Journal of Nuclear Science and Technology
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Thresholds for failure of high-burnup LWR fuels
by Pellet Cladding mechanical interaction under
reactivity-initiated accident conditions
Yutaka Udagawa, Tomoyuki Sugiyama & Masaki Amaya
To cite this article: Yutaka Udagawa, Tomoyuki Sugiyama & Masaki Amaya (2019) Thresholds
for failure of high-burnup LWR fuels by Pellet Cladding mechanical interaction under reactivityinitiated accident conditions, Journal of Nuclear Science and Technology, 56:12, 1063-1072,
DOI: 10.1080/00223131.2019.1637795
To link to this article: https://doi.org/10.1080/00223131.2019.1637795
Published online: 03 Jul 2019.
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JOURNAL OF NUCLEAR SCIENCE AND TECHNOLOGY
2019, VOL. 56, NO. 12, 1063–1072
https://doi.org/10.1080/00223131.2019.1637795
ARTICLE
Thresholds for failure of high-burnup LWR fuels by Pellet Cladding
mechanical interaction under reactivity-initiated accident conditions
Yutaka Udagawa
, Tomoyuki Sugiyama and Masaki Amaya
Nuclear Safety Research Center, Japan Atomic Energy Agency, Ibaraki, Japan
ABSTRACT
ARTICLE HISTORY
To contribute to the future updating on the Japanese safety criteria for pellet/cladding
mechanical interaction (PCMI) failure of light water reactor fuels under reactivity-initiated
accident (RIA) conditions, this paper summarizes the recent important outcomes from
research programs with the Nuclear Safety Research Reactor (NSRR). Applicability of current
criteria, which are defined as a function of fuel burnup and possibility of introducing another
parameter for new criteria were evaluated based on the results of the RIA-simulated pulse
irradiation tests, post-test examinations, and supporting analytical work, such as the reevaluation of fuel enthalpies in earlier NSRR experiments. Failure-threshold curves based on
cladding hydrogen content as a primary measure of fuel degradation have been proposed as
a possible alternative that can be used to judge the occurrence of PCMI failure to ensure
conservativeness in a more pertinent manner.
Received 25 March 2019
Accepted 25 June 2019
1. Introduction
Reactivity-initiated accidents (RIAs) are categorized
as a design basis accident of light water reactors
(LWRs) in many countries, and the regulatory bodies
have formulated their regulatory acceptance criteria
to prevent potential consequences of RIAs, such as
damage to the reactor pressure vessel.
In 1998, the Japanese Nuclear Safety Commission
(NSC) revised the licensing criteria to evaluate the
number of failed fuel rods under RIA conditions [1].
As part of the revision, the NSC added a new fuelrod-failure threshold, namely a decreasing function of
fuel burnup referred to as ‘pellet cladding mechanical
interaction (PCMI) failure criteria’ (see the red line in
Figure 1), since the results of RIA-simulated experiments in 1990s had revealed that highly irradiated
fuels could fail at lower enthalpy increases than that
had been previously assumed [2]. While the revised
acceptance criteria are valid today in Japan, certain
technical issues were left to be addressed through the
extension of the RIA-test database and further discussions, particularly with respect to the PCMI failure
criterion for the burnup range from 65 to 75 GWd/
tHM, which was determined by engineering
judgment.
The Japan Atomic Energy Agency (JAEA) has
since performed extensive research programs thereafter to better understand the transient behavior of
LWR fuels that can be reflected in regulatory judgments. Pulse irradiation tests on advanced LWR fuels
irradiated in commercial reactors were conducted at
CONTACT Yutaka Udagawa
udagawa.yutaka@jaea.go.jp
Naka-gun, Ibaraki, 319-1195, Japan
© 2019 Atomic Energy Society of Japan. All rights reserved.
KEYWORDS
BWR; PWR; PCMI; failure;
fuel; high burnup; RIA;
hydrogen embrittlement;
cladding; zircaloy
the Nuclear Safety Research Reactor (NSRR) primarily in the framework of the Advanced LWR Fuel
Performance and Safety (ALPS) research program
(2002–2009) and in the subsequent ALPS-II program
launched in 2010 [3–10]. The NSRR has a narrow
power pulse and only supports stagnant-coolant system for pre-irradiated-fuel tests, both of which are
apart from the RIA conditions anticipated in LWRs.
A rapid energy injection, however, generally hinders
ductility enhancement of fuel cladding associated
with temperature increase and thus produces conservative test results in terms of PCMI-failure limit.
Besides, the timescale of PCMI is too short to be
affected significantly by fuel-rod-surface heat transfer.
Hence, such differences between the NSRR and
LWRs have been believed to be rather minor factors
in studying the PCMI-failure phenomenon as a safety
issue. These experimental programs have so far added
more than 20 data points to the RIA-test database
and associated post-test examinations and analyses
have resulted in the derivation of important conclusions, regarding effects of fuel initial temperature in
the transient, cladding morphology, fuel burnup, and
cladding irradiation damage with respect to the fuel
failure limit, which had not been clear at the last
revision. Part of the latest ALPS-program output has
been already incorporated into two new acceptance
criteria recently established in the USA and in France
[11,12].
