A comparative analysis of UO2 and MOX fuel behavior under reactivity initiated accident

Pergamon

Ann. Nucl. Energy, Vol. 24, No. 11, pp. 859 870, 1997 © 1997 Elsevier Science Ltd. All rights reserved PIhS0306-4549(96)00063-1 Printed in Great Britain 0306-4549/97 $17.00 + 0.00

A C O M P A R A T I V E A N A L Y S I S O F UO2 A N D M O X F U E L BEHAVIOR UNDER REACTIVITY INITIATED ACCIDENT YANG-HYUN KOO it, DONG-SEONG SOHN ~ and BORIS VOLKOV 2 1Korea Atomic Energy Research Institute, 150 Duckjin-Dong, Yuseong, Daejeon 305-353, South Korea 2IRTM RRC "Kurchatov Institute", Kurchatov Square, Moscow 123182, Russia (Received 18 July 1996) Abstract--Reactivity initiated accident (RIA) is one of the reactor accidents that limit the in-reactor performance of fuel especially at high burnup. In addition, according to recent experiments, RIA at high burnup can lead to fuel failure. In this paper, a comparative analysis was carried out for UO2 and Mixed Oxide (MOX) fuel behavior under RIA as a function of burnup (0, 30, 60 MWd/kgM) and RIA type (realistic and severe) with use of the FRASM code. Comparison of the present analysis with experimental results gives the following three conclusions. First, under realistic RIA which can take place under reactor operating conditions, both UO2 and MOX fuels are intact. Second, under severe RIA which leads to local melting of fuel, there is a possibility that both UO2 and MOX fuel would fail. Third, UO2 fuel behavior under RIA is more favorable than MOX fuel at beginning of life, while MOX fuel is more favorable at the burnups of 30 and 60 MWd/kgM. Fuel failure criterion applied to RIA, however, should be elaborated based on experimental data. © 1997 Elsevier Science Ltd.

INTRODUCTION The in-reactor behavior of UO2 fuel at high burnup under reactivity initiated accident (RIA) in PWR has been a great concern as several recent experimental results show that current fuel failure criteria based on the test results of unirradiated fuel rods are inadequate. For example, CABRI test results (Schmiz et al., 1994) raised the concern that the failure threshold for high burnup fuel would be significantly reduced due to the degradation of fuel cladding properties with burnup. In addition, MOX fuels are beginning to be tAuthor for correspondence. 859

860

Yang-Hyun Koo et al.

used up to high burnup recently in PWRs of European countries. Therefore, integrity of both UO2 and MOX fuels at high burnup under RIA is very important in terms of reactor economy and safety. In this paper the in-reactor behavior of both UO2 and MOX fuel under RIA is analysed with use of the FRASM code (Volkov and Valach, 1994) as a function of burnup and RIA types. Differences in material properties such as thermal conductivity, specific heat capacity, thermal expansion, Young's modulus and melting temperature for two types of fuel are considered in the present analysis (Sohn et al., 1995a). Difference in radial power distributions for the two types of fuel with burnup is also taken into account (Sohn et al. 1995b). Other models for fuel and cladding from the well-known MATPRO 10 library (Bohn et al., 1978) are used for this analysis. Hereinafter, UO2 fuel is marked as 'Fuel A' and MOX fuel as 'Fuel B.'

SIMPLIFIED REACTOR KINETIC MODEL FOR RIA

Thermo-mechanical behavior of a fuel rod under RIA basically depends on the deposition energy that is determined by both peak power and duration of RIA. Therefore, to analyse the fuel rod behavior under RIA by means of a computer code, it is necessary to input the time-dependent power of fuel rod during this accident. As is known, any reactivity transient is characterized by its speed and extent of the power during which it can reach. Based on the method given by Valach and Volkov (1996), we can derive the time-dependent power of a fuel rod in terms of q0 and qmax where q0 and qmax are the average and maximum linear power, respectively. Then the time-dependent linear power of fuel under RIA is transformed into a relative form as follows:

q( t) = qoK( t)

(1)

e-RT K(t) = 16K2ax (1 + 4Kmaxe-R') 2

(2)

where

and

K(t) = q(t)/qo

and

Kmax= qmax/qo.

Similarly, the expression for half-width of reactor power pulse is derived as A~1/2 ~

3.52 R'

(3)

where R is defined below under the assumption that qmax > > q0

(4) In equation (4), P0 and/~ are the magnitude of the step insertion and the delayed neutron

A comparative analysis of UO2 and MOX fuel behavior

861

fraction, respectively. Consequently, if the half-width of reactor power is known, the timedependent power of a fuel rod can be calculated using equations (1), (2) and (3). Then the time-dependent power andenergy release for the step insertion in a fuel rod is calculated as follows provided p0 > fl: E(t) = qo

(5)

K(t)dt.

