts include conditions typical of other types of hypothetical accidents. The
program is also developing information on fission-gas release during steady-
state and load-following operations.
L. R. I<elman presented a talk on Correlations for Transient Fission-
gas Release from High-burnup Fuel (by S. M. Gehl, J. Rest, and L. R. Kelman)
at the Fifth Annual U. S. Nuclear Regulatory Commission's Water Reactor
Safety Research Information Meeting held at the National Bureau of Standards
in Gaithersburg, MD, on November 7-11, 1977. S. M. Gehl visited EG&G Idahoon December 6 and 7 for an informal discussion of the ANL-EG&G crosscheck-
program results.
Recent significant analytical and experimental advances and the status
of the program at the end of this quarter are summarized below.
13
1. GRASS-SST analyses indicate that gas release during DEH tests
is strongly dependent on the induced hydrostatic pressure in the fuel for pres-
sures above ~10 atm (~4 MPa).
2. GRASS -SST was used in conjunction with the LIFE-LWR fuel-behavior code and the DEHTTD code to simulate the irradiation history ofSaxton Rod 843 followed by a DEH test on a single fuel pellet. The steady-
state and transient gas-release predictions are in reasonable agreement withexperimental results.
3. GRASS-SST -Mod 4 and a GRASS-SST driver program were sentto INEL.
4. A topical report on the development and verification of GRASS-SST -Mod 4 has been written.
5. The laser sampling technique was used to determine the radial
profile of fission-gas release during DEH Test 33.
6. The experimental work performed for the ANL-EG&G crosscheckprogram was completed.
7. Three DEH tests were run as part of a study of fission-gas releaseand fuel behavior during loss -of -cooling accidents (LOCA's). Preliminary
evaluations of the transient temperature histories and microstructural evolu-tion have been made.
A. Modeling of Fuel-Fission-product Behavior (J. Rest, MSD)
1. Modeling of Fission-gas Behavior during Transient Heating
Conditions
a. GRASS-SST Calculation of Transient Gas Release during DEHTest of Saxton Load Follower Fuel
GRASS-SST was run for the irradiation of Rod 843 in theSaxton reactor, followed by a DEH transient test of a fuel pellet extractedfrom the irradiated rod. The purpose of the steady-state simulation is toestablish the initial conditions in the fuel pellet before DEH testing (e.g.,bubble -size distributions and amounts and location of retained gas). For thesteady-state simulation of Rod 843, GRASS-SST was run with an experimental
version of the LIFE-LWR fuel-behavior code (see ANL-76-121, pp. 50-61).Table II.1 shows results of the steady-state simulation for total gas releasefrom the rod, and for the quantity of gas retained in a fuel pellet located~46 cm from the bottom of the rod. Also listed in Table II.1 are the experi-
mentally determined values (see ANL-77-34, p. 66, and ANL-78-3, p. 42) forthese quantities.
14
TABLE II.1. GRASS -SST -calculated Values Compared with ExperimentalResults for Total Gas Release from Saxton Rod 843 and for the Quantity
of Gas Retained in a Fuel Pellet Located --46 cm from theBottom of the Rod
Total Calculated Total MeasuredTotal Calculated Total Measured Quantity of Gas Quantity of Gas
Gas Release from Gas Release from Remaining in Fuel Remaining in FuelRod 843, moles Rod 843, moles Pellet, mmol/g Pellet, mmol/g
1.5 x 10-' 0.958 x 10-1 0.017 0.016
Figure II.1 compares the results of GRASS-SST calculations
of transient gas release to the measured results for PCM -type DEH tests on
fuel pellets irradiated in the H. B. Robinson reactor (see ANL-78 -3, Fig. 11.3)and for the fuel pellet irradiated
60 DET T N 26 T in the Saxton reactor. As is evi-
ARE INDICATED dent by comparing with the5- prO B. ROBINSOt-agreement diagonal line,
p5o the GRASS-SST predictions for
4 transient gas release during DEH
I tests are in good agreement with
SURFACE TEMPERATURES ESTIMATED- the experimental measurements127 SPREAD INCALCULATED ET Ifor fuel irradiated in both the
DIFRNE IN THEMAIUSSURACEN TEMPERATMXIMUM H. B. Robinson and the Saxton
25 reactor. The irradiation history20- 35 of the fuel irradiated in the Saxton15 22 f 32 Ireactor was quite different from
324 that of fuel irradiated in the10,-
23 H. B. Robinson reactor. (Saxton
1 fuel had a high-power, load-
0 5 10 IS 20 25 30 35 40 45 50 55 60 6570 following history, compared to
PREDICTED GAS RELEASE (/) the relatively flat, low-powerirradiation of H. B. Robinson fuel.)
Fig. 11.1. GRASs-ss T-predicted Gas Release vs Experimen- The results shown in Fig. II.1tally Measured Values. Neg. No. MSI>-64855. therefore support the contention
that the models used in GRASS -SST
are physically realistic and predictive of fission-gas behavior. As additional
information becomes available, the applicability of the models used in
GRASS-SST to describe physical processes controlling the behavior of fission
gas in irradiated fuel during steady-state and transient heating conditions will
continue to be examined.
b. Effect of Restraint on Fission-gas Release during DEH Tests
All DEH tests run to date have been on fuel pellets that are
radially unconstrained (an axial load is applied). The extensive pattern ofcracks in the form of grain-boundary separations observed in DEH-tested
pellets and the appreciable release of fission gas from pellets showing exten-sive grain-boundary separation could be influenced by the degree of radial
15
constraint on the pellet during transient heating. In an attempt to obtain a
qualitative picture of the effect of radial restraint on fission-gas release,GRASS-SST was run for DEH Test 33 for several values of a fixed hydrostaticstress. (In GRASS-SST all bubbles are assumed to be spherical, and the effectof shear stress is ignored.) The results of this a Iysis are shown inTable 11.2.
TABLE 11.2. GRASS-SST-calculated Transient Gas Release
for DEH Test 33 for Several Values of a
Constant Hydrostatic Stress
Hydrostatic Stress, Calculated Fractional Gas Release,atna
1 39
10 39
100 30
1000 6
aConversion factor: 1 atm = 0.1 MPa.
The results shown in Table 11.2 indicate that the amount ofgas release during transient heating will be moderately reduced for values ofhydrostatic stress between 10 and 100 atm (1 and 10 MPa). For hydrostatic
stresses above ~100 atm (~10 MPa), the transient gas release is predicted tobe significantly limited. The reason for the decrease in GRASS-SST -predictedtransient gas release for high values of the hydrostatic stress is that the in-creased stress reduces the degree of fuel swelling and, thus, reduces the ex-
tent of long-range porosity interconnection.
In GRASS-SST calculations, the fission gas residing along thegrain edges is released via the long-range interconnection of porosity (seeANL-78-3, pp. 23-25). GRASS-SST appears to adequately model transient gasrelease from radially unconstrained fuel where grain-boundary separation hasoccurred (e.g., see Fig. II.1). Whether this model correctly accounts for therelease of gas from the grain edges for high values of stress will be determined
when data become available from planned DEH tests on irradiated pellets havinga high degree of radial constraint.
c. Additional Progress
GRASS-SST -Mod 4 and a GRASS-SST driver program weresent to INEL. A report on the development and verification of GRASS -SST -Mod 4 was written this quarter and is now being prepared for publication.
16
B. Experimental Program (S. M. Gehl and L. R. Kelman, MSD)
The laser sampling technique was used to determine the fission gas
retained in the fuel after DEH Test 33. Three tests in the series of LOCA simu-
lations and preliminary metallographic examinations of LOCA-tested specimens
were performed. The results of the ANL-EG&G crosscheck program were
analyzed, and a report of this investigation is being written.
