EXAMINATION OF PWR SPENT FUEL RODS AFTER 15 YEARS IN DRY STORAGE

BACKGROUND

The original license and the certificate of compliance (CoO) for spent nuclear fuel dry storage

casks are due to expire for many commercial nuclear power plants in the next few years. The

Spent Fuel Project Office (SEPO) of the United States Nuclear Regulatory Commission (NRC)

is developing revisions to interim staff guidance (ISG) for review of technical documents to be

submitted by licensees in support of their applications for renewal of licenses and Co~s. The

renewal of licenses and CoCs for the existing casks would cover an additional period of 20 to

100 years and would require development of a technical basis for continued safe performance

under the extended service conditions. Consistent with the above and the regulatory

requirements of 10 CFR Part 72 for storage and handling of spent fuel, the acceptance criteria

for spent fuel in storage casks in ISG-1 1 are being revised. Revision 2 of ISG-1 1 (Ref. 1 )

recognizes creep to be the dominant mechanism for cladding deformation under normal

conditions of storage and recommends that the maximum cladding temperature during storage

and for short-term vacuum drying and cask backfilling operations be limited to 4000C to ensure

that the extended storage does not result in gross rupture of fuel cladding. The research

program reported herein provides a technical basis for the revision.

OVERVIEW OF THE PROGRAM

The research program, initiated by the NRC Office of Nuclear Regulatory Research (RES) in

November 2000 in response to the user need letter from the SEPO (NMSS-1998-005), involved

post-storage characterization and creep testing of representative pressurized water reactor

(PWR) fuel rods, stored for over 15 years in a single assembly (Ti11) in a CASTOR V/21 dry

storage cask. These rods came from the Surry nuclear power plant, had an initial enrichment

of 3.1 1%, and were irradiated for three fuel cycles to an average burnup of 35.7 GWdIMTU and

a fast (E>1 MeV) neutron fluence of 6.38 x 1025 nrnm2. The fuel cladding was made from coldworked/stress-relieved Zircaloy-4. The rods had cooled for -1300 days before they were stored

for -15 years, during which the peak cladding temperature decreased from 350CC to 1 50CC.

The post-storage characterization consisted of: (1 ) visual inspection of rods for crud, rod bow,

rod bulge, spallation, cracking, or other deleterious rod behavior, (2) profilometry of rods to

determine diametral changes as a result of creep deformation during storage, (3) fission gas

analysis and void volume determination of rods to estimate rod internal pressure and isotopic

composition of fission gas, (4) metallography of rod segments to determine the fuel-cladding

gap and the extent of cladding corrosion (oxide thickness), (5) hydrogen analysis of rod

segments to determine possible deleterious effects of hydrides on cladding integrity, and

(6) microhardness measurements to estimate cladding resistance to deformation. Table 1

below summarizes the post-storage characterization matrix.

Visual inspection and profilometry were performed on 12 rods from a single assembly that had

the highest assembly-average burnup and was stored in the hottest part of the cask. The

profilometry data exhibited very similar cladding diametral profiles for all 12 rods. The

differences found in the diametral profiles were insignificant relative to experimental uncertainty.

Four rods with marginally higher diametral changes were selected for fission gas analysis.

1
Measured gas pressures in all four rods were too close to allow further discrimination for the

remaining characterization work. Thus, two rods with the largest diameters and highest gas

pressures were selected for segmentation and further analysis that included metallography,

hydrogen analysis, and micr6hardness measurements. In selecting specific locations for rod

segmentation, the factors considered included (a) uniformity of burnup and fast fluence,

(b) proximity to the axial regions from which the creep samples were to be prepared,

(c) significant clad oxidation, and (d) highest temperatures experienced during thermal

performance testing and storage.

