ARGONNE NATIONAL LABORATORY, ARGONNE, ILLINOIS

THE PERFORMANCE OF EBR-II MARK-II METALLIC DRIVER FUEL UP TO 675 C

by R. E. Einziger

-NOTICE -

This report WM prepared as an account of work sponsored by the United Staus Covenimeni. Neither the United Stitei nor the United States Department of Energy, nor any of their employees, nor any of their contractors, subcontractor);, or their employees, makes any warranty, express ot implied, or assumes any legal liability or responsibility for the accuracy, completeness ot usefulness of any information, apparatus, product or process disclosed, 01 represents that its use would not infringe privately owned rights.

Prepared for

American Nuclear Society International Conference on Fast Breeder Reactor Fuel Performance

Monterey, California March, 1979

UolCtUtUSOOE

ARGONNE NATIONAL LABORATORY, ARGONNE, ILLINOIS Operated under Contract W-31-109-Eng-38 for the

U. S. DEPARTMENT OF ENERGY

THE PERFORMANCE OF EBR-II MARK-II METALLIC DRIVER FUEL UP TO 675°C

R. E. Einziger

INTRODUCTION

Metallic fuels, used in the early LMFBR's, were abandoned in favor of oxides mainly because the high temperature performance of the metals was questioned. In particular the fuel-cladding eutectic temperature was thought to limit the performance of the fuel element. As a result of the continued decrease in target reactor outlet temperatures, metal fuels can once again be considered viable candidates for LMFBR's. At peak cladding temperatures of 590°C or less, the highly successful irradiation program for the EBR-II Mark-II driver fuel has shown that high fuel burnups are possible for metal fuels.1'2 Until the present experiment, no data were available on performance of metal fuels near the fuel-cladding eutectic temperature.

In this experiment, three subassemblies were designed to operate at 20Z overpower as part of the EBR-II driver-fuel qualification program.

One subassembly operated at a peak inside cladding temperature (PCT) of 675CC, just 40°C below the fuel-cladding eutectic temperature. Although the performance at 675°C was completely satisfactory, it was significantly different than expected. This paper will present the performance for the high-temperature irradiation and compare it with performance at the lower temperatures. In addition, the paper will analyze FCCI, FCMI, and fuel microstructure at 675°C.

EXPERIMENT The Mark-II Fuel Element and Subassembly

The Mark-II driver-fuel element (see.Fig. 1) consists of a 343 mm long fuel pin of 67%-enriched U-5 wt % Fs that is sodium-bonded within tFissium (Fs) is an equilibrium concentration of fission-product elements

left by the pyrometallurgical reprocessing cycle designed for EBR-II.

It consists of 2.4 wt % molybdenum, 1.9 wt S ruthenium, 0.3 wt % rhodium, 0.2 wt % palladium, 0.1 wt % zirconium, and 0.01 wt % niobium.

605

611 32 -

Sodium Level @ 294K

\~ 12.7 — j - 1 . 7

i —( T_

^•tp-

0.30 Clad Wall 0.25 Sodium Bond

F1g. 1. Mark-II Driver-fuel Element Design.

All dimensions are in millimeters.

fully annealed Type 116 stainless steel cladding. Above the bond sodium is a 2.4 cm3 plenum for the collection of fission gas released from the fuel. A spacer wire of Type 316 stainless steel is attached to the element with a 152-mm helical pitch. Thirteen millimeters above the fuel pin, a restrainer dimple is indented to restrict axial fuel growth and liftoff.

Two of the subassemblies in the experiment--X235 and X275--were normal-configuration Mark-II driver subassemblies consisting of 91 fuel elements in an annealed Type 304 stainless steel hexagonal duct. Zhc third subassemb'!y--X212--consisted of 37 Mark-II elements in a hex duct that was similar but which had a flexible inner liner to prevent element- element interaction due to irradiation dilation. The subassemblies were cooled with primary sodium flowing from the bottom to the top, so the peak cladding temperature was at the top of the fuel.

Irradiation Program and Conditions

All three subassemblies were irradiated in EBR-II reactor row 6.