The purpose of this paper is to summarize the test
results, discuss the influence of the updated
Nuclear Safety Research Center, Japan Atomic Energy Agency, Shirakata 2-4, Tokai-mura,
1064
Y. UDAGAWA ET AL.
BWR
PWR
NSRR
ATR
SPERT CABRI
JMTR &PBF MOX UO2
(a)
(b)
failure
200
800
100
400
PCMI failure criteria
0
0
20
40
60
Burnup [GWd/tHM]
80
Fuel enthalpy increase [cal/g]
Fuel enthalpy increase [J/g]
non−failure
0
Figure 1. The current Japanese PCMI failure criteria with the
existing data of dHfail and dHmax for failure and non-failure
cases, respectively, as a function of fuel burnup.
knowledge on the current Japanese licensing criteria,
and consider possible alternatives to the current criteria as a more accurate measure for evaluating fuel
safety performance under RIA conditions. Chapter 2
summarizes the primary results of the NSRR tests
performed after 1998. The chapter also recalls the
reevaluation of fuel enthalpies in the earlier NSRR
tests [13] with some modifications, because the discussion of fuel-rod-failure threshold in the subsequent part is completely dependent on the values of
fuel enthalpy. Chapter 3 then discusses the applicability of the current Japanese PCMI failure criteria in
the burnup range from 65 to 75 GWd/tHM for each
of the conventional and the advanced fuels after all
the updates of the RIA-test database were considered.
The discussion derives expected revisions to the criteria with the current approach and its possible
alternatives.
2. NSRR experiment for high-burnup fuels
2.1. Fuel and test conditions
In the ALPS programs, RIA tests have thus far been
performed on UO2 and mixed-oxide (MOX) fuels
which had been irradiated to 45–84 GWd/tHM
(local burnup) mainly in European commercial reactors [3–10]. The fuels were cut into short test rods
with 50–117 mm fuel stack length at the Reactor Fuel
Examination Facility (RFEF), placed and instrumented in a test capsule, and then subjected to pulse
irradiation tests at the NSRR, which enabled a rapid
energy injection of fission reaction into the test rods
by its sharp power pulse with a full width at half of
the maximum of 4.4–5.5 ms. Detailed pre- and posttest examinations on the tested fuels were performed
at the RFEF, including visual and dimensional inspections, axial gamma scanning, X-ray radiography,
Figure 2. Schematics of the test capsule used for NSRR test:
(a) Room temperature capsule and (b) High-temperature
capsule.
metallographic observations, SEM observations, gas
analysis, etc.
Figure 2(a) and (b) illustrate two different types
of test capsules used for the pulse irradiation tests
performed with stagnant coolant water at (a) room
temperature (RT, ~20ºC) and atmospheric pressure
(~0.1 MPa) and (b) high temperature (HT, ~280ºC)
and high pressure (~6.6 MPa), respectively, around
saturation levels with small sub-cooling less than 10
K. Both types of test capsules adopted a doublecontainer system, fabricated from stainless steel to
ensure airtightness designed to prevent the release
of radioactive material. The internal volumes of
their inner capsules, which had an inner diameter
of 120 mm corresponding to the test section and
contained the instrumented test-fuel rod and coolant water. The fuel-stack length of the test rod was
limited to about 50 mm in the HT capsule because
of its smaller test section in the axial direction.
Strain-gage-type pressure sensors were installed at
the bottom of the test section and acoustic emission sensors were installed on the test fuel’s supporting structure in the vicinity of the test-fuel
bottom part. They detect the pressure waves and
sound waves, respectively, released at the instant of
PCMI failure as a sharp increase in the amplitude
of the output signal, which facilitated a precise
determination of failure time. Detailed descriptions
of the tested fuels, test capsules, test conditions,
and other procedures of NSRR tests for highburnup fuels have been previously reported elsewhere [2].
Table 1 summarizes the test results of the
ALPS program from 2003 to 2017 and it picks
up key data to discuss the PCMI-failure limit of
high-burnup fuels. In the table, the test-rod
burnup was an averaged value for the fuel stack
JOURNAL OF NUCLEAR SCIENCE AND TECHNOLOGY
1065
Table 1. Summary of NSRR test results performed after 2003.