Table 1. Spent Fuel Post-Storage Characterization Matrix

Characterization Type Number of Rods Rod Segment Identification

Visual inspection 12 rods Full rod

Profilometry 12 rods Full rod

Fission gas analysis Rods H7, H9, G6, Full rod

G 9_ _ _ _ _ _ _ _ _ _ _ _ _ _

Metallography and Rod H9 3 rod segments (at fuel centerline, 500 mm

oxide thickness above, and 1000 mm above centerline)

Rod G6 2 rod segments (at fuel centerline, 500 mm

above centerline)

Hydrogen analysis Rod H9 3 rod segments (at fuel centerline, 500 mm

above, and 1000 mm above centerline)

Hydride orientation Rod H9 3 rod segments (at fuel centerline, 500 mm

above, and 1000 mm above centerline)

Rod G6 2 rod segments (at fuel centerline, 500 mm

above centerline)

Microhardness Rod H9 3 rod segments (at fuel centerline, 500 mm

above, and 1000 mm above centerline)

Aside from the extensive post-storage characterization work indicated above, the research

program focused on creep testing of fuel rods that had been in dry storage for -15 years. The

purpose of creep testing was to determine the residual creep capacity of stored fuel rods in

order to assess if the rods can be safely re-stored in a dry cask for an extended period of 20 to

100 years and safely transferred to a permanent repository. The creep data would also provide

valuable information for the analysts to develop creep correlations for incorporation into the

codes used for licensing review.

Five creep tests were conducted spanning a temperature range of 360 0C to 4000C and a stress

range of 190 MPa to 250 MPa. As mentioned previously, specimens for creep testing came

from locations in close proximity to rod segments used for post-storage characterization.

Table 2 summarizes the creep test matrix. The temperature range is prototypic of spent fuel

rod temperatures at the beginning of dry storage. The rod temperature decreases with time

2
during the storage so the temperatures selected for experiments are expected to provide a

conservative estimate of residual creep capacity. The stress range is not prototypic. Fuel rods

in dry storage are expected to encounter much lower stresses (-60 to 70 MPa at a burnup level

of 45 GWd/MTU). However, creep tests at this low stress level are not expected to yield any

measurable creep deformation within the test time frame (<<10,000 hours). Thus, accelerated

testing at higher stress levels was pursued, keeping in mind that the results from such tests

would be translated to stress levels appropriate for dry cask storage by analytical means.

Table 2. Spent Fuel Creep Test Matrix

Test # Temp. (0C) Stress (MPa) Test Time (hr.) Test Purpose

1 380 220 2180 residual creep capacity; primarysecondary creep data

2 380 190 2348 residual creep capacity; primarysecondary creep data

3 400 190 1873 residual creep capacity; primarysecondary creep data

6400 250 693 residual creep capacity; primarysecondary creep data

5 360 2290 3305 residual creep capacity; primary-

. ~~~secondary creep data

RESULTS

The results of post-storage characterization and creep testing are discussed in detail in

NUREG/CR-6831 (Ref. 2) and summarized below under each characterization category. Note

that the visual examinations of the 12 selected rods did not show any appreciable degradation

such as rod bow, spalled cladding, or cladding creep, even after 15 years of dry storage. There

was evidence of crud and oxide buildup from reactor operation but no crud or oxide appeared to

have fallen off. The selected rods are those that came from a 15x15 fuel assembly stored close

to the hottest part of a Castor V/21 dry storage cask (Ref. 3). The assembly was irradiated in

the Surry PWR for three fuel cycles and had the highest assembly average burnup.

Profilometrv

Linear profilometry traces of each of the 12 rods were obtained at 0, 45, 90, and 1350 relative

orientations and at -2.5 mm intervals along the length, starting at -330 mm from the top of the

rod. Averaging the four linear traces at a given axial location generated the cladding outer

diameter profile at that location. An example profile for one of the 12 rods, marked H9, is

shown in Figure 1. Noting that the nominal as-fabricated rod diameter is 10.71 mm, it is seen

that the maximum decrease in cladding diameter of rod H9 is -0.06 mm, which is equivalent to

a creepdown of -0.6% that is typical of PWR rods in the burnup range of 36 GWd/MTU.

3
Prof ilometry traces of other rods were very similar to that of rod H9 and showed no discernible

differences.