CALCULATIONAL DATA In order to analyse the integral behavior of Fuel A and Fuel B, following two RIA types were chosen (see Table 1): - - realistic RIA due to control rod ejection: RIA in a real PWR due to control rod ejection would be self-limited by delayed neutron and negative feedback of fuel temperature. Based on neutron physics calculation for a typical 900 MWe PWR, the typical values of Table 1 were determined (Lee et al., 1995) --severe RIA to determine the limiting fuel behavior: in this scenario, Kmax is chosen in such a way that the melting temperature is reached in the fuel center at maximum power. Moreover, for these two RIA types the following three burnups were considered for the present analysis: --Beginning of life (Burnup: 0 MWd/kgM), --Medium of life (Burnup: 30 MWd/kgM), --End of life (Burnup: 60 MWd/kgM). As a result, a total of 12 cases were analysed by the FRASM code. The initial data for two types of fuel at three burnups were prepared by using a KAERI's (Korea Atomic Energy Research Institute) steady-state code (Sohn et al., 1995b). It is necessary to note here that the purpose of this work is not to examine the separate effects of fuel rod behavior but to analyse the integral behavior under RIA. Fuel rod initial characterization

The as-fabricated fuel rod characteristics are presented in Table 2. The identical initial parameters are chosen for the present RIA analysis. Coolant conditions

The coolant conditions chosen for the present analysis are those for typical 17 × 17 PWRs: ---coolant pressure : 153.0 bar ---coolant mass flux: 2470.2 kg/s. m 2 --inlet coolant enthalpy :1.218 × 106 J/kg Table 1. Parameters for two types of RIA Kmax Realistic R I A

Severe RIA

/90 - ]~

A (S)

R (s-l)

Art/2(s)

50

0.011

1000

0.035

10-4

110

0.032

10-4

350

0.010

Yang-Hyun Koo et al.

862

--outlet coolant enthalpy: 2.610 x 106 J/kg. These conditions were assumed to be constant during RIA. The axial distribution of coolant parameters, however, was considered for the calculation of external heat transfer coefficients during power jump.

Initial data for different burnups The changes in fuel rod conditions with burnup were characterized for the following parameters: --axial and radial distribution of heat generation --axial and radial dimension of fuel and cladding --axial and radial distribution of fuel density --gas pressure in the gap ---cladding oxide layer thickness --increase in cladding strength and decrease in cladding ductility due to neutron fluence. The material properties of cladding in the present analysis are only influenced by irradiation time and do not depend on types of fuel since the typical Zircaloy-4 claddings are assumed to be used for both fuels. The initial fuel parameters just prior to RIA, however, are affected by burnup and fuel type. The axial distributions of linear power at different burnups are assumed to be the same for both fuels. The fuel densities in pellets and the relative radial power distributions are given as a function of burnup and fuel type in Figs 1 and 2. In these two figures, comparison is made between the KAERI code's and FRASM's calculation for these parameters. The main initial data for both fuels at different burnups are presented in Table 3 for the axial segments 6 and 9 when a fuel rod is divided into l0 axial segments with equal length. The reason for the selection of axial segments 6 and 9 as analysis segments is that we can get the highest fuel temperature in axial segment 6 at beginning of life and in axial segment 9 at the burnup of 60 MWd/kgM. Table 2. Fuel rod initial characterization Parameters Enrichment (A) or Pu-fissile content (B) O/M ratio Fuel roughness Clad roughness Fuel density Cladding outside diameter Cladding inside diameter Wall thickness Fuel outside diameter Central hole diameter Fuel stack length Fuel-cladding diametrical gap Upper plenum volume Gas pressure of He in the gap

Fuel A(UO2)

Fuel B (MOX)

3.5%

3.5%

2.00

2.01 2.0 lam 1.0 ~tm 93.5% 9.45 mm 8.18 mm 0.635 mm 8.06 mm 0.0 mm 3658.0 mm 0.12 mm 7.0 cm3 21.5 bar

A comparative analysis of UO2 and MOX fuel behavior

863

C O M P A R A T I V E A N A L Y S I S OF FUEL A A N D FUEL B U N D E R RIA

Using the data o f Table 3, 12 calculations were made and some of the results are presented in Figs 3-5 for realistic RIA at beginning of life. In addition, Figs 6-8 represent the calculation results for severe RIA at the burnup of 60 MWd/kgM. On the basis of these calculations, the impact of changes in radial power distribution and material properties on the fuel behavior was analysed. Since the material properties of both fuels are not so different from each other except slightly lower thermal conductivity for Fuel B, the radial power distribution is a main factor that induces different fuel behavior. As can be seen in Figs 1 and 2, there is a marked difference in radial power distributions in both fuels especially at the beginning of 1.0