The subassemblies were orificed to reduce the coolant flow and make the subassemblies operate hotter than normal. The peak burnup occurred 171 mm above the bottom of the fuel column. The ratio of fast fluence (E>0.1 MeV) to burnup was 0.76. All the elements operated at at peak power of 21 kW/m. Table I gives the peak subassembly burnup and beginning- of-life (BOL) peak inside cladding temperature* for all the subassemblies.

End-of-life temperatures ^re approximately 30°C less.

FUEL ELEMENT PERFORMANCE

For purposes of comparison, we will first summarize the performance of a Mark-II fuel element operating at 590°C in row 6 (normal Mark-II).

The performance has been previously documented.1'2'3 Early in life, the fuel pin exhibits some liftoff. The fuel continues to swell anisotropically until it contacts the cladding. By this time, open porosity has developed, fission gas has begun releasing from the fuel, and the bond sodium has been pushed into the plenum region. As irradiation continues, the fuel swelling in the radial direction is limited by the cladding dilation. The fuel swelling in the axial direction continues past the dimple restrainer at a rate of 2 mm per at. % burnup. The bond sodium in the plenum

infiltrates the fuel, so the plenum space gets larger. Concurrently,

*A11 temperatures in this paper will be peak inside cladding temperatures unless stated otherwise.

TABLE I. Operating Parameters for High-temperature Mark-II Elements P-PJ-_ Peak Inside Cladding Temperatures Interior Elements Exterior Elements Subassembly Burnup, at. % T(°C), AT(°C/mm) T(°C'), AT(°C/mm)

X212 6.7 675, 0.8 590, 0.55 X212A 8.0 675, 0.8 590, 0.55

X235a 9.5b 604, 0.6 532, 0.4

X275C 9.9 604, 0.6 532, 0.4

Normal Row 6 — 590, 0.55 528, 0.4 Reconstituted from X235 at 5.8 at. % burnup.

Breached.

Reconstituted from X275 at 7.9 at. % burnup.

the fuel chemically bonds to the cladding, thereby causing a nickel depleted interaction zone in the cladding. The cladding strain is predominantly swelling but it contains a small creep component due to the hoop stress produced by the pressure of the fission gas. The cladding eventually breaches in the dimple restrainer at approximately 10 at. % burnup.2 At end of life, the elements have little bow and almost no ovality.

Limited surveillance of elements operated at 604°C indicated that their performance was similar to that observed for the normal row-6 Mark-II element. Examination of the elements operated at 675°C indicated a substantially different performance, although none of the elements were breached when they were removed from the reactor at 8 at. % burnup.

Differences were observed in the following: the FCCI zone; swelling and creep of the cladding; bond-sodium penetration into the fuel; fission- gas release; axial fuel growth; fuel-porosity distribution; fuel-pin liftoff; and element ovality and bow.

Whole Element Performance

Ovality and Bow. - After irradiation, elements operated at 590°C or less were straight and had less than 1% ovality. In contrast, the elements operated at 675°C exhibited approximately 6.5 mm of bow and

were quite oval (2.5%).l( Even though the subassembly hardware was designed to prevent element-element interaction, interaction did take place because of the unexpectedly large element strain. The large bow and ovality were probably caused primarily by the elements deforming to minimize the interaction. The effect of the element distortion on the coolant flow in the subassembly is unknown.

Swelling and Creep of the Cladding. - On an element irradiated at 675°C the cladding strain is double-peaked (Fig. 2) in contrast to a single peaked profile for an element operated at 590°C.2 The peak at the core midplane is approximately the same magnitude as observed in a normal Mark-II element. (AD/D = 1.6* compared to AD/D = 1.7ft) at the same burnup. The peak at the higher elevation is significantly larger than would be expected.

I

0

O SWELLING

• TOTAL

X CREEP STRAIN

200 300 400 500

-a.

DISTANCE FROM BOTTOM OF FUEL ELEMENT, mm

Fig. 2 Element cladding strain, creep strain and swelling at 675°C and 8 at.% burnup. The total element strain (AD/D ) was determined from four axial profiloinetry traces fit to an ellipse. Creep and swelling were determined by immersion density. The core centerline is at 172 mm.