Fission Gas Analysis

The void volume, gas composition and inventory, and fission gas isotopic composition were

measured on four of the 12 rods to determine internal rod pressure and to estimate the extent

of maximum possible fission gas release during the storage period. The void volumes ranged

from 19.5 x 104 to 20.4 x 10 in3, which is typical for 15x15 PWR rods. The gas composition

measured in all four rods were essentially the same, i.e., 96-98% helium fill gas with small

amounts of oxygen (<0.01%) and nitrogen (<0.03%). The fission gas release values ranged

from 0.4 to 1.1%, well within the range reported in the literature for the rod type and burnup

range, indicating that no significant amount of fission gas was released from the fuel during the

dry storage period. The Xenon fraction in the fission gas was between 1.5 to 3.2 mol%, and the

Krypton fraction was between 0.1 to 0.4 mol%. The Xenon to Krypton ration ranged from 9.0 to

11.4 for the four rods, again typical of values reported for PWR fuel rods.

E ID75

E

C0, E 070

Cu*

O ~~~~~~~~~~~A Grid Spacers

0 B0O ID00 too 2000 2500 2000 2500 4000

Location (nm, from rod bottom)

Figure 1. Average Outer Diameter Profile of Rod H9 (from Surry-2 Assembly Ti11)

After 15 Years of Dry Cask Storage

Metalloaraphy and Oxide Thickness

The thickness of the oxide layer on the cladding outer surface was measured from micrographs

(metallographs) taken at eight azimuthal locations, spaced evenly around the cladding and

covering -45% of the cladding circumference. For rod H9, the oxide layer thickness was

measured at three axial locations: fuel centerline, 500 mm above centerline, and 1000 mm

above centerline. For another rod G6, the thickness was measured at two axial locations: fuel

centerline and 500 mm above centerline. The oxide thickness measurements are shown in

Table 3.

4
Table 3.Cladding Oxide Thickness (pm) at 0, 500, and 1000 mm Above the Fuel Centerline

'Turkey Point - D assemblies data provided for comparison. The fuel in D assemblies had lower bumup than

the Surry fuel and hence, smaller oxide thickness.

It is seen from the table that at higher elevations, the oxide layer is thicker because of the higher

in-reactor temperatures at these elevations. Figure 2 shows the micrographs of oxide layer at

the fuel centerline and 1000 mm above the centerline at one azimuthal location in rod H9. The

oxide layer at the centerline is thin and adherent whereas the oxide layer at 1000 mm above the

centerline is thicker and has isolated cracks and/or microvoids. Again, this is consistent with

oxide growth morphology whereby thicker oxide layers are more prone to cracking. Figure 3

shows the cross-sectional views of rod H9 at the fuel centerline and 1000 mm above the

centerline, depicting a fuel pellet with multiple cracks and voids that are prototypic of this type of

medium burnup PWR fuel.

oxide -l VI oxide Al lz

F

(A) (B)

Figure 2. Cladding Oxide Layer of Surry-2 Rod H9 from Assembly T11 at:

(A) Fuel Centerline, and (B) 1000 mm Above Centerline

5
(A) (B)

Figure 3. Cross-Sectional Views of Surry-2 Rod H9 from Assembly T1 I at:

(A) Fuel Centerline, and (B) 1000 mm Above Centerline

Hydrogen Analysis and Hydride Orientation

The hydrogen content of samples from eight azimuthal locations at each axial elevation (fuel

centerline, 500 mm above, and 1000 mm above) was measured using a Leco analyzer. The

hydrogen content was also calculated from the oxide thickness measurements. Both measured

and calculated values are shown in Table 4 for rods H9 and G6. Note that the calculated values

are based on a pickup factor (fraction of available hydrogen migrating into cladding) in the range

between 16 and 24%, typical of medium burnup PWR fuel rods, and assuming no axial migration

within the cladding.