1. Determination of Radial Fission-gas Profiles for DEH -tested
Fuel (S. M. Gehl, MSD; and D. G. Graczyk, CEN)
The laser sampling technique is used at ANL to determine pro-
files of retained fission-gas concentration in LWR and LMFBR fuels. In this
technique, a focused laser pulse is used to vaporize a small volume of fuel,
and the released fission gas is collected. Part of the fission gas in a heat-
affected zone surrounding the crater left by the laser is released and is also
collected. The 8 5Kr activity in the combined sample is determined, and the
total amount of fission gas is calculated from the known isotopic and Xe/Kr
ratios.
The laser sampling technique has already been used to examine
the radial profile of fission -gas retention in the as -irradiated Robinson fuel
(see ANL-76-121, p. 75). Large confidence intervals were assigned to the
fission-gas concentrations because of the uncertainties in the measurements
of the volume of the laser crater. The confidence intervals were large enough
to mask the radial variation in concentration.
The laser system and the accuracy of crater-volume measurement
have been improved since the early measurements were made. The improved
equipment and techniques were used in a recent determination of the radial
fission-gas profile of the Robinson fuel from DEH Test 33.
The Test 33 specimen was selected because it contained examples
of all the microstructure types observed in the DEH-tested Robinson fuel. A
transverse section through the midplane of this specimen is shown in ANL-78-3
(Fig. 11.9). The specimen was remounted, thinned, and repolished to reduce
the radioactivity level before laser sampling. The condition of the specimen
before laser sampling is shown in Fig. 11.2. The missing arc of fuel was
ground away during the thinning process.
The appearance of the specimen after laser sampling is shown in
Fig. 11.3. Several laser samples were taken at each of 10 radial positions,
including the melt zone. At higher magnification, shown in Fig. 11.4, a laser
hole is visible, surrounded by a rim of solidified fuel. Several rings of heat-
affected and/or stained material are present out side of the crater. ("Crater"
refers to the hole plus the solidified rim.)
F-I
* [
l, l~ -ll
SIC i*L. -. ,~ ~ ", h l c I.-
Fi.I. . Spec1imn , \ ( ( \1S)- 1:18 l er
- * S
Fig'. II1.;. Speccime1Ln 1 5 ::os a tr l
I plng WY.,.,AlD- m3
ILA~* ~ E 'P9'~
Pig. 11,1
Detail of Typical Lascr Crattr and Ilcm--
affcctCd Zonc. Ncg. No. NISI}-_.I:.15:1
100 p1m
17
18
The radial profile of retained fission gas, as determined by lasersampling, was normalized to the volume-averaged retention obtained from
the measurement of gas release during DEH testing (see ANL-77-59, p. 88).The pretransient radial gas distribution was assumed to be uniform for thiscalculation. The retained fission-gas concentrations were converted to frac-tional fission-gas release values to be consistent with the previously reportedgas -release data.
The radial profile of fission-gas release during DEH Test 33 isshown in Fig. 11.5. This figure illustrates several important features of the
gas-release processes. First, moreoo( -'I1 ' than 90% of the ,as was released from
the fuel that melted during DEH test-
ing. Second, a like amount of gas wasMELTED released from the unmelied fuel that
UNMELTED was immediately adjacent to the melt
zone. Third, the gas-release fractiondecreased smoothly with increasing
1 radius for the rest of the specimen.
40o 4 About 20% of the gas was released fromthe outermost radial position.
20 The gas-release data were also
plotted as a function of the maximum
local temperature in each radial zone,
0 0.2 0.4 0.6 0.8 1.0 as determined by the DEI-ITTD codeFRACTIONAL RADIUS (J. C. Vogelwede, ANL, private com-
munication). As shown in Fig. 11.6,Fig. 11.5 gas release increased most rapidly
Radial Profile of Krypton Release from Speci- for temperatures in the rangemen 155CC1513. ANL Neg. No. 30-78-230. 2000 -2400 C. For temperatures in
excess of ~2700 C, gas release ap-
peared to become saturated at ~92%. Little additional gas release wasobserved for temperatures up to 2900 C.
When the local krypton-release fraction is plotted against the lo-cal volume swelling, as shown in Fig. 11.7, gas release shows a saturationeffect for swelling values in excess of 50%. The smooth variation of gas-release fraction with local temperature and swelling values indicates the self -consistency of these experimental data, obtained from three independent
sources.
We wish to emphasize that the laser sampling data presented here
were obtained from a single specimen, operated under a particular set oftransient conditions. Therefore, caution should be used in extending theseresults to other test conditions, in particular, to transients in which the heatingrates are significantly higher than the ~15*C/s of Test 33. However, the pres-ent results do indicate the type and precision of the information obtainable
19
with the laser sampling technique. Additional laser sampling studies can beused to provide information needed for model verification and to answer spe-cific safety questions.
i 00o1
80[
WQj
W-JW
60
-- -
80 - -
W -SWI
60-1
a-
40i --
20. -
V - --- - - -1500 2000 2500 30
MAXIMUM LOCAL TEMPERATURE ( C)
Fig. II.G. Fractional K rypton Release vs
Maximum Local Temperature.
ANL Neg. No. :30-78-234.
20
00 0 20 40 60
% VOLUME SWELLING
Fig. 11.7. Fractional Krypton Releasevs Local Volume Swelling.
ANL Neg. No. 20G-78-2:33.
2. ANL-EG&G Crosscheck Program (S. M. Gehl, MSD)
a. Summary
The experimental portion of the crosscheck program has beencompleted. A final report describing all aspects of this program is being
prepared. The principal findings of the program are summarized here in ad-
vance of the final report.
(1) The ANL-EG&G crosscheck program demonstrated thatplanar intergranular separations are formed during nuclear- and electrical-heated transients of the Saxton fuel. Therefore, the separations are not anartifact of the DEH technique.
(2) Three types of transient-tested fuel were examined aspart of the crosscheck program: DEH-tested Saxton fuel, PBF-tested Saxtonfuel, and DEH-tested Robinson fuel. Microstructural differences among these
WU/)
WJW
80
I M
40[
.1
20
fuel types after transient testing were primarily the result of the pretransientirradiation history and the pretransient distribution of retained fission ,gases.
The transient heating method, i.e., nuclear or electrical, had a secondary ef-
fect on microstructural evolution; this effect was tentatively ascribed to thelack of radial constraint in the DEH tests.
(3) Gas -analysis results indicated that the fractional rod-
averaged release from Rod 843 during the Saxton irradiation was ~0.085 and
that, at the peak power position, the release was -0.27. About 50% of the re-
tained gas was released from the fuel just above the film-boiling boundary
during the PBF test of Rod 008.
(4) A fractional gas release of -0.16 occurred during a DEHtest of the Saxton fuel. Transient heating of the Robinson fuel to the same
centerline temperature would give release fractions of -0.40. The lower re-
lease fraction for the Saxton fuel is consistent with the low pretransient gas
burden in the central zone of the fuel. The difference between the fractional
gas release of DEH- and PBF -tested Saxton fuels is due to the differing tran-
sient temperature histories.
b. SEM Examination
Recently, efgrtin the crosscheck program has been concen-trated on the interpretation of the fuel microstructures as revealed in the SEM.
Transverse sections through Saxton-irradiated fuel specimens
showed two distinct microstructural zones: an outer unrestructured region,
and an inner zone in which grain growth had occurred and intergranular poros-ity was present. Examples of these structures were given in Fig. II.10 of
ANL-77-34.