Fuel enthalpy increasee (J/g)
Cladding
Test IDa
VA-1
VA-2
VA-3
VA-4
MR-1
RH-1
RH-2
BZ-1
BZ-2
BZ-3
BZ-4
LS-1
LS-2
LS-3
DW-1
VA-5
VA-6
VA-7
VA-8
GR-1
Test rod burnup
(GWd/tHM)
71
77
71
77
71
67
67
48
59
59
59
69
69
69
45
81
78
81
78
84
Material
Oxide thick.b (μm)
H cont.c (wtppm)
Initial coolant temp.d
Peak: dHmax
At failure: dHfail
TM
73
70
82
80
39
6
6
30
20
20
20
25
25
25
10
30
60
33
68
15
660
760
670
760
210
70
70
340
160
160
140
300
290
300
50
312
683
318
703
95
RT
RT
285°C
249°C
RT
RT
278°C
RT
RT
281°C
RT
RT
283°C
281°C
RT
RT
RT
281°C
282°C
RT
556
546
454
457
408
462
378
688
644
528
451
469
371
390
505
587
617
483
500
452
268
231
344
318
545
222
301
146
175
-
ZIRLO , SR
MDATM, SR
ZIRLOTM, SR
MDATM, SR
NDATM, SR
M5TM, RX
M5TM, RX
Zry-4, SR
Zry-4, SR
Zry-4, SR
Zry-4, SR
Zry-2(LK3), RX
Zry-2(LK3), RX
Zry-2(LK3), RX
Zry-2, RX
M-MDATM, SR
M-MDATM, RX
M-MDATM, SR
M-MDATM, RX
M5TM, RX
a
All the tests were performed with stagnant coolant water and rod-filler gas of 1 atm He.
Absolute uncertainty is estimated to be less than 1 μm.
Relative uncertainty is estimated to be less than 3%.
d
Absolute uncertainty is estimated to be less than 3 K.
e
Relative uncertainty is estimated to be less than 10%. (See Section 2.2)
b
c
of the test rod, estimated from the average
burnup and gross gamma-scanning profile of
each mother rod, which was provided by the
fuel vendors. The oxide thickness provided in
the table was determined by taking the arithmetic
mean of the oxide-thickness values measured in
metallographic observations on the horizontal
cross sections of the cladding at the top- and
bottom-side vicinities of the test-rod section. An
essentially linear profile for the oxide thickness
within the axial section of the test-rod fuel stack
was confirmed for each test rod by eddy current
measurement, where oxide thickness is supposed
to correlate with the measured distance between
eddy current probe and the outer surface of the
conductor, namely cladding metallic layer. The
hydrogen content was also calculated as the arithmetic mean of the measurement results by intergas fusion thermal conductivity detection, while
information of its axial profile, such as the eddy
current measurement used to support the determination of oxide thickness, was not available to
estimate hydrogen content. The specimens subjected to the destructive examinations were
extracted from mid-pellet positions.
All the test cases with the values of ‘dHfail’ indicated in the table were those identified as of PCMI
failure cases. The evaluation of peak fuel enthalpy,
dHmax, was conducted based on the total amount of
fissile materials evaluated by mass analysis (MA) and
computation by the neutron transport calculation
model for the NSRR reactor core, referred to as
‘PROC2’ [13] and described in the next section.
2.2. Evaluation method of fuel enthalpy
The evaluation method, reported as ‘PROC2’ in our
previous work [13], was based on the total amount of
fissile materials (U-235, Pu-239, and Pu241) in the
test-fuel rod, which was quantified primarily by MA.
Elemental and isotopic compositions of the test fuel
were evaluated from a set of burnup calculations with
the SWAT code [14] that adopted the neutron spectrum adjusted to reproduce the fissile ratio found by
MA and also take account of base irradiation history,
coolant temperature at the MA sampling elevation,
and the cooling period after unloading. The fission
number density per unit NSRR power, RPROC2, was
then evaluated by neutron transport calculations
using the TWODANT code [15,16], which takes the
configurations of the NSRR core, test capsule, coolant, and the test-fuel rod with the isotope composition obtained in the aforementioned processes into
account, with a two-dimensional (R-Z) geometry
description. Finally, dHmax was calculated by the use
of several empirical correlations formulated by
Fuketa et al. [17] based on the evaluation by
Ohnishi and Inabe [18]:
dHmax
Qinf 1 Qp
;
¼ Qp ¼ Qinf
Q1
Q1
(1)
Qinf =Q1 ¼ 5:55847 101 þ 6:53146a 7:90842a2
þ4:78526a3 1:63376a4 þ 3:20404 101 a5
3:37394 102 a6 þ 1:48040 103 a7 ;
(2)
1066
Y. UDAGAWA ET AL.
Qp =Q1 ¼ 4:55351 þ 1:09706 101 a 9:51210a2
þ4:52288a3 1:26063a4 þ 2:04792 101 a5
1:78692 102 a6 þ 6:42444 104 a7 ;
(3)
Qinf ¼ NVTinf RPROC2 qinf ;
Table 2. Specifications of NSRR tests modeled for verification
of PROC2: test-fuel design [13].