Both the creep and swelling components of the strain are enhanced at the higher temperatures, the swelling by 80% and the creep by a factor of 10 over that which would be expected in a normal Mark-II element. Because of the large enhancement of the creep at 675°C, creep accounts for 60% of the cladding strain in contrast to 15% of the strain at 590°C.

Fuel Cladding Chemical Interaction

Characteristics. - At 59O°C, the FCCI zone has two regions: a

nickel-depleted ferritic zone in the cladding and a larger nickel enriched zone in the fuel.3 The zone in the fuel will be described later in the sections on fuel porosity. The zone in the cladding with a PCT of 675°C and a burnup of 6.8 at. % represents 25% of cladding thickness, substantially more than the 5% of the cladding thickness the zone occupies at 590°C

PCT and 10 at. % burnup (Fig. 3 ) . At both temperatures the zone has a uniform front with no deep penetration along the grain boundaries. At 675°C the zone is cracked in spots, with the cracks extending the full width of the zone. No cracks were observed at 590°C. Since the zone is cracked and occupies such a large fraction of the cladding thickness at high temperatures, it should be considered as strengthless cladding wastage in determining ultimate cladding lifetimes and cladding stress.

yse_Jj}_J_eH)j)erature_ JteJ;ernnjiation.' - The width of the FCCI zone in the cladding was used to determine the operating temperature of the

high-temperature elements. Temperature calculations5 based on preirradia- tion flow measurements and known pin power ratings indicated that the subassembly was operating at the designated temperature of 604°C. All the surveillance data presented in this paper, especially the FCCI zone width, indicated that the calculated temperatures were too low. At

590°C, the interaction zone1 in the Mark-II cladding agrees with the out-of-pile interaction zone6 if one assumes that the temperature does not decrease during irradiation and is known within + 25°C. The depth of the interaction zones for various elements within the high temperature cubassemblies (X212 and X212A) and for a number of positions along the elements were compared with the out-of-pile interaction zones (Fig. 4) to determine the temperatures and temperature gradients that were listed in Table I.

Fuel__Per f ornw nee

Fuel_-j)in Liftoff. - At 675°C, the fuel pin lifts off 12 mm until it contacts the restrainer dimple designed to prevent liftoff. At 590°C and lower, the fuel pins lift off 2 mm or less on the average; a maximum liftoff of 6 mm has been observed. Since the cladding breach at 590°C

590 °C PCT 10.3 at.% BURNUP

54

675 °C PCT 6.8 ot. % BURNUP

Fig. 3 FCCI zone in the cladding at two different temperatures.

1

ini Q

w

(O

m

100

10

LU

o

. 1

050

000

550 C

^ AH*

X

o

1

130-100 230 305 175

mm INNER ROW mm INNER ROW mm OUTER ROW mm INNER ROW mm INNER ROW OUT-OF-PILE

i

10

10

10

10 8

TIME AFTER FUEL-CLADDING CONTACT.

c .

Fig 4 Depth of FCCI zone in high temperature elements measured by optical metallugraphy at various axial elevations to determine the peak cladding temperature and axial temperature gradient. Inner and outer row refer to position in subassembly.

always occurred in the restrainer dimple, it was recommended that a

restrainer was not necessary.1'2 At the elevated temperatures, a restrainer of some type is probably necessary, because the extent of the fuel-pin liftoff without a restrainer has not been determined.

Bond Sodium Infiltration. - The bond-sodium level above the fuel, which is a measure of sodium infiltration (or logging) of the fuel, was measured by neutron radiography on a number of elements operating at various peak element fuel center!ine-temperatures (FCT). After fuel- cladding contact, the bond sodium infiltrates the open porosity in the fuel at the rate indicated in Fig. 5. This infiltration accomplishes

two things. First, it increases the thermal conductivity of the fuel and thus reduces the thermal gradient across the fuel. Second, it increases the plenum volume available for occupation by fission gas As a result, the greater the sodium infiltration at any burnup, the smaller the resulting cladding stress due to fission gas pressure. For the high-temperature elements, neither of these beneficial aspects of sodium logging occurs because all the bond sodium resides in the plenum region and does not infiltrate the fuel. In contrast, for a normal Mark-II element, 7% of the bond sodium infiltrates the fuel per at. % burnup.