0.9

o,8

0,7

0.6

0.5

o,o

o12

o14

o:8

1.o

Fuel relative radius

Fig. 1. Radial power distribution of Fuel A for different burnups. 1.0

Fuel B

/~

-4- 30 M'~M/kgM : KAERI Code

/ ,~

-*-

0.9

~ ~ M

: KAem

Code

•... 30 I,t'~kM~M : FII~6M

/~ /4 I

/r

,,~

0.8

~

0.7 [3" - -El-

-

0.6

0.5

00

012

014

016

018

I0

Fuel relative radius

Fig. 2. Radial power distribution of Fuel B for different burnups.

864

Yang-Hyun Koo et al.

fuel life. For this condition, enthalpy increase in Fuel A is much lower than that in Fuel B (see Fig. 3). This can be explained by the gap conductance's dependence on the gap size and contact pressure. As is well known, stored energy is approximately proportional to the power density and is also affected by both thermal resistance of pellet-cladding gap and fuel material properties. The initial data in Table 3 gave rise to the gap contact in Fuel A only as shown in Fig. 5. The reason that the gap contact occurred only in Fuel A can be explained, first of all, by Fuel A's radial power distribution leading to higher average temperature (see Fig. 4) at the beginning of life. Accordingly, the gap conductance in Fuel A is much higher than that in Fuel B. Consequently this leads to more heat Table 3. Main initial data for Fuel A and Fuel B Burnup (mean) (MWd/kgM)

0

Fuel type Gas pressure (bar) Released gas fraction (%) Cold gap width at axial segment 6 (mm) Cold gap width at axial segment 9 (mm) Maximum neutron fluence (E + 20/crn2) Oxide thickness (Ixm) Cold plenum volume (era3) Maximum linear power, (W/cm)

30

A

B

67.4 0

71.1 0

60

A

B

A

B

74.7 0.20

78.0 0.32

81.9 0.80

86.3 1.04

0.120

0.120

0.0277

0.0400

0.0217

0.0336

0.120

0.120

0.0239

0.0357

0.0195

0.0310

0 0 7.00

0 0 7.00

321.4

321.4

57. I 29.4 6.15 258.5

57.1 29.4 5.92 258.5

119.4 142.4 4.48

119.4 142.4 4.20

216.7

216.7

30.0

2.0E+6

25.0

Axial segmem 6 ( B ~ m p : 0 MWO/kgM)

1.5E+6

.mmm

Linear power

i

m =- Enet'~" Release ,,= Fuel A enll~py increar,e ~ ~ Fuel B ¢nthall~J increase

\\ \

1.0E+6

J

\

20.0

15.0

._=

10.0

.1 !

5.0E-H~ 5.0

0.0 31.90

32.00

32.10

32.20

32.30

0.0 32.40

Tilll¢,s

Fig. 3. Comparison of enthalpy increase during realistic RIA at 0 MWd/kgM.

A comparative analysis of UO2 and MOX fuel behavior

865

transfer from pellet to cladding in Fuel A, which in turn reduces enthalpy increase in Fuel A compared with Fuel B. At the burnup of 60 MWd/kgM, axial segment 9 was chosen for the present analysis since it has maximum linear power at this burnup. In this case, the radial power distribution and therefore temperature in Fuel A are not so different from those in Fuel B as shown in Fig. 7. However, Fig. 8 indicates that the lower thermal conductivity of Fuel B 2200 /

1600

2100

J

1500 1400

2000 /

T.

1300

/

1900

/ Axial segment 6 (Burnup : 0 MWd/kgM) ~ Fuel A - - - Fuel B

/

1200

~J

1100

1800 1700 1600

/N\

1000

~

c~

~" E u "6

1500

900

o

800

1400

700

1300

60O 31.90

i 32.00

z 32.10

3;.20

3;.30

r.j

1200 32.40

Time, s

Fig. 4. Comparison of centerline fuel temperature (T~), outside temperature (Tfo), inside cladding temperature (Td), outside cladding temperature (Tco) at the beginning of life during severe RIA. 9.60E-3

2.0E+7 A

x

i

a

l

Pcont

segment 6 (Burnup : 0 MWd/kgM) Fuel A

1.5E+7

~=

9.56E-3

1.0E+7

,

___J

31.00

i Dco

5.0E-~

0.G

E

9.48E-3

z

32.00

....J 33.00

I 34.00

9.44E-3 35.00

Time, s Fig. 5. Comparison of contact pressures, gas pressures and diameter changes at the beginning of life during RIA at 0 MWd/kgM (contact pressure for Fuel B is 0 at this burnup).