U235/
U
Pu/
(U+ Pu)
Gd2O3/
UO2
Pellet outer
diameter
wt%
5
10
20
0.7
10
20
10
20
2.6
10
10
wt%
0
0
0
6.3(74.5a)
10(82.7a)
0
0
0
0
0
0
wt%
0
0
0
0
0
6(98.6b)
0
0
0
0
0
mm
9.29
9.29
9.29
9.30
12.4
8.19
10.57
9.29
10.66
9.29
9.29
Test group
F1
F2
F3
F4
F5
F6
F7
F8
F9
F10
IJc
(4)
NVTinf ¼ 6:78989 101 þ 1:65876 102 a
1:72017 102 a2 þ 1:17553 102 a3 4:66650
101 a4 þ1:07311 101 a5 1:32201a6
þ6:74461 102 a7 ;
(5)
a
1
2
a ¼ þ8:89813 10 þ 3:97948 10 NVT1s
3:47212 104 NVT1s 2 þ 1:42752105 NVT1s 3
2:71172 107 NVT1s 4 þ 2:50669 109 NVT1s 5
1:08531 1011 NVT1s 6 þ 1:64532 1014 NVT1s 7 ;
(6)
1
where Qp [J/g] is the prompt part of the energy
deposited in the test fuel; Qinf [J/g] and Q1 [J/g] are
the energies deposited in the test fuel at an infinite
time and at 1 s after the pulse operation, respectively;
NVTinf [MJ] and NVT1s [MJ] are the NSRR power
integrated to an infinite time and to 1 s, respectively;
qinf is the energy deposited in the test fuel per fission
at an infinite time and assumed to be 186.4 MeV ≈
2.986 × 10−11 J; and a [$] is the inserted reactivity in
the test. In the case of fuel-rod failure, dHfail was
evaluated as
dHfail ¼ NVTf RPROC2 qp ;
(7)
where NVTf [MJ] is the NSRR power integrated to
measured failure time; and qp is the prompt part of
the energy deposited in the test fuel per fission and
assumed to be 175.3 MeV ≈ 2.808 × 10−11 J. The
values for qp and qinf employed above were reported
by Ohnishi and Inabe [18].
The NSRR-core modeling in PROC2 was verified
against the results of another procedure referred to as
Table 3. Specifications of NSRR tests modeled for verification
of PROC2: test capsule design and test condition [13].
Outer capsule
Inner diam.
Test group
F1
F2
F3
F4
F5
F6
F7
F8
F9
F10
IJ
mm
120
120
120
120
120
120
120
120
130
100
130
Inner capsule
Outer
diam.
Inner
diam.
Outer
diam.
Coolant
temp.
mm
134
134
134
134
134
134
134
134
136
126.4
138
mm
65
65
72
120
72
mm
73
73
76.8
126
81
K
RT
RT
RT
RT
RT
RT
RT
RT
RT
549
RT
‘PROC1’ [13] applied to the tests performed at the
NSRR with various test-fuel designs, test capsule
designs, and coolant temperatures, as presented in
Tables 2 and 3. PROC1 is a fully experimental procedure to measure the amount of short-life fission
products (typically Ba-140) which are produced in
fuel matrix during an NSRR test by gamma-spectrometry. The amount of fuel specimen subjected to
PROC1 had been well quantified in the case of the
8
6
11
4
2
0
0
Calculated [10 fissions/g/MJ]
(b)
11
Coupling factor [10 fissions/g/MJ]
(a)
Fissile isotopes in Pu [wt%].
Gd-160 in Gd [wt%].
c
Pre-irradiated fuels in the JMTR.
b
calc.
exp.
IJ
F1,F2,F3
0.1
Fissile ratio [−]
0.2
F6
6
F7
4
F5
F9
2
3
F8
F10
F4
2
1
0
0
2
4
6
Evaluation by PROC1
11
[10 fissions/g/MJ]
Figure 3. Comparison of coupling factors between neutron calculation of PROC2 and evaluation by PROC1 [13].
JOURNAL OF NUCLEAR SCIENCE AND TECHNOLOGY
tests in the verification matrix (Tables 2 and 3), since
they were fresh fuels whose weight was easy to measure, except the IJ group, and the analyzed amount of
the IJ-group fuel had been carefully evaluated by an
MA-based method. Tests listed in the tables were
thus considered to be a reliable part of PROC1 results
and selected as a counterpart of PROC2 which is to
be verified. Figure 3(a) and (b) compare the results by
PROC2 with those by PROC1 in the form of a coupling factor, which is the ratio of heat-generation rate
in the test fuel to that in the reactor core. The relative
error between the results of PROC1 and PROC2 was
less than 7% for the F1 group. The scatter in the
coupling factors evaluated by PROC1 appears to
increase as the fissile ratio increases. The IJ group,
which is the sole pre-irradiated fuel group, exhibited
rather large scatter in the PROC1 results. The coupling factors of the IJ group calculated by PROC2
were still close to the mean value of the scatter range.
From the comparison, we expect the uncertainty
involved in the NSRR-core modeling adopted in
PROC2 is less than 7% for fissile ratio less than 5%,
which is the case of the earlier NSRR experiments
with high-burnup fuels discussed in the next section.