E

§

0

-1

O - AXIAL FUEL GROWTH X - No LOGGING

8

-2

1

m

8?

I

o o o

3?

400 500 600 700

PEAK ELEMENT FUEL-CENTERLINE TEMPERATURE. °C

Fig. 5 Axial fuel growth and bond sodium infiltration (logged) into the fuel. The apparent increase in nonlogged sodium volume at 720°C is due to some solid fission products, especially 1 3 7C s , leaching out of the fuel into the bond sodium. This occurs at all temper- atures but is measurable only when there is no sodium infiltration.

For the Na logging curve, burnup is measured from the time of fuel- cladding contact.

Axial Fuel Swelling. - Contrary to unrestrained swelling,7 measure- ments of anisotropic fuel swelling has been observed in Mark-11 elements operating at 590cC, with the swelling rate in the radial direction being much larger than that in the axial direction.1 Anisotropic swelling is observed at the higher temperatures also, but in a larger degree because there is practically no axial fuel swelling (see Fig. 5 ) . •

Fission-gas Release. - Two elements with 7.7 at. % burnup were laser-punctured to determine the amount of gas in the plenum and the percentage of the gas that was released from the fuel. At 675°C only 40 cm3 of gas were in the plenum, corresponding to a gas release of 45%.

At 590°C and 8 at. 1 burnup, 68% of the fission gas were released from the fuel which resulted in 60 cc of gas in the plenum with a pressure of 12.0 MPa.8 When both the smaller plenum volume due to lack of soc ium infiltration into the fuel and the higher operating temperatures tire taken into account, the fission-gas pressure in the high-temperature element is the same as that in a normal temperature Mark-II.

Fuel Porosity. - At elevated temperatures, the fuel appears divided into five different zones. One is the previously mentioned fuel zone next to the cladding. The location and interpretation of the other four zones, which are axially segmented, are based on phase changes and

porosity distribution.

The zone next to the cladding, shown in Fig. 6(a), decreases in width and increases in porosity as one nears the bottom of the fuel pin, where the temperature decreases. At the top of the fuel pin this zone represents 45* of the fuel volume and has very little porosity. At

about 224 mm from the bottom of the fuel pin, the temperature has decreased to the same temperature as the top of a normal temperature row-6 Mark-II element. From this point down to the bottom of the pin, the width and porosity characteristics of the interaction zone are the same as those in normal temperature Mark-II element.

The first of the four axial zones occupies the bottom 100 mm of the fuel pin. This zone consists of a band of rather coarse porosity sur- rounding a zone of finer open porosity and is characteristic of a normal Mark-II element. The coarse porosity appears to line up on the fuel-

element isotherms. The cause for this pattern has not yet been investigated.

The second axial zone, shown in Fig. 6{b), occupies the next 124 mm from the pin bottom. It has the rather uniform extensive open porosity that is characteristic of the porosity to be expected in a normal Mark-II ele- ment. These two zones are probably all a-uranium at operating temperature.

The third axial zone, shown in Fig. 6(c), is marked by the phase

change a •» y + l^Ru- Out-of-pile studies indicate that the ILRu precipitates

a.) INTERACTION ZONE C.) PHASE CHANGE ZONE 3

b.) OPEN POROSITY ZONE 2 <*•> CLOSED POROSITY ZONE 4

F i g . 6 Fuel zones o f d pin i r r a d i a t e d at 675°C to 8 a t . * burnup.

should be the light acicular structures* shown in Fig. 6(c). By 8 at. % burnup, the fuel has a composition of U-9 wt % Fs. For this composition there is a complete change to the y phase at a fuel- centerline operating temperature of 620°C.7 The majority of the y-p

likely transforms back to the a-phase during the prolonged post-irradiation residence of the subassembly in the storage basket at 370°C. The ghosts of this probable phase-change mark the bottom of the third axial zone.

The location of this phase change is consistent with the temperatures deduced from the FCCI zone. The top of a normal temperature fuel pin would also have the transformation, but it is usually not seen because that area has extensive open porosity.