The intergranular pore and bubble mor'phologies in the central
grain-growth region of Saxton -irradiated Rod 843 are shown in the SEM fracto-graphs of Fig. 11.8. The predominantly intergranular fracture surface in
Fig. II.8a has exposed grain-face bubbles with diameters in the range .0.5-2.0 pm; small, isolated grain-edge bubbles with diameters <1.0 m; and large
and apparently multiply connected channels of intergranular porosity. Simi-
lar features are present in Fig. II.8b. The multiply connected nature of theintergranular porosity is accentuated in the latter micrograph.
A typical SEM fractograph of the outer, unrestructured fuel
zone, shown in Fig. 11.9, shows an almost exclusively transgranular fracture,
in contrast to the intergranular fracture observed in the grain-growth region.Because of the difference in fracture mode, grain-surface bubbles would not
be as evident in the outer region, even if equal densities of bubbles were pres-
ent in both zones. However, close inspection of the grain boundaries inFig. II.9 reveals no indication of grain-surface bubbles down to the -35-nm
limit of resolution. Several intergranular pores with maximum dimensions
.t ' 7>
a. 10pmtl h. 10 lm
Fig. 11.8. SE[MI Fractographs from Grain-growth Region near Peak-power Po-
sition of Rod 843. Spec. No. AG178A2. Neg. No. MSI.931418.
r " 4r
Fig. II.9
S EM Fractograph from Unrestructured Zone
near Peak-power Position of Rod 84:3. Spec.
No. AG178A2. Neg. No. MST>-204302.
10pm
21
22
in the range 1-5 m are visible in Fig. 11.9. These features were formed
during pellet fabrication and were apparently little changed during irradiation.
The examination of the unrestructured fuel zone disclosed no evidence for an
interconnected network of porosity.
Figure II.10 shows theexisting crack in the as -irradiated fuel
- 4
10pn
Fig. 11.10. Surface of Preexisting Crack in Unre-
structured Lone near Peak Power Posi-tion of Rod 84:3. Spec. No. A G178A2.Neg. No. MSD-204.306.
Figs. II.lla and II.1lb are covered with
5EM image of the surface of a pre -
at a fractional radius of -0.8. High-
temperature operation of the fuel
after formation of the crack resulted
in deep thermal grooves at the grain
boundaries. However, the fuel tem-
peratures were apparently low enough
to preclude crack healing. The grain
surfaces are relatively flat at positions
removed from the grain boundan..1 s.
The morphology of the surface of a
preexisting crack is presented to
allow comparison with the crack
morphologies formed during transient
heating.
The SEM examination of frac-ture surfaces prepared from a speci-men from the film-boiling region ofa PBF -tested rod exhibited importantdifferences from the fracture surfacesin the as -irradiated material. The"unre structured" zone of PBF -te stedfuel, shown in Fig. II.11, fracturedintergranularly, in contrast to thepredominantly transgranular fractureof the as-irradiated fuel shown inFig. 11.9. The grain surfaces in
bubbles with diameters in the range
35 nm (i.e., the resolution limit of the SEM) to -0.1 pm. Slightly closer to the
fuel centerline (Figs. II.llc and II.lld), the bubble size range is -50 nm to
-0.5 pm, with most bubbles larger than 0.1 pm.
Large pores formed during fabrication are also visible on thegrain surfaces in the unrestructured region.
At somewhat hotter positions in the specimen, i.e., in the outerportion of the powdery fuel zone, further development of the bubble structureoccurred. (See ANL-77-34, pp. 67-68, for a description of powdery fuel.)Figure II.12a shows a region in which interlinkage of the grain-surface bubblesto form sinuous channels has occurred on some of the grain surfaces. Severalof the smaller grain surfaces are smooth, indicating that complete separation
l "V
A*f 4 'A
a.
I A Ihal0ilp 97-
I ** @
c.
eq -"40It
Vr '4
aI
1011m d.
Fig. II.11.. 51:\ Fract raph> ofl "Lnrciriutiuiu" R 4gfin n [ilmt-w Alin; /_O1n )I
PBF- L>tCLd Rol COT. Spcc. No. A G1 781 . *N. :). N IS I: .9214.
23
4
eta
10 jim n. 1 .(i p
1 .0 1111
I 71tJW .A
E *
1.4
r /4 , N
~ ..
/4
e> -~
~v w~4
A
-t -
-YS -t.F---
_. r
-I
V'S
of two grains has O()curred. Similar featTaTres are resentt . . cx -
cept ;fat the bubbles and channels are iaceted. Tnhe faceted appearanceprobably the result of the predomin.ance of surface and or vapor--transnortmechanisms. Figure II.l1e shows a region in which bubble . \ th as o crr!
4> a greater extent than in Fig. II. va. Note. however, that the formation ofsinuous tunnels was almost nonexistent at the location shown in Fig. !1.1 I.
The three diverse structures in ig. II. 12 have in comnitn areduction in grain-boundary strength that nay have contributed to the powderv
nature of the fuel. For example. extensive grain-boundary cracking may have
occurred during the rapid cooldown following the end of film boiling. In addi -
tion, the smooth grain surfaces in Fits. 11. 1 .a and 11. 121b indicate that some ra in-
boundary separation occurred while the fuel was still at elevated tcmperat.re.
Figure II. 13 shows the formation of separations in a tempe ra-ture range where diffusional prices ses are import ant'. Ithe upper surface ofthe larger grain in the center of Fig. II.13a shows a transition from a smooth.nearly featureless surface in the foreground to sinuous channels in the bacK -ground. Figure II.13b is a high-rmagnificat ion view of this surface.. Ihe sinothsurface is an area of total separation of the grains, and may be viewed as an
intergranular crack that was propagating from the front to the rear of the grain
surface. This interpretation suggests the presence of a tensile-stress colmpt.)o-
nent normal to the grain surface. The crack did not reach the portion of tihe
grain surface in which the sinuous channels are present.
The fact that the remnants of the channels are not. vi silc in
the portion of the grain boundary through which the crack front passed in>i ate sthat diffusional processes acted to smooth the surface after the grain surfaces
separated. The small particles visible on the smooth surface in Fig. II.13bU arecondensed fission products (dark gray) and loose debris deposited during spec i -men preparation and handling (light gray).
The front surface of the large grain in Fig. I1.1 3a evolved bya somewhat different sequence of events than the upper surface. Consider r a
grain boundary in which interlinkage of the bubbles and vent ing of the gas hasoccurred. Without a tensile stress of sufficient magnitude, further separation
of the grains will not occur. In fact., since the grain-boundary channels we rCdepressurized when the gas was vented, further heating of the fuel will cause
the partially separated grains to sinter back together and the existing channels
to pinch off.
The front surface of the large grain in Fig. II.13a, shown athigh magnification in Fig. II.13c, appears to have evolved by this sequence.Vestiges of the prior interconnected channels are visible and occupy ~30" ofthe projected area of this surface. The narrow necks or filaments that join
the larger-diameter portions of the structure are in the process of pinching off.
_ ~.
'ir~
N r
I,.
wv;f, v~
4
I-
* 4j
~~Y,
This mo rphology is superficially similar to that present onseveral of the grain surfaces in Fig. II.12a, a structure ascribed to the forma-
tion stage of the interconnected channels. IIowevez, Fig. II. 1a indicates thatisolated bubbles grow until they occupy about half of a given grain contact be-fore interlinkage occurs. Note also that, in Fig. II.l1ia, the connections betweenthe bubbles have the same diameters as the bubbles themselves.
The necking down of originally straight-sided channels is the
thermodynamically predicted response to surface -tension-driven phenomena
such as sintering. Therefore, the necked channels in Fig. I1.13c are taken as
evidence of resintering of the grains following venting of the fission gas. The
relatively low areal coverage of the bubbles and channels is consistent with this
hypothesis.