Note that the actual error, namely deviation from the
true coupling factors, of the neutron-calculation
results themselves should be smaller because the scatter which is attributed to PROC1 was not discriminated in the present verification. Taking account of
the uncertainty entailed in the fissile-ratio quantification by MA which is estimated to be about 1–3% thus
1067
far, we expect the overall uncertainty of fuel enthalpy
evaluated by PROC2 is less than 10% at the present.
2.3. Reevaluation of fuel enthalpy in tests before
2003
Regarding the earlier NSRR experiments with highburnup fuels performed before 2003 [19–21], the
discussion of failure limit in this paper is based on
the results of the reevaluation of fuel enthalpies [13],
using the procedure PROC2 described in the previous
section. For several pressurized water reactor (PWR)
fuel cases with low burnup for which MA sampling
was not available, the elemental and isotopic composition was estimated by using the ORIGEN2 code
with PWRUS and PWR34J33 cross-section libraries,
which are included in the original and ORLIBJ33
packages [22,23]. Table 4 shows the results of the
reevaluation by PROC2 with certain modifications
from the previous report. These modifications
included the addition of the reevaluated enthalpies
of the test FK-9, and a small revision to the enthalpies
of the test TK-2, for which necessary correction of
Pu-241 decay effect had been missed. Table 4 also
provides the values of ‘NVT1s’, which complete the
information necessary to compute other important
test-condition parameters, such as the energy deposited in the fuel.
Among the enthalpy-reevaluation results presented in Table 4, the significant differences were
seen between the values of the PWR rods and those
Table 4. Fuel enthalpy increase in earlier high-burnup fuel tests reevaluated using a consistent procedure.
Fuel enthalpy increaseb (J/g)
Fuel enthalpy increase (J/g)
Test ID Integrated NSRR power at 1s (MJ)*
MH-1
68.9 ± 0.6
MH-2
79.6± NA
MH-3
100.0 ± 0.8
GK-1
98.3 ± 0.2
GK-2
94.8 ± 1.6
OI-1
104.2 ± 0.7
OI-2
106.3 ± 0.5
HBO-1
108.5 ± 0.4
HBO-2
57.9 ± 2.0
HBO-3
110.4 ± 0.1
HBO-4
76.3 ± 0.8
HBO-5
108.7 ± 1.2
HBO-6
108.6 ± 0.6
HBO-7
111.9 ± 0.3
TK-1
110.6 ± 0.9
TK-2
111.0 ± 1.0
TK-3
110.8 ± 1.2
TK-4
108.6 ± 0.6
TK-5
111.8 ± 1.1
TK-6
111.4 ± 0.8
TK-7
110.3 ± 0.3
TK-8
61.6 ± 0.3
TK-9
110.6 ± 0.4
TK-10
58.7 ± 3.1
Peak:
dHmax
290
338
415
421
407
603
615
494
252
490
330
534
482
507
753
716
645
654
694
768
673
359
649
409
At failure:
dHfail
354
506
398
607
-
Test ID Integrated NSRR power at 1s (MJ)a
OI-10
77.0 ± 2.2
OI-11
111.3 ± 2.6
OI-12
109.1 ± 1.3
TS-1
73.0± NA
TS-2
81.0 ± 0.6
TS-3
96.1 ± 0.4
TS-4
97.3 ± 0.1
TS-5
106.8 ± 0.2
FK-1
111.0 ± 0.2
FK-2
61.7 ± 0.3
FK-3
109.2 ± 0.2
FK-4
108.9 ± 0.9
FK-5
56.4 ± 0.2
FK-6
110.8 ± 0.2
FK-7
110.9 ± 0.1
FK-8
57.6 ± 0.6
FK-9
78.4 ± 0.6
FK-10
88.0 ± 1.3
FK-12
78.1± NA
Peak:
dHmax
448
674
622
370
367
431
447
460
579
311
626
535
261
522
522
262
363
413
350
At failure:
dHfail
517
271
244
349
316
280
Average value of those from two reactor-signal lines available at the NSRR for integrated reactor power. ‘NA’ denotes the case that valid output was
attained only from one reactor-signal line.
b
Relative uncertainty is estimated to be less than 10%. (See Section 2.2).
a
Y. UDAGAWA ET AL.
(b)
3
PROC2
PROC1(S1−MA)
PROC1(S1−6)
2
3
2
PROC1(S1−MA)
PROC1(S1−6)
y=1.07x
11
11
Coupling factor [10 fissions/g/MJ]
(a)
PROC1 [10 fissions/g/MJ]
1068
1
0
0
0.01
0.02
Fissile ratio [−]
1
y=0.93x
0
0
1
2
3
11
PROC2 [10 fissions/g/MJ]
Figure 4. Comparison of coupling factors evaluated by PROC1 and PROC2. The legends ‘S1–6’ and ‘S1-MA’ denote the two
different quantification methods of fuel-specimen amount subjected to PROC1: based on weight measurement and mass
analysis, respectively [13].
