ML21144A135

TLR-RES/DE/REB-2021-02: Assessment of Non-Gross Breaches in Spent Fuel Cladding During Drying Operations

ML21144A135
Person / Time
Issue date:	05/26/2021
From:	Patrick Raynaud NRC/RES/DE/CIB
To:
Raynaud P
Shared Package
ML21167A218	List: ML21144A135
References
UNR NMSS-2017-001 TLR-RES/DE/REB-2021-02
Download: ML21144A135 (30)

Text

Technical Letter Report TLR-RES/DE/REB-2021-02 Assessment of Non-Gross Breaches in Spent Fuel Cladding during Drying Operations Date:

Drafted: June 2019 Updated: March 2021 Completed: May 26, 2021 Prepared in support of task 3b-3-b in UNR NMSS-2017-001 by:

Patrick Raynaud Senior Materials Engineer Component Integrity Branch Brendan Dowling, and Matt Bisbee Summer Interns Component Integrity Branch Division of Engineering Office of Nuclear Regulatory Research U.S. Nuclear Regulatory Commission Washington, DC 20555-0001

DISCLAIMER This report was prepared as an account of work sponsored by an agency of the U.S. Government.

Neither the U.S. Government nor any agency thereof, nor any employee, makes any warranty, expressed or implied, or assumes any legal liability or responsibility for any third party's use, or the results of such use, of any information, apparatus, product, or process disclosed in this publication, or represents that its use by such third party complies with applicable law.

i

This report does not contain or imply legally binding requirements. Nor does this report establish or modify any regulatory guidance or positions of the U.S. Nuclear Regulatory Commission and is not binding on the Commission.

ii

1 Background

This report summarizes the initial RES/DE/CIB efforts to address task 3b-3-b from UN NMSS-2017-001:

Task 3b: Spent Fuel Cladding Performance during DCSS Draining and Drying

3. FY17-FY19: Assess whether non-gross breaches (hairline cracks, pinholes):

b. can grow and develop into gross ruptures during drying operations (i.e., crack growth is self-limited and arrested before exceeding the gross rupture criterion of 1 mm, per ISG-1, Rev. 2).

The task will provide confirmatory analyses that cladding gross ruptures are not expected during drying operations and that fuel being loaded will retain its analyzed configuration. The integrity of the cladding is to be protected against gross ruptures per 10 CFR 72.122(h)(1) and the fuel configuration is relied on in the safety analyses for providing assurance of compliance with various system-related requirements (e.g.

l10 CFR 72.122(h)(5) and 10 CFR 72.124(a)).

2 Analytical Approach In order to assess the potential evolution of non-gross breaches in spent fuel cladding during drying operations, the potential breaches were divided into two categories: hairline cracks and pinholes. The analytical approach consisted of two parallel tasks:

• One task focused on axial hairline cracks and consisted of calculating the stress intensity factors associated with various postulated axial cracks in the cladding, subjected to a temperature-dependent bounding rod internal pressure that would result in a bounding hoop stress for drying operations at 400°C, 425°C, and 450°C. The calculated stress intensity factors were used to determine which postulated cracks may be expected to grow.

• The other task focused on pinhole defects and consisted of creating a finite element model of a simplified fuel rod subjected to a bounding rod internal pressure that would result in a bounding hoop stress for drying operations. From this model, stress concentration factors were calculated for a variety of pinhole sizes to show that a pinhole defect would either evolve into an axial crack or be inconsequential.

In both cases, the bounding pressure and stress values were chosen based on the results from reports PNNL27418 [1] and PNNL-30430 [2]. All stress intensity factor calculations were performed using the solutions from API-579 Appendix C [3]. All finite element modeling was performed using the ABAQUS 2018 software [4].

The maximum stress intensity factor values calculated for the scenarios above were compared to the expected range of fracture toughness of the cladding at the corresponding temperature. Such 1

toughness is highly dependent on burnup and hydride configuration, as well as the temperature during the drying transient. Since relevant toughness data was difficult to find, a conservative approach was adopted.

3 Assessment of Non-Gross Breaches 3.1 Axial Hairline Cracks 3.1.1 Stress Intensity Factor Calculations An axial hairline defect in a cladding tube is analog to an axial through-wall crack in a cylinder. Stress intensity factor (KI) solutions for such a configuration are well known, and NRC has approved the solutions from API-579 [3] (used by ASME) for such evaluations. For a cylinder (i.e. tube) with a through-wall crack in the axial direction (Figure 1), the stress intensity factor is given by Equation 1.

Figure 1: Cylinder - Through-wall Crack, Axial Direction For the purpose of this analysis, it is assumed that the hoop stress is due to internal pressure in the rod, with the pressure being adjusted to ensure that a bounding stress (and thus a bounding KI) is obtained (from [1] and [2]). Under these assumptions, the equation simplifies to Equation 2.