,6
A r
*ti
'ig. I1.14a. SmiiI aI nr'i-", nr(' 1 It n
(0ver7 1he field of viev. (xcAp ( n I l -1 rIn) .
\lost of the grain surfaces a1nd( ( ' 1- Ca1 IA 'C (
able rounding O. these' f atu re s s < 1rred. 1 ! '1 - . l' .- (
are id:ratifiable, this si r1u11,tu.re ftrh e c c?)lnI !1,.
gui shed froin the structure tf a r)Ix('X !i tn' Cr n
61 r
T'he fracture surf ace OI unihiel tedi fel 11 S: < n :4
zone is shown in Fi . II. lb. Ih;is strut1u1rc is si I1 l r u 117r n 5 n(
str u ctu re of fuel from the gra1 n -g r lv th regI uon shown in l\. ."!.
SEM exami nations 1f fracture s1rUfaccI repar('l ;I: . Ill
tested Robinson fuel revealed Iany of the same foal ures )hs5r (d ii n w II
tested fuel. As shown in Vit. II.1 . isolated s;pleri(a1 hubiles. lC'ImCrihn e<I'ti
channels, and complete separations are visible on t1e grain fI acls 1: ? I b ;11
types. In both cases, these three st ructure type s occur cu di ff rent Les
the same grain.
Several. structures were Ohserved (In the SaxtIn fl ut nut
in the Robinson fuel. These structures are the faceted biibbles shown :n
Fig. II.1 1b and the resintered channels shown in ig. 11.1 . T he rifrm ti on :
both of these morphologies is thought to leyCuire relatively lcon' times at e -vated temperatures. These conditions were mo re typical Of the PI l test. ' T'
structures may have been formed during the high-power ope rat i on bn ore 1hestart of film boiling.
\~ 1-4
AS
;v A
"
I"
i
" 1
i
. wr
A.-A
DEH-tested Roino ~ -Robinson2 pim
w, +
t ,
'I
t
PBF-tested Saxton
AIIhou21h ith'se St rl1ctli ,s are pre ferntIllyIv fb)r (r(l linier
s5rnIlar tonditrions, lhewv :ar rv simIllewhalbdt di ffe rent imllohicat ins about the rt:-
waSc of ::s1 on gas. lh' 1 ;lt'irilh of grPa1n1-fa((e channllcls is part of 1h1(
".r(I n1 ar ' Sequence ( ))st(rv(d n bthi )11 (l tvp)('. Tlat is, r''sintering follows
the g'owthi. coal (sn('fct. dl >ie r rlina12e o f the I)bibb les. an(] thil ventn11gT 0)
the contained _,as. The relative priominante of the resintered strilcture in
the P1,) -tested hiel is eviriel(c that the same imlelchanismr of Las release by(hilllnel oalescenye m)( 'urS I n bo;1 l uel ;'l es.
Th, (x.ln e ffa eed bubblMes dlots 110 slterth 1("1'l t 11hat
bu1lAc intr. c iKag n a til .i nla a role in gias reI aLs( . I lO(w'\.'r. biibblie
iac'eti 1ngL in>IlicatLts that biibble ! l wtI'P ; h is OCCu rI i at a slower rat thanlll \'whfei
the bubbles are sLheri al. (See ANL-78-. pp. 16-".) I' ower as-releasefractions, or retarded gas release, may therefore occur for bubble -facete(I
structures. lor the IT' test fuel examined, there were fewer a reas with
faceted bubbles. TLerfor(, the effect of this structure on gas release is ex-pected to hav( been ninor.
3. IAC;- Sim lat ions (S. M . GeMh. ID. R. Pepal is, and J1. A. fluzzel I,MlSl)
Fou r D11H tests have been performed as part of a p relin i narystudy of fission-gas release during DEIH simulations of LOCA thermal condi-
tions. The first two experiments were long,-time (>1500-s), near -steady -state
tests at surface temperatures of ~1800 and 1600 C, respectively. Tlhe tests
used a combination of DEI and external heating by a. tungsten-mesh element.
17 1 aktor D ( l',1 ) n r:. t tr-1S '' w'r 2 i . I :i -
HP >I) -7. pp. 3.1 -4 10 (lar 7. 1.1781. Ihese tests (Im (enunstrated ihat radial
~f-
; -.1 ..
*
- 4- - *
I \ Kjf . . V4
4.
*'J,4. a -- - 'V .4
- -*' *
* 0 tie
Ic / j)V|ipm
Ii .tth \
Ii.. 1.
rain -hbi tvria y' 5'wV11j!in. \ smial lt
a t (ili)ct . at r t t li lil \I, s \' not.' a i so
)t r r cc . xe CCV t S PIr Vri- : ed f
ii nl._lc r t'Iihi e( re I ri2 1i p th. \ ( .
.II.1 a n e b e ne
S iO wa(5\\S 21n ta 1010 rI niig C)(111 1 C
I'e ten rtu e ttU i II sitr () y rl 1an1 1K \1 ,t1
: I(1'.'A cal c i ati i s 02 Stadldari 0 U)r1 '1b-
S 1. l5 PIt) 1(1 )i I u .r)
78 - .) D i ing I mws I th D )I1r CIt? iiiat 1,i117 i(( te:)1 0 rat Xi1ix
V V ' ' t I' X t'i 1 !.I'tt , ric Vt 1 t . ( 1d f
t1ti OS dt , S ... l ' (:. AS Ct r7C5ul!I.
I It i 510a I I ,ct r th r 1) ( 1.1 at 'in S
11)1i no Ie ptri tr: CC tI QC *' t
= (: ' t () iC Vi i S i' l l 1t i ('(3 : t' ,U I hi '
ol ,' t'as nt' nt (' . I tit1 fr I. I -
e S imi la I lted iiCut' 1r - ot' rt'" rcifl' l It })r
ft(i' titIo f rth TA)C1A test he nt't.' rI1ne tetCIm.-irat lrkts w"1r 174.' ' lc latld(l 1)',
assui ng , a craclking fa tlor (t 0.L. (See A 1.-78 - i. pp. -11 1. or ;a (iscISSI
Of cracking factors.) loer',' 1 was raIpC1ed in the sai.' ImlannCcr ;Is tOr 1 C.
simulation t() p)ru)(IIte a t(C'lt'rlinlO t mpiI)era'Luri e of - ;O (: C at thet. SIaI'l ot tK
blowdown sinolat ion. At the start (f )iOv.down, 11lw surface tmnpe 111r u:t i II
the 1)1-M test exceeded that of tIhe standard iO roblem No. 1. ThR 1)1II surtaco -
temperature history approached the standard 1 probtIlen No. I hi stttr' wi ti in -
creasing time in the transient. 'The centerline temipe raturc in the DIIII test
)k i l
flt x,1 '1r t 1ac 1)' '-ttteratur I h st (lry ;an1( t
30
decreased from 2250 to 1750 C in ~48 s, a'drop that occurs in .5 s in standardproblem No. 1. Although higher rates of temperature decrease can be achievedby changing the DEH power input, the rate called for in standard problem No. 1is probably not attainable.
2500--
2000 / CENTER
WN1500
0 20 40 60 80 100 120
TIME (s)
Fig. U.17. Centerline and Surface Temperature Histories forDEH Test LOCA 4. ANL Neg. No. 306-78-235.
The gas release and microstructural evolution during the fourthLOCA test have not yet been determined.