from the earlier evaluations by PROC1 as plotted in
Figure 4. In the figure, the data series with the legend
‘PROC1(S1–6)’ corresponds to the results produced
by PROC1 with its standard methods of quantifying
fuel-specimen amount using direct weight measurement. In particular for the HBO- and TK-test rods,
the discrepancy between ‘PROC1(S1–6)’ and PROC2
was significantly larger than 10%, which is an error
level expected from the comparison in Figure 3. The
reevaluations of the two test series were performed
using results of mass-analyses given in Table 5, performed on the fuel specimens cut out from their
mother rods. Both agreed well with the predictions
by the ORIGEN2 code. The comparison described in
the previous section confirmed the reliability of the
remaining portion of PROC2, namely computation
by the neutron transport calculation model for the
NSRR reactor core. It is important to note that the
two evaluation methods show reasonable agreements
in the comparisons for rather new tests as OI-10 and
OI-12, which is in contrast to the HBO- and TK-tests.
It is also noticeable in Figure 4 that the results by
PROC1 strongly depend on the quantification
method of fuel-specimen amount treated, which
implies a source of large uncertainty in PROC1 at
its application to high-burnup fuels.
The above assessment justifies the use of the
PROC2 results rather than the PROC1 results as the
more reliable evaluation of the fuel enthalpies
referred to in the following discussion of PCMI failure limit in this study, while the PROC2 results
showed a certain level of uncertainty of about 10%
Table 5. Mass analysis results used in the reevaluation of fuel
enthalpies in HBO- and TK-tests.
Sample ID
HBO-MA
TK-MA
Fissile ratio [-]
Pu/U ratio [-]
Mass-analysis date
0.01143
0.01693
0.01164
0.01169
4 April 2013
30 Mar 2010
based on the available data. In the verification matrix
for PROC2, there were a limited number of available
experimental data points whose fuel enrichment was
directly comparable with those of the interested
HBO- and TK-fuels. Specifically, there was only one
data point with fuel enrichment less than 3%.
Additional experimental efforts to obtain such data
of relatively low enrichment fuels (e.g., around 1.0%),
would hence be valuable in a more critical assessment
on PROC2.
3. Summary of PCMI failure database and
discussion
Figure 1 shows the current Japanese PCMI failure
criteria with the existing data points of dHfail and
dHmax for failure and non-failure cases, respectively,
as a function of fuel burnup [1]. The criteria provided
a conservative judgment except for two test cases
including REP-Na1 and VA-6. The criteria treated
the former case as an exception because such excessive corrosion behavior had never been observed in
the nuclear power plants in Japan. The latter may also
be treated as an exception, because the commercial
use of recrystallized (RX) M-MDATM cladding is not
planned.
The threshold shown as the broken line in
Figure 1, 40 cal/g for the 65–75 GWd/tHM range,
was determined by engineering judgment with an
additional safety margin of 10 cal/g applied to the
threshold for the 40–65 GWd/tHM range because of
lacking experimental data at the time the licensing
criteria was revised in 1998. The failure and nonfailure data points obtained in the ALPS tests as
VA-1 to −4, MR-1, RH-1 to −2, and LS-1 to −3
demonstrated the conservativeness of the threshold
up to about 80 GWd/tHM. A majority of the failure
data points above 65 GWd/tHM shows certain margins, namely excess of the failure-enthalpy increases
JOURNAL OF NUCLEAR SCIENCE AND TECHNOLOGY
Recrystalized
HT
RT
Stress−relieved
HT
RT
failure
non−failure
threshold
Fuel enthalpy increase [J/g]
evaluated at the measured failure times compared to
the threshold line, while the absolute values of the
failure-enthalpy increase are on average lower than
those found below 65 GWd/tHM. As a result, the
reduction of the threshold by 10 cal/g was adequate
to absorb the influence of fuel degradation with
burnup extension above 65 GWd/tHM, which has
been actually observed in the ALPS programs. A
possible option in the future revision of the PCMI
criteria, in which the updated knowledge is reflected,
may be hence to simply retain the current approach
with burnup as the primary parameter and keep also
the current threshold curve, including the highburnup range 65–75 GWd/tHM. In this case, however, judgment regarding the applicability of the criterion for 65–75 GWd/tHM would be alloy-specific,
and the experimental data now support this burnup
range for Zry-2, MDATM, M-MDATM(SR), NDATM,
ZIRLOTM, and M5TM, while there is no data for Zry-4
thus far.
The advantage of the burnup-dependent failure
threshold is the simplicity in its practical application.
The possibility of failure for a given fuel rod can be
evaluated directly based on the burnup calculated from
its heating-rate history. On the other hand, it has been
long pointed out that the burnup-based threshold does
not reflect the effect of cladding degradation well as a
result of hydrogen absorption into the metallic layer in
cladding, which is regarded as the primary parameter of
PCMI failure limit, which is discussed below. Since the
amount of hydrogen absorbed in a given fuel burnup
strongly depends on water-side corrosion resistance of
the cladding material and irradiation conditions, the
threshold is generally thought to be more conservative
for the cladding materials with improved corrosion resistance and less conservative for irradiation conditions
with a higher coolant temperature or higher heating
rate, for example. Moreover, the current criteria are not
capable of reflecting another important conclusion from
the ALPS programs, namely the temperature effect. As
demonstrated through the direct comparison of the
results between the tests VA-1 and VA-3 [4], the PCMI
failure limit is now believed to increase as fuel temperature (or more specifically cladding temperature) increases
at the beginning of a reactivity insertion event anticipated
in safety evaluation. The current approach, which bounds
the threshold simply by the lower limit of the failure data
points from both RT and HT tests, is thus thought to be
overly conservative in the case of PWR.