The influence coefficient Gp is calculated with Equation 3 and Equation 4. The constants within Equation 3 are given in Table 1 (at tube ID) and Table 2 (at tube OD). The constants were interpolated to the proper t/Ri ratio, and then averaged between the ID and OD values to get a value for A0 through A6 at the cladding mid-wall, which is the relevant value for a through wall crack.

In this study, the worst case from [1] was chosen: a fuel rod from a PWR 17x17 assembly design. The dimensions chosen are those typical for such designs: a rod internal radius Ri = 4.15 mm and a cladding wall thickness of t = 0.6 mm, resulting in t/Ri = 0.1446.

2

Equation 1: Stress Intensity Factor for an Axial TWC Equation 2: Reduced Stress Intensity Factor for an Axial TWC Equation 3: Influence Coefficient Equation 4: Influence Coefficient Parameter 3

Table 1: ID Influence Coefficient for Axial TWC 4

Table 2: OD Influence Coefficient for Axial TWC From the equations presented above, the stress intensity factor KI was calculated for crack lengths from 2c = 0.2 mm to 2c = 200 mm. It is important to note that a 200 mm axial crack is well beyond the threshold for a gross rupture, however, these calculations were performed to assess the critical crack length that might evolve into a gross defect, based on assumed values of toughness for the high-burnup spent fuel cladding.

3.1.2 Fracture Toughness of High Burnup Spent Fuel Cladding In order to assess whether an axial hairline crack would evolve into a gross breach, the applied stress intensity factor must be compared to the fracture toughness of the material for crack growth in the axial direction. The exact fracture toughness of the cladding for axial growth of a through-wall crack depends, among others, on the temperature, fluence, hydrogen content, hydride orientation, and oxide thickness. These parameters can vary significantly in spent fuel depending on the fuel design, cladding material, and operational history of each fuel rod. As a result, it is impossible to know exactly the relevant values of fracture toughness, and a conservative approach was used in this assessment, based on available literature and engineering judgment.

A detailed literature survey for cladding fracture toughness was performed in 2009 [5], covering a variety of test specimens, crack growth directions, test temperatures, and material conditions. Overall, 5

for axial crack growth, the reported fracture toughness values above 300°C were in the range of 89 MPam to 150 MPam. In general, hydrogen, hydride orientation, oxygen, and fluence all tend to reduce the fracture toughness as they increase. It was observed that the influence of these parameters generally decreases with increasing temperature, and in particular, because it is mostly dissolved at 400°C, the deleterious effects of hydrogen on toughness are diminished at this temperature.

Furthermore, the reported fracture toughness in a different direction of crack growth (through-thickness) did not change significantly between 300°C and 375°C.

Based on the above information, it is conservatively estimated that the fracture toughness of spent fuel cladding at 400°C in the axial crack growth direction is at least 75 MPam and could be as high as 125 MPam or more. It is hoped that planned testing at ORNL on actual spent fuel cladding will provide better estimates of relevant toughness values.

The toughness of the material was assumed to decrease above 400°C, due to the material progressively losing its strength but remaining highly ductile. The rate of toughness decrease should be roughly equivalent to the rate of strength decrease in this temperature range. Based on the material properties of zircaloy as implemented in the FRATRAN code, the cladding strength decreases by a little less than 10% per 25°C between 400°C and 450°C, thus it is conservatively assumed here that the toughness will decrease at a rate of 10% per 25°C between 400°C and 450°C.

Using a range of 75-125 MPam at 400°C results in toughness ranges of 67.5-112.5 MPam at 425°C, and 60.75-101.25 MPam at 450°C.

3.1.3 Results and Discussion Figure 2 through Figure 4 show the calculated stress intensity factor as a function of axial through-wall crack length, respectively assuming a cladding hoop stress of 90 MPa at 400°C, 94 MPa at 425°C, and 98 MPa at 450°C (corresponding to rod internal pressures of 11.368 MPa, 11.874 MPa, and 12.379 MPa),

and Table 3 shows the corresponding calculated stress intensity factor values. The stress values were chosen because they bound the calculated maximum hoop stresses at the corresponding maximum temperatures, as reported by PNNL in [1] and [2]. Cracks shorter than 10 mm result in an applied KI of 35-40 MPam, a 25 mm crack results in KI of around 105-115 MPam, and the applied KI continues to increase with crack length beyond this point.

In Figure 2 through Figure 4 and in Table 3, the color coding in the stress intensity factor KI column is as follows:

• Green indicates that crack growth is not expected to occur

• Yellow indicates that crack growth may occur based on conservative fracture toughness estimates

• Red indicates that crack growth is likely to occur Based on this assessment, it is expected that any axial hairline crack that is less than about 15-18 mm in length will not evolve during drying operations. Conversely, any crack that is initially longer than about 22-27 mm is likely to grow further during drying operations. In summary, assuming the cladding is initially free of hairline cracks less than 15-18 mm in length, drying operations are not expected to result in the creation of longer cracks with sufficient opening widths to give rise to gross defects.