31
III. MECHANICAL PROPERTIES OF ZIRCALOY CONTAINING OXYGEN
H. M. Chung, A. M. Garde, and T. F. Kassner, MSD
A. Summary
The impact properties of homogeneous and composite Zircaloy material
with a range of oxygen concentrations are being determined at temperatures to
-800 K. The -ynarr fracture toughness KID, ductility index, and equivalent
yield stress have been calculated from the curves of load and energy versus
time. Initial results show that the dynamic fracture toughness decreases asthe o:vygen content of homogeneous Zircaloy specimens increases. The ductility
index of homogeneous and composite specimens increases as the test tempera-
ture increases, at a given oxygen content, and the nil-ductility temperature
increases with oxygen concentration in homogent -us alloys. For a given oxygen
content, composite specimens exhibit a lower nil-ductility temperature than
homogeneous Zircaloy specimens because of the higher ductility of the centraltransformed-S region of the former mater..
Additional tests have been performed to determine the capability ofZircaloy cladding to withstand thermal shock during quenching from ~810 K
by bottom-flooding the tube with water after rupture and isothermal oxidationin steam for various times at temperatures up to ~1550 K. The results indicate
that the cladding will fail by thermal shock after isothermal oxidation of the
inner and outer surfaces for periods of -10, 2, 0.7, and 0.3 ks at 1300, 1400,
1500, and 1600 K, respectively. These time-temperature conditions producedtotal oxide layer thicknesses greater than 17% of the cladding wall, which is
the present limit in the acceptance criteria for emergency core-cooling sys-
tems in light-water reactors (LWR's). The thermal-shock and impact data
will be used to establish a more quantitative cladding-embrittlement criterionfor Zircaloy fuel cladding under postulated loss-of-coolant-accident (LOCA)
situations in LWR's.
B. Impact Properties of Zircaloy-Oxygen Alloys
The impact-testing system, specimen geometry, test procedures, and
data analysis methods were described in Ref. 1. Initial results have been ob-tained for the impact properties of as-received specimens (0.14 wt % oxygen)
with longitudinal and transverse textures and homogeneous and composite Zir-
caloy specimens with up to 1.1 and 1.6 wt % oxygen, respectively.
Figure III.1 shows the dependence of the equivalent-energy fracturetoughness KID and the ductility index DI on the impact velocity at 300 K foras-received Zircaloy-2 specimens with a transverse texture. Although the
ductility index decreases considerably with an increase in impact velocity, the
dynamic fracture toughness is not highly dependent on velocity over the range
of 1.5-4.0 m/s.
0.9 -
-759
0.8
0.7 F0.6
0.5
0.4 -1.0
0.14WT % OXYGEN
2.0
IMPACT
3.0
VELOCITY (m/s)
58 E
0a
c1-1
56 '
55 I-444
54
0 O
L 534.0
Fig. 111.1. Dependence of Ductility Index andEquivalent-energy Fracture Tough-
ness on Impact Velocity for As-.received Zirealoy-4 Specimens at
300 K. The subsize V-notch charpy
specimens had a transverse texture
(i.e., perpendicular to the rollingdirection of the sheet). Neg.
No. MSD-64997.
equivalent stress than specimens
-1
fracture occurs in the latter material before significant plastic deformation
takes place; i.e., the equivalent stress
material at this oxygen level.
900
U)U
F-
zw'44
0WA
--
800--
700 -
600
500 -- 0C'
400 0.14 WT % OXYGEN
300 - ---_---- ------ I1.0 2.0 3.0 4.0
IMPACT VELOCITY (m/s)
Fig. 111.2. Variation of Equivalent YieldStress with Impact Velocity forAs-received Zircaloy-4 Speci-mens with Transverse Texture at300 K, Neg. No. MSD-64996.
is much lower than the yield stress of
700
- 600Q-
500
U,W 400U,
Z 300
2J
>200
0LaJ
.--- 1. r[ 1 I 1
IMPACT VELOCITY.5 rn/s
0.14 WT % OXYGENo LONGITUDINAL ZIRCALOY-2
0 TRANSVERSE ZIRCALOY-20 LONGITUDINAL ZIRCALOY-4
---1
4 - -
0 100 200 300 400 500 600 700 800
TEMPERATURE ( K)
Fig. 111.3. Equivalent Yield Stress as a Func-tion of Temperature for As-receivedZircalcy Specimens with Transverseand Longitudinal Textures. Neg.No. MSD-65000.
32
0z
0
The results in Fig. 111.2 show that
the equivalent yield stress ay, which is
computed from load-versus-time data and
the dimensions of the three-point-bend im-
pact specimen, is essentially independent
of the impact velocity. Based upon these
results, most of the tests were conductedat an impact velocity of 1.5 m/s. (That is,the ductility index appears to be almostindependent of the impact velocity for val-
ues1 3.0 im/S.)
Figure 111.3 shows the temperaturedependence of the equivalent stress (i.e.,yield stress) for as-received Zircaloy-2and -4 specimens with different textures.Figure 111.4 shows the effect of oxygen con-
centration in homogeneous Zircaloy- 2
specimens on the yield stress at several
temperatures. In these tests, the materialwith 0.6 wt % oxygen exhibits a higher
with 1.1 wt % oxygen. This suggests that
1
i
-1- ----__i
33
700
600
500
400
300
200ZIRCALOY -2
100300 400 500 600 700 800 900
TEMPERATURE (*K)
Figure III.5 shows the temperature dependence of the dynamic fracture
toughness of as-received Zircaloy-2 and -4 specimens with longitudinal and
transverse textures. Above 300 K, the fracture toughness increases as the
temperature increases. 'Ihs is consistent with data in Fig. 111.3 for the tem-
perature dependence of the yield stress. The higher fracture toughness of the
Zircaloy-2 specimens at 77 and 193 K, compared with the room-temperature
value, may be due to the greater extent of deformation twinning that occurs at
lower temperatures. Enhanced ductility of zirconium at subambient tempera-
tures has been attributed to deformation twinning.? Specimens tested at the
lower temperatures will be examined metallographically for evidence of twin
formation.
Fig. 111.5
Equivalent-energy Fracture Toughness as aFunction of Temperature for As-received Zir-
caloy Specimens with Transverse and Longitu-dinal Textures. Neg. No. MSD-64999.
N
E0
CL
H
}
zwz
LJ
90 -
80-
70-
60-
50
IMPACT VELOCITY 1.5 m/s0.14 WT % OXYGEN
O LONGITUDINAL ZIRCALOY-4J TRANSVERSE ZIRCALOY-4 --
0 LONGITUDINAL ZIRCALOY-20 TRANSVERSE ZIRCALOY-2
7
I I I I40 0 100 200 300 400 500 600 700
TEMPERATURE ("K)
0
n
I-H
LJ
IMPACT VELOCITY 1.5 m/s
- 0.14 WT % OXYGENA 0.6 WT % OXYGEN0 1.1 WT % OXYGEN
TRANSVERSE
Fig. 111.4
Temperature Dependence of Equivalent Yield
Stress for Hlomogeneous Lirealoy-2 OxygenAlloys with 0.14, 0.6, and 1 .1 w"t %,o Ox ygen.Neg. No. NS[>-G5001.
O A t1 --- r--
34
Figure 111.6 shows the influence of temperature on the dynamic fracture
toughness of homogeneous Zircaloy-Z
3 0 '- ~- - '-~ ~-- - -'-'-
N
E0
a
z1-z
w
IMPACT VELOCITY 1.5 m/sO 0.6 WT% OXYGENA 1.1 WT% OXYGEN
25K
20 -
400 500 600 700 800 900
TEMPERATURE ('K)
Fig. 111.6. Temperature Dependence of Equivalent-
energy Fracture Toughness of -omoge-
neous Zircaloy-2 Specimens with 0.6 and1.1 wt %Oxygen. Neg. No. MSI>64998.
alloys with 0.6 and 1.1 wt % oxygen. Thefracture toughness decreases as thetemperature and oxygen content in-crease, whereas the fracture toughness
of as-received Zircaloy is greater at
higher temperatures (see Fig. 111.5).