JAEA has thus discussed different failure thresholds based on the hydrogen content absorbed in fuel
cladding as an alternative to the burnup-based
threshold. Figure 5 shows four failure thresholds
with bounding experimental data points, proposed
in the present work. Table 6 presents the hydrogen
content of the test rods subjected to the earlier NSRR
tests considered in the present work. The four
1069
150
a (80,452)
600
d
BZ−3
Na−7
400
VA−3
100
VA−2
200
b
(95,402)
VA−8
VA−6
c (148,235)
0
0
50
250
500
750
Hydrogen content [wtppm]
0
Figure 5. PCMI failure thresholds: fuel enthalpy increase at
PCMI failure expressed as a function of cladding hydrogen
content.
Table 6. Hydrogen content in test rods subjected to MH,
HBO, TK, and OI tests.
Test ID
HBO-1
HBO-2
HBO-3
HBO-4
HBO-5
HBO-6
HBO-7
TK-1
TK-2
H cont. (wtppm)
187
152
148
89
576
331
370
50
196
Test ID
TK-3
TK-4
TK-5
TK-6
TK-7
OI-10
OI-11
OI-12
H cont. (wtppm)
73
125
163
100
138
338
418
355
Relative uncertainty is estimated to be less than 3%.
thresholds are defined for RT and HT, and for cladding types of recrystallization (RX) and stress-relief
(SR) annealing at fabrication, since previous studies
have revealed that the coolant temperature and the
final heat treatment of cladding at fabrication are also
influential parameters on the PCMI failure limit.
Here ‘HT’ curves are intended to show the thresholds
applicable to reactivity insertion events under hotzero-power conditions of PWRs and operating conditions of both PWRs and BWRs. While the potential
influence of irradiation damage accumulated in the
fuel-cladding matrix had also been considered in the
discussion of the definition of the current criteria, the
effect does not appear evident even in the latest RIAtest database. The proposed thresholds only implicitly
take the irradiation effect into account by adopting
experimental results conservatively, as described
below.
The threshold curves were determined as directly
as possible based on dHfail evaluated at experimental
failure time. The basic idea is to form lower-limit
envelops of the failure-limit data thus far obtained
in the RIA-simulated experiments as a function of
hydrogen content by linear interpolation. Taking
Y. UDAGAWA ET AL.
account of the uncertainty involved in the evaluation
processes of dHmax, we have adopted values reduced
by 10% to define the curves from the reference values
that the curves were based on. The margin of 10% is
reasonable considering the error involved in the current procedure for enthalpy evaluation as discussed in
Section 2.2.
The y-axis value of breakpoints ‘a’ and ‘d’ in
Figure 5 is based on the dHfail values from the early
RIA tests as SPERT and PBF, where hydrogen content is not always well identified and/or the failure
mode is not always a typical PCMI failure characterized by an axial split of the cladding tube. Here the
data point from the test SPERT-859 has been
excluded since our analysis suggested that a relatively
low failure limit observed in the test was caused by
excessive corrosion during its base irradiation, in
accordance with other previous analyses by more
simple approach. Thus, the hydrogen content should
have been beyond the range that the proposed threshold curves are defined [24].
The x-axis value of the breakpoint ‘a’ and x- and yaxis values of breakpoints ‘b’ and ‘c’, which define the
low hydrogen content portion of the RX/RT curve,
are not supported by any dHfail data, but by the test
results of dHmax from the non-failure cases FK-4, GR1, and FK-8, respectively. Since the range below 150
wtppm is thought to be practically important for fuel
rods irradiated in commercial BWRs, further efforts,
such as out-of-pile tests and/or numerical analyses, to
support the conservativeness of the threshold curves
are desirable. The x-axis value of breakpoint ‘d’ for
the SR/RT curve corresponds to the hydrogen content of the TK-7 test rod. While the number of the
failure data points directly bounding the SR/HT and
RX/HT curves is small, all the existing data points of
dHmax from non-failure test cases performed under
HT conditions, such as RH-2, LS-2, LS-3, BZ-3, VA4, VA-5 and VA-7 also support the conservativeness
of the two curves.