866

Yang-Hyun Koo et

al.

induces the higher fuel temperature and, as a consequence, the more thermal expansion leads to higher contact and gas pressures than in Fuel A. By the same explanation given above, enthalpy increase in Fuel B is lower than that in Fuel A (see Fig. 6) although the difference is not significant. The result of comparative analysis of fuel behavior under RIA is that while the radial power distribution plays a key role at lower burnup in distinguishing both fuels' response, at higher burnup the difference in material properties is the main factor causing the difference in both fuels' behavior. 4.0E+7

250

, . . . . . .

200

J

3.0E+7 B

150

8 2.0E+7 100 I

1.0E+7

I (B=~: 1 ~ I ....

t L

0.0--31.00

33.00

32.00

6o M W d P , ~ ~ w e r ~ - q y re.s~

so

34.00

35.00

Time, s

Fig. 6. Time-power distribution, energy release and comparison of enthalpy increase during severe RIA at 60 MWd/kgM. 3200

2800

Axial segment 9 (Bumup : 60 MWcl/kgM) ~Fuel A - - - Fuel B

2400

f /f

2000

1600

/t

1200

.0 L _ _ j ~ J j

32.0°

~'.02

3,'.o,

32'.~ ' ~.0.'

32.10

"time, s

Fig. 7. Comparison of centerline fuel temperature (T~) outside temperature (Tfo), inside cladding temperature (Td), outside cladding temperature (T~o) during severe RIA at 60 MWd/kgM.

A comparative analysis of UO2 and MOX fuel behavior

867

RESULTS AND DISCUSSION IN TERMS OF BURNUP At the beginning of life the contact between pellet and cladding can play the positive role in decreasing fuel enthalpy increase via improved gap conductance. At high burnup, the contact pressure increases significantly due to narrower gap just prior to RIA. As a consequence, the pellet is strongly bonded to the cladding resulting in the strong mechanical interaction of pellet and cladding. This means that the pre-transient's conditions of the fuel rod exert significant influence on its behavior during RIA. In addition, two other features related to burnup were drawn from the present analysis. First, the embrittlement and strengthening of cladding due to fast neutron fluence induce the higher contact pressure. Therefore at high burnup cladding failure can occur even at small deformation. Second, the fast increase in fuel temperature during RIA leads to free volume decrease due to large thermal expansion of pellets. As a result, the internal gas pressure increases significantly and induces additional stresses in the cladding. This gas pressure can exert greater influence on the behavior of cladding at the beginning of life than the contact pressure which is expected to be very small due to wide gap at this time. COMPARISON OF CALCULATION RESULTS WITH FAILURE THRESHOLDS In this section an attempt is made to compare some of experimental data (Fujishiro and Ishijima, 1994, Yanagisawa et al. 1994) with present calculations. Figure 9 shows the comparison of calculated peak fuel enthalpies for different burnups with experimental data obtained from several test reactors (Fujishiro and Ishijima, 1994). This illustration may be considered as a failure criterion of fuel rod undergoing RIA. As this figure displays, the calculated values of peak fuel enthalpy for fuel rods with burnup less than about 30 MWd/kgM are placed in the region of no failure. For higher burnups, however, 1.20E-2

1.2E+8

1.15E-2

E

8.0E+8

,I

o

i IB,,~,,p:60 Mwd1~Ml I I -~°~"

1.10E-2

I 1.05E-2

P~s

4.0E+8

2Z/ 0.0 32,00

1.00E-2

I 32.02

I 32.04

; 3 .06

i 32,08

9.50E-3 32,10

Time, $

Fig. 8. Comparison of contact pressures, gas pressures and diameter changes for Fuel A and Fuel B during severe RIA at 60 MWd/kgM.

868

Yang-Hyun Koo et al.

the deficit of experimental data does not allow to define the proper failure threshold for severe RIA. As for realistic RIA, the experimental data except one obtained at CABRI suggest that the failure of fuel rods is unlikely to occur for all the range of burnups up to 60 MWd/kgM. Since the fuel behavior under RIA conditions is strongly affected by reactor kinetics, we should pay attention to reactor kinetics [that is, Kmax and R in equation (2)] in evaluating the CABRI data obtained at the burnup of about 65 MWd/kgM. In Fig. 10, a 400 * - - Real RIA UO~ Real RIA MOX Severe RIA UO= Severe RIA MOX • PBF failure • NSRRfailure + CABRIfailure • CDC failure

30(

] MOX

/ 100 r

•

/

+ i

0

10

20

30

40

70

80

Burnup (MWd/kgM)

Fig. 9. Comparison of calculated peak fuel enthalpy with experimental data for different burnups. 800

~11• [] 600

S e ~ r e RIA.