Figure III.7 shows the depen-dence of the ductility index on tempera-
ture for homogeneous and compositeZircaloy-2 specimens with different
oxygen contents. The ductility index
increases as the temperature increases,and, as the oxygen content of the homo-
geneous specimens increases, thecurves shift to higher temperature. Ifthe nil-ductility temperature is definedby extrapolation of the curves to zero
on the ductility-index axis, oxygen
clearly increases the nil-ductility tem-
perature. Also, the nil-ductility tem-perature for the composite specimensis lower than that of the homogeneous
alloys with similar oxygen concentrations. This results from the high ductility
of the central transformed-P region of the specimen, which has a low oxygen
content. Additional data are being obtained to determine the effect of oxygen
concentration and distribution on the impact properties of Zircaloy-oxygen
alloys.
Fig. 111.7
Ductility Index as a Function of Temperaturefor Homogeneous and Composite Zircaloy-2Specimens with Several Oxygen Contents.Neg. No. MSD-64995.
x
2
HA
V
C
9I. 1 _IMPACT VELOCITY 1.5 m/s
8 -WT % OXYGEN
0 0.147 0 0.6 HOMOGENEOUS
v 1.1
6 - } 'COMPOSITE
5
4 -
3
2
00 100 200 300 400 500 600 700 800
TEMPERATURE ( K)
15F -
35
C. Integral Tube-burst/Thermal-shock Properties of Oxidized Zircaloy-4
Cladding
1. Failure of Oxidized Zircaloy-4 Cladding by Thermal Shock
Reference I describes the experimental methods used to determine
the capability of Zircaloy-4 cladding to withstand thermal shock by bottom-
flooding with water after rupture under internal pressure and oxidation in steam
for various times at temperatures between 4170 and 1500 K. Additional data
have been obtained at temperatures between ~1300 and 1580 K to define the
oxidation time that results in thermal-shock failure during bottom flooding of
the tubes with water. The results in Fig. 111.8 indicate that isothermal oxidation
periods greater than ~10, 2, 0.7, and 0.3 ks at 1300, 1400, 1500, and 1600 K,
respectively, will result in failure of the tubes. Each datum point represents
OXIDATION TEMPERATURE ('G)1300 1200 oo 1000
0 CLADDING INTACT
0V
0
FAILED UPON QUENCHINGSURVIVED QUENCHINGBUT FAILED DURING 0HANDLING 0 00
".00 00
Soo 0S Ooo(o 00
" 00
" ocpo 0 O J0
4f 0 c0 0na99
" 0 0 0
"" 0 020 000
CD 0000 O
0 0
0 0
2h-
6.0 6.5 7.0 7.5
Fig. 111.8. Capability of Zircaloy-4Withstand Thermal shodto 430 K by Bottom-flWater and "Normal HaRoom Temperature afterIsothermal Oxidation inVarious Times at Tempetween 1140 and 1550K.No. 306-77-598 Rev. 1.
00
0
00-
0
0(o1
the constant temperature of the tubesection during oxidation. The figure
does not include results for tubes in
which the temperature in the hottest
portion of the tube varied by more than
7 C. Since oxidation periods required
to produce thermal-shock failures at
temperatures <l300 K are several
orders of magnitude greater than those
expected in hypothetical LOCA situa-
tions in LWR's, the failure boundary
will not be established quantitatively.
However, additional data are being
obtained at temperatures between ~1550
and 1750 K.
The amount of oxidation in eachtube was determined by measuring
0the thicknesses of the oxide, a', andtransformed-p layers depicted inscanning-electron micrographs offracture surfaces of failed specimens
6.0 8.5 and optical micrographs of polishedcross sections of the intact tubes. This
Cladding to information was used to calculate the
k from ~810 equivalent- cladding- reacted (ECR)coding with parameter, i.e., the total thickness ofndling" at oxide on the cladding if all the oxygenRupture and in the a and transformed-p layers wasSteam for
ratures be- converted to stoichiometric ZrO2 -ANL Neg. Figure III.9 is the failure map relative
to the ECR parameter and the oxidation
4
2
104a
6
4
6
4
Ai
V
W.
1-
<iNW
I - I -I--- - -
C
ki
36
temperature in steam. No failures occurred for specimens oxidized to 17% of
the total cladding thickness, based on the ECR parameter, and temperatures
up to 1477 K, which are the present oxidation limits in the acceptance criteria
for emergency core-cooling systems in LWR's.
ISOTHERMAL OXIDATION TEMPERATURE (C)
900 1000 i100 1200 1300 1400 1500100 -- T T - -T _ _1 -
90 o CLADDING INTACT" " FAILED ON QUENCHING
80 -" SURVIVED QUENCHING BUT
S70 FAILED DURING HANDLINGW7 o-0
0 60 o "-Z
0 50--
40 -17% LIMIT "ii
W 30- 0
o 0 20 --. Q1 o o o
W 10L - g 1477K LIMIT
L L --- 1. 1 1 1 -1200 1300 1400 1500 1600 1700
ISOTHERMAL OXIDATION TEMPERATURE (K)
Fig. 111.9. Failure Map for Zircaloy-4 Cladding by Thermal Shock or Normal
Handling Relative to ECR Parameter and Oxidation Temperatureafter Rupture in Steam. ANI Neg. No. 30-78-213.
2. Measurements of the Leidenfrost Temperature during Quench of
Oxidized Zircaloy-4 Tubes by Bottom-flooding with Water
Temperature-versus-time profiles were obtained during the tube-
burst thermal-shock tests from several thermocouples (250-p m-dia Pt-
Pt 10 wt % Rh wire) attached to the outer surface of the cladding at various
axial positions. Figures I1.10 and III.11 show typical time-temperature rec-
ords for two specimens ruptured in steam at ~1040 K (heating rate of 10 K/s)
and oxidized at maximum temperatures of 1473 (for 1.38 ks) and 1550 K (for
0.78 ks). Figures 111.12 and 111.13 show the thermocouple locations and the
types of fracture that occurred in the tubes (temperature-time profiles corre-
spond to Figs. III.10 and III.11, respectively) after slow cooling (-5 K/s) from
the oxidation temperatures to ~980 and ~1100 K, respectively, and bottom-
flooding with water at rise rates of -2.5 and ~6.5 mm/s. Under the low rates
of water rise, the quench front moves up the tube in a uniform manner, i.e.,
as indicated by the response of thermocouples 1-4 in Fig. 11.10. A transition
in the heat-transfer mode from film boiling to nucleate boiling is responsible
for the abrupt decrease in the cladding temperature at -650-790 and 750-840 K
in Figs. 111.10 and 111.11, respectively.
16 -
14 - CHART SPEEDCHANGE
12 -- 5
-0 -- ' TRANSFORMATION BOTTOM-F0DING -1RANGE N /BOTTOM-FODN
8 - CLADDING V6- BURST6-K-- #-4
N>-
TIME SCALE CHANGE2 TC*I #2
0 -L LW i-18 -__-_0 600 1200 1800 1962 1986 2010 2034 2048 2072 2096
TIME (s)
Fig. III.10. Temperature-vs-Time Curves from a Tube-burst/Thermal-shock Test That Show an Abrupt Changein Cooling Rate at Temperatures between ~650and 790 K. The Zircaloy-4 tube was ruptured insteam at ~1040 K , oxidized at 1473 K for 1.38 ks,slow-cooled t'980 K, and bottom-flooded withwater at a rise rate of ~2.5 mm/s. ANL Neg.No. 306-78-216.