It is important to note that the ‘hydrogen content’,
taken as the horizontal axis, indicates an average
value for a given axial length of fuel section, neither
a local nor peak value. On the other hand, the experimental dHfail values should reflect the cladding degradation state of a specific axial location at which
hydrogen content should have been relatively high,
because that part should have failed first in the test
rod. It should also be noted that the axial power
profile during the pulse irradiation in the NSRR
reactor core is almost flat. The local hydrogen content
at the failure location is actually not known and solely
the average value is available. This uncertainty may
partly explain the large scatter seen in the plot of the
SR/RT test results in Figure 5. Another important
reason would be the probabilistic nature of the
phenomenon, in particular, the crack formation process in hydride-rim [25].
PCMI loading on the cladding tube for a given
enthalpy increase, which generally depends on the
fuel pellet design, can be explained by fuel thermal expansion for the majority of failure cases
bounding the threshold curves. If UO2 fuel tests
were the sole focus, all the important failure cases
can be explained by fuel thermal expansion, and
the SR/HT threshold could be shifted upward to
some extent based on the result of scaling the
NSRR test results using a fracture mechanics
approach, shown in Figure 6 [26]. When MOX
fuel tests are involved, the impact of additional
PCMI loading by gas-induced fuel pellet swelling
cannot be excluded from the discussion, which is
potentially a cause of the fuel failure observed in
the REP Na-7 [27]. As the recent MOX fuel tests
performed at the NSRR, tests BZ-1 to −4 and
DW-1, did not present clear signs of such MOXspecific behavior, the bounding condition or
mechanism remains unclear. This is one of the
primary reasons that fuel enthalpy must be maintained as the vertical axis of the threshold,
whereas there can be a more direct measure of
loading, like cladding mechanical strain, which
can be correlated with hydrogen content. Note
that the proposed SR/HT curve also gives a conservative judgment on the failure of the REP Na-7
test case. For the same reason, the present discussion rules out the applicability of the current and
proposed criteria to the UO2 fuels with additives,
Fuel enthalpy increase at failure[J/g]
1070
RANNS criteria for HT
RANNS criteria for RT
800
400
0
0
200
400
600
800
Hydrogen content [wtppm]
Figure 6. PCMI failure thresholds for SR-annealed cladding:
scaling of the NSRR test results by fracture mechanics
approach [26].
JOURNAL OF NUCLEAR SCIENCE AND TECHNOLOGY
whose specific behavior during an RIA, in particular fission gas behavior, has not been known.
Hence, it is desired and expected that more
details as bounding conditions and mechanisms
of the fission-gas-assisted additional PCMI loading are elucidated through the continued ALPS
tests and the Cabri International Program tests,
which involve fission gas dynamics tests to clarify
the different kinetics in fission gas behaviors
between different types of fuels [9]. All of these
experimental programs include both MOX and
doped fuels in their test matrices currently
planned.
4. Conclusion
This paper summarized the main results of the NSRR
tests performed after the establishment of the current
Japanese PCMI failure criteria in 1998, recalled the
reevaluation of fuel enthalpies of the earlier NSRR
tests with some modifications, and discussed the
applicability of the current criteria to the burnup
range from 65 to 75 GWd/tHM for each of the conventional and the advanced fuels in light of the
updated knowledge, expected revision of the criteria
using the current approach, and its possible
alternatives.
None of the newly added or updated RIA-test
results presented in this paper challenge the validity
(i.e., conservativeness) of the current Japanese PCMI
failure criteria, except for a few cases that can be
excluded from discussion, with fuel design and irradiation experiences in Japanese commercial nuclear
power plants taken into consideration. It can hence
be concluded that, given that the current approach
using fuel burnup as the primary parameter will be
retained, the current PCMI criteria is applicable to
fuels consisting of UO2 or MOX fuel pellets and
cladding tubes of Zry-2 for BWR and MDATM, MMDATM(SR), NDATM, ZIRLOTM, and M5TM for
PWR, including the burnup range from 65 to 75
GWd/tHM.
Failure-threshold curves based on cladding hydrogen content as a primary measure of fuel degradation
have been proposed as a possible alternative to the
current PCMI criteria, which can be used to judge the
occurrence of PCMI failure during an RIA. The
threshold curves were formulated based on the evaluation of fuel-enthalpy increases at the measured
failure times as far as possible. Ensuring similar safety
margins to the current criteria for each data point,
they could enable more performance-based safety
evaluation reflecting important factors lumped in or
neglected in the current criteria, such as improvements in the cladding materials with respect to their
corrosion resistance, irradiation conditions, and
1071
anticipated fuel temperature at the beginning of a
reactivity insertion event.
Acknowledgments
The RIA-simulated experiments in this study were performed under the research entrusted by Secretariat of
Nuclear Regulation Authority. The test fuels for the HBO,
TK, and OI experiments were provided by The Kansai
Electric Power Co., Inc. The information on fuel fabrication and irradiation was provided under cooperative
research agreements between JAEA and Mitsubishi Heavy
Industries, Ltd., and between JAEA and Nuclear Fuel
Industries, Ltd.
Disclosure statement
No potential conflict of interest was reported by the
authors.
ORCID
Yutaka Udagawa
http://orcid.org/0000-0001-9267-7606
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