The fourth axial zone, shown in Fig. 6(d), occupies the top 30 mm of the fuel pin. In a normal temperature Mark-II element, most of the porosity in this zone is open; in a high-temperature Mark-II element, only 20-40% of the porosity is open.

DISCUSSION Fission-gas Release

The lack of sodium infiltration of the fuel, lack of open porosity, and reduced fission g^.s release are consistent with each other. The initial irradiation performance is thought to be the same in both the Mark-II elements operating at 675°C and 590°C. The fuel swells out to the cladding and pushes the bond sodium up into the plenum region. In an element operating at 590°C, open porosity would develop at this stage, gas would be released to the plenum, and sodium would begin to infiltrate the fuel.

The fission-gas-release characteristics of the high-temperature elements are different from those described above, but can be interpreted rather simply. Consider the fuel to be composed of three parts. The first part, consisting of axial zones 1 and 2, has gas-releise character- istics similar to those of a normal Mark-II element. It contains 2.55 cm3

of fuel, which releases 68% (15.86 cm3 of gas/cm3 of fuel) of its fission gas. The second part of the fuel is the upper 120 mm of the zone next to the cladding. It contains 0.6 cm3 of fuel, but releases no gas since it has no open porosity. The third part, composed of axial zones 3 and 4, contains 0.75 cm3 of fuel having a mixture of open and closed porosity

in an unknown ratio. With a 45% fission-gas-release ratio for the full fuel pin, 43 cm3 of gas are released. The gas release from the fuel in zone 1 and 2 can account for 40.4 cm3 so 2.6 cm3 of gas would have to be produced by the region of mixed porosity. This amount of gas corresponds

to 3.4 cm3 of released gas/cm3 of fuel, which is only 21% of the release in a Mark-II element operating at 590°C. With possible uncertainties in the measurements this release can range from 20 to 40% of the normal Mark-II release. This range of gas release corresponds to open porosity of 14 to 28% which agrees with the open porosity of 20-40% observed in axial zone 4. Hence the lower fission-gas release is duetto the lack of open porosity in the top of the fuel pin. This lack of open porosity at the top of the fuel pin is also the reason that the sodium does not infiltrate the fuel.

FCMI

One of the primary beliefs concerning metal fuels has been that there is little or no FCMI with such fuels. For normal Mark-II tempera- tures, there is a wealth of evidence to justify this assertion but similar evidence is not available at higher temperatures.

The high temperature elements exhibit a double peaked diameter profile with a peak above the core centerlin? which is not found at 590°C. Since

the flux peaks at the core midplane and the cladding temperature peaks at the core top, the temperature dependence of in-reactor creep may dominate the flux dependence so a second peak in the diameter profile appears.

Application of available creep correlations12'13, with plenum pressure as the only source of stress, resulted in an underprediction of the second peak creep strain. This may result from the combination of the following three phenomena.

1. The position of the second peak is at the point of maximum carbide precipitation in the cladding.11* The creep and swelling components of the strain may have been enhanced by the carbon depletion from the matrix.

2. The FCCI is 25% of the cladding thickness in the region of the second peak compared to 5" of the cladding thickness at the core midplane. If the cracked interaction zone has no strength, then the stress due to fission gas pressure would be higher in the vicinity of the second peak, and more creep strain would result.

3. It has been shown that the interconnected porosity decreases above the core centerline. This implies that the fuel may retain some strength and exert stress on the cladding due to FCMI as was observed on the Mark-IA elements where inter- connected porosity and gas release did not occur.

At this stage in the study it is neither possible to identify the dominant mechanism for the double peaked profile nor is it possible to

estimate the magnitude of the FCMI if it is present. At the high temperatures though the cladding is able to successfully respond to any FCMI that is present, since none of the elements breached at the attained burnup of 8 at.%.

CONCLUSIONS

The performance of Mark-II fuel at 675°C is significantly different than that at 590°C. Despite the differences, the elements were all successfully irradiated to 8 at. % burnup before being removed from the reactor unbreached.

Some of the more significant differences in performance at the higher temperatures are:

1. The fuel-element cladding showed a double-peaking strain profile, with enhanced swelling and creep due to extensive carbide precipitation, an FCCI enhanced fission gas stress, or FCMI.