G 8 401

[]

E Realistic RIA

]

20~

Fuel B l I_. Fu'A 200

400

600

800

1000

1200

Claddinghoop su~ss (MPa) Fig. I0. Comparison of calculated cladding hoop stresses versus temperatures with PCMI failure threshold.

A comparative analysis of UO2 and MOX fuel behavior

869

comparison is made between PCMI failure threshold determined by ring tensile test (Yanagisawa et al. 1994) at out-of-pile conditions with calculated cladding stress and temperature. According to this comparison a fuel rod does not lead to failure up to 60 MWd/kgM under realistic RIA. For severe RIA, fuel rods with 30 and 60 MWd/kgM exceeded greatly the PCMI failure threshold, although fresh fuel does not appear to exceed it. However, note that the conditions of out-of-pile ring tensile test for the determination of failure threshold was not given in this reference (Yanagisawa et al., 1994). The test conditions such as stress and strain rate, content of hydride, and irradiation time are very important for the determination of PCMI failure threshold. For example, the stress and strain rate obtained in the present analysis reached values larger than 60,000 MPa/s and 500%/s, respectively. This very fast loading would not have been achieved in establishing the failure threshold given in Fig. 10. Therefore it should be noted that comparison in Fig. 10 does not provide adequate prediction for fuel behavior under RIA.

CONCLUSION The main results from the present analysis for fuel behavior under RIA can be summarized as follows: • The comparative analysis shows that MOX fuel behavior during RIA conditions has worse characteristics than UO2 at beginning of life due to radial power distribution leading to lower gap conductance (lower contact pressure) compared with UO2 fuel. On the other hand, at 60 MWd/kgM where the radial power distributions for both types of fuels are not so different, fuel behavior of both fuel under RIA are very similar. Therefore, the radial power distribution is one of the most important factors determining fuel behavior during RIA. • The degradation of fuel material properties with burnup leads to relative increasing (relatively of energy deposition) of fuel enthalpy peak during RIA. • The embrittlement and strengthening of cladding due to irradiation exert the main influence on increasing of fuel-cladding contact pressure and therefore on failure threshold.

REFERENCES

Bohn, M. P. et al. (1978) MATPRO-VERSION 10. A handbook of material properties for use in analysis of light water reactor fuel rod behavior. TREE-NUREG-1180. Fujishiro, T. and Ishijima, K. (1994) NSRR experiments to study the effects of burnup on the fuel behavior under reactivity initiated accident conditions. High-Burnup Fuel Behavior Session of 22nd Water Reactor Safety Information Meeting 24-26 October. Lee, C.-B. et al. (1995) Analysis of the fuel behavior under rod ejection accident in the pressurized water reactor. OECD Specialist Meeting on Transient Behavior of High Burnup Fuel Cadarache, France, 12-14 September. Schmiz, F. et al. (1994) Investigation of the behavior of high burnup PWR fuel under RIA conditions in the CABRI test reactor. High-Burnup Fuel Behavior Session of 22nd Water Reactor Safety Information Meeting, 24-26 October. Sohn, D.-S. et al. (1995a) Compilation of material properties of UO2 and MOX fuel. KAERI/TR-556/95.

870

Yang-Hyun Koo et al.

Sohn, D.-S., Koo, Y.-H. and Lee, B.-H. (1995b) A computer code for the thermomechanical analysis of fuel rod during steady-state operating conditions. KAERI Internal Report, September. Valach, M. and Volkov, B. (1996) Fuel rod modelling under LOCA and RIA conditions. Enlarged Halden Programme Group Meeting on High Burnup Fuel Performance, Safety and Reliability and Degradation of In-Core Materials and Water Chemistry and Man-Machine Systems Research, Leon, Norway, 19-24 May. Volkov, B. and Valach, M. (1994) FRASM (version 94): computer code for analysis the thermomechanical behavior of LWR fuel during LOCA and RIA. Vol 1: Code manual. UJV, Z-18-T,M, November. Yanagisawa, K., Katanishi, S. and Fujishiro, T. (1994) Transient FGR and PCMI of PWR fuels preirradiated to 42 MWd/kgU. Proc. of 1994 Int. Topical Meeting on Light Water Reactor Fuel Performance, ANS, April, p. 248.