6

Figure 2: Calculated KI for an Axial Crack using API-579, for a maximum hoop stress of 90 MPa, corresponding to a maximum temperature of 400°C. The color coding represents the likelihood of crack growth occurring, and the horizontal dashed lines represent an estimate of the conservative lower and upper bounds of the cladding toughness at this temperature.

7

Figure 3: Calculated KI for an Axial Crack using API-579, for a maximum hoop stress of 94 MPa, corresponding to a maximum temperature of 425°C. The color coding represents the likelihood of crack growth occurring, and the horizontal dashed lines represent an estimate of the conservative lower and upper bounds of the cladding toughness at this temperature.

8

Figure 4: Calculated KI for an Axial Crack using API-579, for a maximum hoop stress of 98 MPa, corresponding to a maximum temperature of 450°C. The color coding represents the likelihood of crack growth occurring, and the horizontal dashed lines represent an estimate of the conservative lower and upper bounds of the cladding toughness at this temperature.

9

Table 3: Calculations for Stress Intensity Factor in Axial Direction Stress Intensity Factor for Axial Crack KI (MPam) KI (MPam) KI (MPam)

Crack Length [hoop=90 MPa] [hoop=94 MPa] [hoop=98 MPa]

2c (mm) [PCT=400°C] [PCT=425°C] [PCT=450°C]

0.2 1.6 1.7 1.8 0.5 2.6 2.7 2.9 1 3.9 4.1 4.2 2 6.3 6.6 6.9 5 15.6 16.3 17.0 10 35.4 37.0 38.6 15 57.6 60.2 62.7 20 80.9 84.5 88.1 25 104.9 109.5 114.2 30 129.4 135.2 140.9 40 179.8 187.8 195.8 50 231.6 241.9 252.2 60 284.6 297.2 309.9 70 338.7 353.8 368.8 80 394.0 411.5 429.0 90 450.5 470.5 490.5 100 508.3 530.9 553.5 120 628.2 656.1 684.1 140 755.2 788.7 822.3 160 890.7 930.3 969.9 180 1036.8 1082.9 1129.0 200 1195.9 1249.1 1302.2 3.2 Circumferential Hairline Cracks 3.2.1 Stress Intensity Factor Calculations A circumferential hairline defect in a cladding tube is analog to a circumferential through-wall crack in a cylinder. Stress intensity factor (KI) solutions for such a configuration are well known, and NRC has approved the solutions from API-579 [3] (used by ASME) for such evaluations. For a cylinder (i.e. tube) with a through-wall crack in the circumferential direction (Figure 5), the stress intensity factor is given by Equation 5.

10

Figure 5: Cylinder - Through-wall Crack, Circumferential Direction For the purpose of this analysis, it is assumed that the hoop stress is due to internal pressure in the rod, with the pressure being adjusted to ensure that a bounding stress (and thus a bounding KI) is obtained (from [1] and [2]). Under these assumptions, the equation simplifies to Equation 6.

The influence coefficient G0 is calculated with Equation 3 and Equation 4. The constants within Equation 3 are given in Table 4 (at tube ID) and Table 5 (at tube OD). The constants were interpolated to the proper t/Ri ratio, and then averaged between the ID and OD values to get a value for A0 through A6 at the cladding mid-wall, which is the relevant value for a through wall crack.

In this study, the worst case from [1] was chosen: a fuel rod from a PWR 17x17 assembly design. The dimensions chosen are those typical for such designs: a rod internal radius Ri = 4.15 mm and a cladding wall thickness of t = 0.6 mm, resulting in t/Ri = 0.1446.

Equation 5: Stress Intensity Factor for a Circumferential TWC Equation 6: Reduced Stress Intensity Factor for a Circumferential TWC 11

Table 4: Inside Diameter Influence Coefficient for Circumferential TWC 12

Table 5: Outside Diameter Influence Coefficient for Circumferential TWC From the equations presented above, the stress intensity factor KI was calculated for crack lengths from 2c = 0.2 mm to 2c = 26 mm. It is important to note that a 26 mm axial crack implies the fuel rod is almost completely ruptured. However, these calculations were performed to assess the critical crack length that might evolve into a gross defect, based on assumed values of toughness for the high-burnup spent fuel cladding.

3.2.2 Fracture Toughness of High Burnup Spent Fuel Cladding Data on fracture toughness of spent fuel cladding above 400°C for circumferential crack growth are not available. However, based on a comparison of toughness of zircaloy-2 for different specimen orientations, the fracture toughness in the circumferential crack growth direction appears to be approximately 75% of the fracture toughness for axial crack growth [6].