2. The FCCI zone in the fuel was cracked and represented 25" of the cladding thickness. This zone was substantially larger than observed at 590°C.

3. The fuel pin lifted 12 mm to the restrainer dimple, so some type of restrainer may be necessary in future designs. A restrainer had previously been deemed unnecessary based on irradiations at 590°C.lf2

4. The fission-gas release is less at a fuel-centerline temperature of 720°C compared to 650°C. This difference is primarily due to the upper third of the element having very little open porosity. Owing to the lack of bond sodium infiltration into the fuel at the higher temperature, the fission gas pressure is the same at both temperatures.

ACKNOWLEDGMENT

The author thanks Drs. J. Flinn, G. McVay, L. Walters, and B. Seidel for many helpful discussions. The efforts of Mr. E. Wood for the carbide analysis, Mr. B. Carlson and Mr. C. Wood for metallography, Mr. G. Hudman for data reduction, and Mr. R. Keyes of the Hot Fuel Examination Facility hot cells have been invaluable to this work and are greatly appreciated.

This work is supported by the United States Department of Energy.

REFERENCES

1. R. E. Einziger and B. R. Seidel, "The Performance of Reference Design Mark-II Metallic Driver Fuel Elements in EBR-II to 10 at. % Burnup," to be submitted to Nuclear Techoiogy.

2. B. R. Seidel and R. E. Einziger, "In-reactor Cladding Breach of EBR-II Driver-fuel Elements," Radiation Effects in Breeder Structural Materials, M. L. Bleiberg and J. W. Bennett, Eds., TMS-AIME New York, p. 139 (1977).

3. G. L. Hofman, W. N. Beck, R. V. Strain, G. 0. Hayner, and C. M.

Walter, "Irradiation Behavior of Unencapsulated EBR-II Mark-II Driver Fuel to a Maximum Burnup of 6 at. %," ANL-8119 Feb. 1976.

4. R. E. Einziger, "Surveillance of Subassembly X212," Reactor Development Program Progress Report, March 1977, ANL-RDP-59, p. 1.12 (May 20", 1977).

5. J. Gillette, private communication (April 1977).

6. S. T. Zegler, H. V. Rhude, and J. P. Lahti, "Compatibility of Uranium-5 wt "!,- Fission Alloy with Types 304L and 316 Stainless Steel," ANL-7596 (Sept. 1969).

7. J. H. Kittel, J. A. Horak, W. N. Beck, and R. J. Fousek, "Irradiation Behavior of Uranium-Fissium Alloys," ANL-6795 (Oct. 1971).

8. R. E. Einziger, "Surveillance of Plenum Pressures in Reference- design Mark-II Driver Fuel," Reactor Development Program Progress Report, July-August 1976, ANL-RDP-52", p. 1.24 (Oct. 19, 1976).

9. W. D. Wilkinson, Uranium Metallurgy, Vol. II, Interscience Publishers, New York, (1962).

10. John Flinn, private communication (1978).

11. G. L. Hofman, L. C. Walters, and G. L. McVay, "The Effect of Phase Instability of Stainless Steel on the Cladding Deformation Profiles

• of LMFBR Fuel Elements," J. Nucl. Mater 67., p. 289-294 (1977).

12. E. R. Gilbert and A. J. Lovoll, "Temperature Dependence of In- rcactor Creep of 20> Cold Worked 316 Stainless Steel," Radiation Effects iji Breeder Reactor Structural Materials, M. L. Bleiberg and J. "W~ Bennett',' EdV.'f~TMS-"ATMEY"New York", pV 269 (1977).

13. J. L. Straalsund, "Irradiation Creep in Breeder Reactor Structural Materials," Radiation Effects jj^ Breeder Reactor Structural Materials, M. L. Bleiberg and J. W. Bennett, Ed"s., TMS-AIF1E, New York, (1977) p. 191.

14. G. L. McVay, R. E. Einziger, G. L. Hofman, and L. C. Walters, "The Relationship Between Carbide Precipitation and the.In-reactor Deformation of Type 316 Stainless Steel," accepted for publication in J. Nucl. Mater.