Table 4. Hydrogen Content (wppm) at 0, 500, and 1000 mm Above the Fuel Centerline

Rod ID 0 mm (fuel centerline) 500 mm 1000 mm

Measured Calculated Measured Calculated Measured Calculated

H9 250 ± 40 200 ± 70 300 ± 25 270 ± 100 255± 14 330 ± 100

G6 not 180 ± 70 not 210 ± 60 not not

measured measured _ _ measured calculated

TP-D01/046 40-60 40-85 50-70 40- 100 75-90 65- 150

aTurkey Point - D assemblies data provided for comparison. The fuel in D assemblies had lower burnup

and the cladding had smaller oxide growth than the Surry fuel and, hence, lower hydrogen content.

A hydrogen content of 250 to 300 wppm in cladding is typical of medium burnup fuel. Almost all

of the hydrogen (200 to 250 wppm) stays in solid solution in the zircaloy matrix at temperatures

4001C (pre-storage temperature) and above. As the cladding temperature decreases during the

storage period, the hydrogen in solid solution eventually precipitates as hydrides. At about

150'C corresponding to the end-of-storage temperature of fuel rods after 15 years in dry storage,

essentially all of the hydrogen would have precipitated as hydrides. Figure 4 shows hydride

6
structure in rods H9 and G6, at 500 mm above the fuel centerline and at the fuel centerline,

respectively.

(A) (B)

Figure 4. Hydride Structure in Surry-2 Fuel Rod Cladding:

(A) 500 mm Above Fuel Centerline of Rod H9; (B) Fuel Centerline of Rod G6

As seen in the above figure, hydrides are oriented circumferentially and though very few smaller

hydrides in some samples may be oriented at angles between 0 and 900, no radial hydrides were

observed at the cladding/oxide interface. Hydrides, which result from clad oxidation during inreactor operation, contribute to a loss of clad ductility if oriented radially in large quantities in

cladding.

Microhardness Measurements

Cladding microhardness was measured at three axial locations of rod H9 and at four azimuthal

points (-90° apart) per each axial location. The hardness numbers at three locations were

calculated to be 234 + 18, 239 + 5, and 236 ± 10, respectively. The microhardness is a measure

of cladding resistance to permanent or plastic deformation. During reactor operation, the

cladding undergoes radiation hardening. When out of the reactor, the same cladding

experiences thermal excursions at 4000C and above during pre-storage operation. Most of the

irradiation damage (hardening) is annealed out if the cladding is exposed to such temperatures

for a reasonable length of time. At temperatures -350 0C and below, hardly any annealing takes

place. Little, if any, annealing occurred during the pre-storage operation or during storage of the

Surry cladding under consideration. This is supported by the uniformity of the calculated

hardness values as above at three different locations and also by a comparison of the Surry

hardness data with the Turkey Point hardness data.

Post-Storage Thermal CreeR Tests

As mentioned previously, five creep tests were conducted under conditions summarized in

Table 2 above. In these tests, changes in cladding outer diameter were measured periodically

at several locations (9 axial locations and 5 azimuthal points at each axial location) in a 76-mm

long creep test specimen as a function of time. The hoop strain was determined from these

measurements as a function of time (see, for example, Figure 5 for Test #1).

7
2.0

1.8 Nominal Test Conditions 0.232% @ 335 h

Temperature: 380'C 0.370% @ 671 h

1.6 - Engineering Hoop Stress: 220 MPa 0.514% @1028 h 0.818% @1820 h

1.4 -0.972% @ 2180 h

_ 1.2

Average

0.

o 0.8

0.6 0

0.41

0.24

0.0

0 1,000 2,000 3,000

Time (h)

Figure 5. Thermal Creep Data (Test #1) for Irradiated, Dry-Cask-Stored Surry Cladding

Diamond: hoop strain from measurements; Triangle: wall-averaged strain

Figure 6 shows the hoop strain data for tests #2 and 3 combined. Note that Test #3 was a

continuation of Test #2 beyond 1873 hours when the stress level was increased from 190 MPa

to 250 MPa. The hoop strain increased significantly at the elevated stress level and at the

conclusion of Test #3, the total creep strain was -6%. The creep sample appeared to be still

within the secondary (steady-state) creep regime and no creep failure was observed. Table 5

summarizes the creep data from all 5 tests.