13 - K- CHART SPEED
CHANGE
B0TTOM- FLTC # 2 VALVE OPEN
TC #4 -
r CLADDING BURST -
%TIME5 /SCALE TC 5
CHANGE
3 - - ! - --. i- 1 i 1- ll_0 120 240 360 480 600 720 840 960 1080 1144 1168 1192 1216
TIME (S)
oo700
LiJ30
100z
500
-j1600
;1500TODINGN -- 1400
1-~300TC#6
1200
T 100
-' 900
800
700
1240 1264
Fig. III.11. Temperature-vs-Time Curves from a Tube-burst/Thermal-shock TestThat Show an Abrupt Change in the Cooling Rate at Temperatures be-tween ~750 and 840 K. The Zircaloy-4 tube was ruptured in steam at~104C :, oxidized at 1550 K for 0.78 ks, slow-cooled to ~-1100 K, andbottom-flooded with water at a rise rate of ~6.5 mm/s. ANL Neg.No. 306-78-212.
37
R-
-JCL
0O
U.'
E
00
1-
Y
0
W
a
crW
CLNW
1--
U
Z
D
a
J
U
,TC #5
o
C -
JG -
O -
G
T -
O -
C ""
G
G
oGo-
4
C "
G -
C '
G
O
J O
O
~ O
O
m
#4
#2
Wn
1'lg. [11.1:'
Zi rca\-'1 I.lad dding afitr Thitrnia1- ihock FaiI-
urc Slowinu lo.'ai n of Tlhlrm.o'liplc That
Pioduicd ilit l'mpcralulre-v -T1inm. ('I: rvc ill
Fig . III.1.0. ANI Nc.. No. 2C0-78-2i:
TC #6
CGo
4 * 00# 5
-#4
#2
*
I'il. I11.1:i
Zira -dincd -It 1c riipt' aftir F Ii niat. a i rhck Fa il-
u1r' Shoingll Location of Th1'1erm (oIoupI ' Thal
P'roduic'cd the'I tl) T Imp lrturl'-Vsi-T m I :re I ll I nV'
Figure 111.14 shows the tine-trnperature response of two the rrno-
couples on a Zircaloy-4 tube that was ruptured in steam at -1050 K, oxidized
at 1273 K for 7.2 ks, and quenchecd from 1273 K by interrupting the power and
500
500
TIMF AF ; .TART OF BoTu TrM H 001DiNG ()
Fig. III.14. Tcmpcraturc-vs-1'imc curves from a Tlube-burstTI'hcrmal-shck
Test in Which the Tube Was Quenched from 1273 K iV Bottom-
flooding with Water at a Rise Rate of 50 mm/s. Ihe abrupt changein the cooling rate occurred at temperatures between 784 and 841 K.ANL Neg. No. :306-78-214.
38
39
bottom-flooding the 200-mm-long tube within 3 s. In this case, violent boilingoccurred, and a distinct boundary between the water level and the two-phase
mixture could not be observed. The temperature-time profiles, recorded by
a high-speed Visicorder, indicate that the transition temperature from film to
nucleate boiling, i.e., the Leidenfrost temperature, 3 ' 4 ranged from -780 to 840 K
in this experiment. Based upon the results from numerous tube-burst/
thermal-shock experiments, we conclude that the Leidenfrost temperature for
heavily oxidized Zircaloy cladding is in the range of 650-840 K, and the cooling
rate of the cladding during the quench is -970 K's. Since the Leidenfrost tern-
perature increases with an increase in pressure, 5 the quench temperature for
Zircaloy cladding during the reflood stage of hypothetical LOCA's (systempressure of ~0.4 MPa) may be slightly higher than the values determined in
our experiments at a system pressure of ~0.07 MPa.
The information on the Leidenfrost temperature and the quench rate
will be used in an analysis' of crack growth in oxidized Zircaloy cladding during
thermal-shock conditions and to define the temperature range for fracture-toughness measurements on oxidized and homogenized Zircaloy specimens.
3. Oxidation Characteristics of Inner and Outer Surfaces of Zircaloy-4
Cladding in Steam
In Ref. 1, local variations in the thickness of oxide layers on the
inner and outer surfaces were reported for Zircaloy-4 tubes that were ruptured
in steam and oxidized for extended periods of time (2 ks) at temperatures be-
tween 1170 and 1300 K. The anonalous oxidation behavior, which could not be
attributed to circumferential temperature differences in the cladding, was cha r-
acterized by a porous white oxide layer (under polarized light) in contact with
the normal columnar oxide at the outer surface and a comparatively thick scale
at the inner surface with numerous short wavelike cracks in the circumferential
direction.
The white oxide at the outer surface is similar in many respectsto the oxide layers that form on zirconium and zirconium alloys during break-
away oxidation. 7 Figure 111.15 shows regions of the fracture cross section ofZircaloy tubes that were ruptured under internal pressure in steam, heavilyoxidized at 1270-1370 K, and quenched by bottom-flooding with water. The
SEM fractograph in Fig. III.15A shows the initial stage of crack formation inthe oxide at the inner surface after the normal columnar oxide attains a thick-
ness of ~0.035 mm. Figure III.15B shows a region of anomalous oxide growthat the inner surface. Apparently, the thick porous scale forms by repeated
cracking when the adherent oxide attains a critical thickness. Figure III.15C
shows regions of anomalous oxide growth at the inner and outer surfaces in
which the initial stages of cracking occurred at oxide thicknesses of 0.019 and0.035 mm, respectively. Since the wavelike cracks are always parallel to L:eoxide/ce interface, a tensile stress in the radial direction apparently initiates
the cracks. (Note the absence of cracks in the external region of the scale atboth surfaces, i.e., near the oxide/steam interface.)
A
On
40
PREV IOUS
BET A
f P HA-
1I.PHIA
ID OXIDE
(A )
Fig. IlI.L5
SE. Fractographs Showing Anomalous Oxidation ofZircaloy-4 Tubes Ruptured under Internal Pressure,
Oxidized at 1270-1370 K, and Quenched by Iottom-
flooding with Water. (A) Shows the initial stage ofcrack formation in the oxide layer at the inner (If))surface when the columnar oxide layer is ~0.035 mmthick; (B) shows a region of anomalous oxide layergrowth at the inner (In) surface that results from re-
peated cracking of the oxide layer when a criticalthickness is attained, and (C) shows regions of anom-alous oxide at the inner (II)) and outer (OD) surfaces
in which the initial stages of cracking occurred at
oxide thicknesses of -0.019 and 0.035 mm, respec-tively. ANL Neg. Nos. 30G-78-221, -222, and -220.
200 y m
It)A.
-
50pm,
OD OX IDE
4V ALPHA
PREVIOUSBETA
1 ALPHA
I D OXIDE
(C)
j**.
.. ,
41
Figure 111.16 shows a more complete characterization (than re-
ported in Fig. 111.27 of Ref. 1) of anomalous oxide-layer formation relative to
OXIDATION TEMPERATURE (C)
1300 1200 1100 1000 900
O ANOMALOUS POROUS OXIDE OBSERVED AT ID 8 OD- , ANOMALOUS POROUS OXIDE OBSERVED AT IDO OLY- 0 ANOMALOUS POROUS OXIDE NOT OBSERVED
@ @@
000c;e e so06 CO 00 C
C o0oo jo o o Do
o o b 0 " 00 00o o o O0
Or ?9;o C-10
Q n
2
102 -60 6.5 7.0 75
I/T X10 (I)80 85
Fig. III.1. Time and Temperature Condi-
tions Resulting in AnomalousOxidation at Inner and OuterSurfaces of Zircaloy-4 TubesRuptured at ~1050 K in Steamat a Heating Rate of 10 K /s and
an Internal Pressure of 6.89 MPa.
Pellet-cladding axial-gap dis-tance was 2.5 mm. ANL Neg.No. 30(-77-G03 Rev. 1.
10
e
6
4
2
A comparison of the time and temperature conditions required for
the formation of an anomalous oxide layer at the inner surface of pellet-
constrained (12.5-mm axial gap) and unconstrained tubes in Figs. 111.17 and
111.18, respectively, with results in Fig. 111.16 indicates that the time for the
onset of anomalous layer growth increases as the degree of axial constraint
decreases. For example, at 1258 K, anomalous oxide formation occurred at
11, 4, and 2 ks for unconstrained tubes and tubes with pellet-cladding axial-gap
distances of 12.5 and 2.5 mm, respectively.
the oxidation temperature and time
for the experimental data in Fig. 111.8.
The data indicate that the anomalous
porous oxide forms at the inner and
outer surfaces at oxidation times : . 5
and 7 ks, respectively, and tempera-
tures near 1270 K. For oxidation
times il ks, no anomalous oxide-layergrowth was observed on either surface
of the tube between ~1200 and 1600 K.
In addition to the time and tem-
perature of oxidation, microstructural
results (Figs. 111.28 and 111.29 of Ref. 1)
indicate that local stress in the clad-
ding also appears to be a factor in
anomalous oxide-layer formation at
high temperatures. To evaluate this
hypothesis, several Zircaloy-4 tubes
with a larger pellet-cladding axial-gap
distance (12.5 versus 2.5 mnm) and
several unconstrained tubes (empty
tubes) were ruptured in steam and
oxidized under conditions similar to
those in the previous experiments
(Fig. 111.16). As the degree of axialconstraint decreases, axial contraction
and bending of the tube occur to a
greater extent during rupture at tem-
peratures in the o-phase region (i.e.,
<1083 K). Therefore, regions of the
tube with high localized stress areless likely to occur.
I0 K40
2
F-
z0C
0
0
1o'
8
6
4
H
6
4
3
2
8
6
4
3
2
1038
6
4
3
OXIDATION TEMPERATURE (C)
1150 1100 1050 1000 950 900 850
7.0 7.4 7.8 8.2
104 - T1 (K-')
8.6 9.0
86
4
3-"|"" -
- I I I I I I I -
" 0 0 0 al.- 00
0-0 0 -
0 aoo o
c600oC0 o0o 00
" ANOMALOUS POROUS OXIDE OBSERVED-- AT ID ONLY
0 ANOMALOUS POROUS OXIDE NOT -OBSERVED
- -I
2F
10486
4
3
2
103
8
6
4
3
2
Fig. 111.17
Time and Temperature Conditions Resulting inAnomalous Oxidation of Inner Surface of Pellet-constrained Zircaloy-4 Tubes (12.5-mm axialgap) after Rupture in Steam (~1050 K) at a
Heating Rate of 10 K /s and an Internal Pres-sure of 6.89 MPa. ANL Neg. No. 306-78-217.
OXIDATION TEMPERATURE ( C)1150 1100 1050 1000 950 900
7.0 7.4 7.8
10 4"-T~'
8.2
( K-1 )
850
8.6 9.0
Fig. 111.18
Time and Temperature Conditions Resultingin Anomalous Oxidation of Inner Surface ofUnconstrained Zircaloy-4 Tubes after Rupturein Steam ('1050 K) at a Heating Rate of10 K /s and an Internal Pressure of 6.89 MPa.ANL Neg. No. 306-78-218.
These results indicate that the conditions during deformation and
rupture at temperatures in the a-phase region have a significant effect on
anomalous oxide-layer growth at higher temperatures and oxidation times
?2 ks. Additional evidence concerning the role of deformation on the subsequent
oxidation behavior of the cladding was obtained from a comparison of the oxida-
tion characteristic of undeformed cladding (Fig. III.19B) with the previous re-
sults for tubes ruptured in steam. No anomalous oxidation behavior was
observed at either the inner or outer surfaces after oxidation at temperatures
between 1220 and 1330 K and times to ~-21.6 ks (Fig. 111.20).
In view of the results in Figs. 111.16-111.18 and 111.20, it is not sur-
prising that Westerman et al.8 could not reproduce the results of Furuta et al. 9
for anomalous oxidation at the inner surface of Zircaloy tubes oxidized at
1273 K for 300-420 s.
42
4N
IL.4
0
Q-4
H
CLJ
z0
0
0W-4
00 0
00 00 00 0 0 0 0
0 ANOMALOUS POROUS OXIDE NOTOBSERVED
" ANOMALOUS POROUS OXIDE OBSERVEDON ID SURFACE ONLY
_ ~~
L 1
TOP FREE
DIAMETRAL GAP - --
0.2 mm
Al2 03
PELLETS
RUPTUREDAND BENT
(A)
THERMOCOUPLE
CIRCULAR HOLESFOR STEAM INLET
NO PELLETSNO PRESSURIZATION
(B)
Fig. 111.19
Schematic of Zircaloy Tube Geometriesfor Investigation of Anomalous OxidationBehavior in Steam at High Temperatures.(A) Pellet-constrained tubes ruptured undertransient-heating conditions and (B3) unde-
formed tubes with a 2.2- or 6.2-mm2 holefor steam access to inner surface. ANLNeg. No. 306-78-215.
BOTTOM FIXED
7
5
3
Fig. 111.20
Time and Temperature Conditions, Similar toThose in Figs. 111.17 and 111.18, That Do NotResult in Anomalous Oxide-layer Growth onUndeformed Zircaloy-4 Cladding. ANL Neg.No. 306-78-219.
Q.2U
0
0
I-
2
104
7
5
3
2
7
5
3
05
OXIDATION TEMPERATURE ( C)
1150 1100 1050 1000 950 900 850
0 ANOMALOUS POROUS OXIDE NOTOBSERVED
~ $~STEAM INLET AREA 2.2 mmr 2
~ N STEAM INLET AREA 6.2 mm2
- -
7.0 7.5 8.0
104 .T 1- ("K-1 )
8.5 9.0
43
7. ,
44
References
1. Light-water-reactor Safety Research Program: Quarterly Progress Re-
port, July-September 1977, Sec. III, "Mechanical Properties of Zircaloy
Containing Oxygen," ANL-78-3.
2. A. M. Garde, E. Aigeltinger, and R. E. Reed-Hill, Relationship betweenDeformation Twinning and the Stress-Strain Behavior of Polycrystalline
Titanium and Zirconium at 77 K, Met. Trans. 4, 2461 (1973).
3. J. G. Leidenfrost, De Aguae Communis Nonnullis Qualitatibus Tractatus,
Duisburg (1756).
4. L. S. Tong, Boiling Heat Transfer and Two-Phase Flow, John Wiley and
Sons, Inc., New York (1967).
5. S. Yao and R. E. Henry, Effect of Pressure and Materials to the Quenching
of Solids, Trans. Am. Nucl. Soc. 27, 611 (1977).
6. Light-water-reactor Safety Research Program: Quarterly Progress Re-port, April-June 1977, Sec. III, "Mechanical Properties of Zircaloy Con-
taining Oxygen," ANL-77-59.
7. T. Ashmed and L. H. Keys, The Breakaway Oxidation of Zirconium and Its
Alloys - A Review, J. Nucl. Mater. 39, 99 (1975).
8. R. E. Westerman, G. M. Hesson, and C. R. Hann, Zircaloy Cladding ID/OD
Oxidation Studies, Battelle Pacific Northwest Laboratories, EPRI NP-525
(Nov 1977).
9. T. Furuta, M. Hashimoto, T. Otomo, S. Kawasaki, and K. Homma, Defor-mation and Inner Oxidation of the Fuel Rod in a Loss-of-coolant Accident
Condition, Japan Atomic Energy Research Institute, JAERI-M6339
(Nov 1975).