6.0 -

Nominal Test Conditions

Temperature: 400'C I0 Strain

5.0 Engineering Hoop Stresses: 1901250 MPa 0.079% @ 48 h

Average 0.136% @ 120 h

0.209% @ 240 h

- 4.0 0.266% @ 360 h

0.351% @ 552 h

c l0.494% @ 912 h

0.656% @1248 h

( 3.0 0.729% @1489 h

0 Stress increased 0.920% @1873 h

0 from 190 to 250 MPa POD1.588% @2041 h

2.0 : f2.056% @2162 h 2.0 2.918% @2330 h

5.161% @2566 h

1.0

0.0

0 1,000 2,000 3,000

Time (h)

Figure 6. Thermal Creep Data (Tests #2 & 3) for Irradiated, Dry-Cask-Stored Surry Cladding

Diamond: hoop strain from measurements; Triangle: wall-averaged strain

8

CON
Table 5. Hoop Strain and Secondary Creep Rate Data from Five Creep Tests

Test # Temp. Stress Test Time Hoop Strain Secondary Creep

(OC) (MPa) (hr.) (%) Rate (%/hr.)

1 380 220 2180 1.10 4.5 x 104

2 380 190 2348 0.35 8.8 x 10' 5

3 400 190 1873 1.04 4.9 x 104

4 400 250 693 5.83 4.9 x 10'3

5 360 220 3305 0.22 4.2 x 1 O5

Inasmuch as no bulging or asymmetric wall thinning, a precursor to creep failure, was observed

in any of the creep specimens, the results indicate that significant residual creep capacity

remains in fuel cladding after the initial storage period of 15 years. With decreasing

temperature and stress during the extended storage beyond the initial 15 years, the cladding is

not likely to experience a significant amount of additional creep. Noting that two of the five tests

were conducted at 400CC, the results support the recommendation of 4000C as the maximum

allowable beginning-of-storage cladding temperature for dry cask storage of spent fuel.

REGULATORY IMPLICATIONS

The research results reported herein suggest that the PWR spent fuel cladding after 15 years in

a dry storage environment retained sufficient structural integrity to remain as a barrier to fission

product release. There was no evidence of hydrogen pickup or hydride reorientation during the

storage period. Little, if any, cladding annealing occurred during the pre-storage operation and

15-year dry storage period. Post-storage creep data indicate significant creep capacity

remaining in the cladding, thus suggesting that very little thermal creep occurred during the

storage period. These observations suggest the viability of spent fuel cladding to remain as a

barrier to fission product release during extended storage up to 100 years in a dry cask

environment. Further, these suggest the likelihood of spent fuel cladding to retain sufficient

structural integrity after the extended storage period for safe transfer to a final repository. The

results provide a sound technical basis in support of Revision 2 of ISG-1 1, which recommends

that the beginning-of-storage cladding temperature be limited to 4000C. The results also

provide a sound technical basis for granting an extension of cask licenses and CoCs, thereby

reducing regulatory burden on licensees while maintaining the safety of spent fuel storage.

Note that the results are valid for spent fuel with a burnup limit of 45 GWd/MTU. RES is

currently conducting a similar experimental program for high burnup spent fuel cladding (burnup

in excess of 45 GWd/MTU). The results from the high burnup program will be reported

separately in the future.

9
REFERENCES

1. "Interim Staff Guidance 11 - Revision 2," Spent Fuel Project Office, Office of Nuclear

Materials Safety and Safeguard, July 30, 2002.

2. R.E. Einziger, H. Tsai, M.C. Billone, and B.A. Hilton, 'Examination of Spent PWR Fuel

After 15 Years in Dry Storage," NUREG/CR-6831, USNRC (Prepared by Argonne

National Laboratory), September 2003.

3. W.C. Bare and L.D. Torgerson, "Dry Cask Storage Characterization Project - Phase 1:

CASTOR V/21 Cask Opening and Examination," NUREG/CR-6745, USNRC (Prepared

by Idaho National Engineering and Environmental Laboratory), September 2001.

10