Based on this limited information, it is conservatively estimated that the fracture toughness of spent fuel cladding at 400°C in the circumferential crack growth direction is at least 56.25 MPam and could be as high as 93.75 MPam or more. Similarly, the toughness range is estimated to be 50.684.4 MPam at 425°C, and 45.675.9 MPam at 450°C.

13

3.2.3 Results and Discussion Figure 6 through Figure 8 show the calculated stress intensity factor as a function of circumferential through-wall crack length, assuming a cladding axial stress of 90 MPa at 400°C, 94 MPa at 425°C, and 98 MPa at 450°C (corresponding to rod internal pressures of 22.737 MPa, 23.747 MPa, and 24.758 MPa, double that required to get a similar hoop stress, and thus a very conservative approach). Table 6 shows the corresponding calculated stress intensity factor values. The stress values were chosen because they bound the calculated maximum hoop stresses at the corresponding maximum temperatures, as reported by PNNL in [1] and [2] (corresponding to a 17x17 IFBA fuel rod design). It is important to note that in a pressurized cylinder (i.e. a closed fuel rod), the axial stress is normally half of the hoop stress, so assuming that the axial stress is equal to the hoop stress is highly conservative. Cracks shorter than 12-13 mm result in an applied KI of less than 45.6-56.25 MPam (the lower bound estimated toughness),

a 1618 mm crack results in KI of around 75.9-93.75 MPam (the upper bound estimated toughness),

and the applied KI increases very rapidly beyond this point, as the remaining fuel rod cross circumference rapidly diminishes as the crack length continues to increase.

In Figure 6 through Figure 8 and in Table 6, the color coding in the stress intensity factor KI column is as follows:

• Green indicates that crack growth is not expected to occur

• Yellow indicates that crack growth may occur based on conservative fracture toughness estimates

• Red indicates that crack growth is likely to occur Based on this assessment, it is expected that any circumferential hairline crack that is less than about 12-13 mm in length (representing about half of the fuel rod circumference) will not evolve during drying operations. Conversely, any crack that is initially longer than about 16-18 mm is likely to grow further during drying operations. In summary, assuming the cladding is initially free of circumferential hairline cracks less than 12-13 mm in length, drying operations are not expected to result in the creation of longer cracks with sufficient opening widths to give rise to gross defects.

14

Figure 6: Calculated KI for a Circumferential Crack using API-579, for a maximum axial stress of 90 MPa, corresponding to a maximum temperature of 400°C. The color coding represents the likelihood of crack growth occurring, and the horizontal dashed lines represent an estimate of the conservative lower and upper bounds of the cladding toughness at this temperature.

15

Figure 7: Calculated KI for a Circumferential Crack using API-579, for a maximum axial stress of 94 MPa, corresponding to a maximum temperature of 425°C. The color coding represents the likelihood of crack growth occurring, and the horizontal dashed lines represent an estimate of the conservative lower and upper bounds of the cladding toughness at this temperature.

16

Figure 8: Calculated KI for a Circumferential Crack using API-579, for a maximum axial stress of 98 MPa, corresponding to a maximum temperature of 450°C. The color coding represents the likelihood of crack growth occurring, and the horizontal dashed lines represent an estimate of the conservative lower and upper bounds of the cladding toughness at this temperature.

17

Table 6: Calculations for Stress Intensity Factor in Circumferential Direction Stress Intensity Factor for Circumferential Crack KI (MPam) KI (MPam) KI (MPam)

Crack Length

[axial=90 MPa] [axial =94 MPa] [axial =98 MPa]

2c (mm)

[PCT=400°C] [PCT=425°C] [PCT=450°C]

0.2 1.6 1.7 1.8 0.5 2.6 2.7 2.9 1 3.8 4 4.1 2 5.7 6 6.2 3 7.5 7.9 8.2 4 9.5 9.9 10.3 5 11.6 12.1 12.6 6 14 14.6 15.2 7 16.7 17.5 18.2 8 19.8 20.7 21.6 9 23.3 24.4 25.4 10 27.4 28.6 29.8 11 32 33.4 34.8 12 37.3 39 40.6 13 43.4 45.4 47.3 14 50.5 52.8 55 15 58.9 61.5 64.1 16 68.7 71.8 74.9 18 95 99.2 103.5 20 136 142 148 22 208.3 217.6 226.9 24 371.3 387.8 404.3 26 1094.5 1143.1 1191.8 3.3 Pinhole Defects 3.3.1 Finite Element Modeling Because pinhole defects do not constitute a sharp crack, stress intensity factors cannot be calculated for such defects. Consequently, it was decided to create a finite element model of a fuel rod with a circular pinhole defect and study the stress concentration resulting from such a defect, to understand how the pinhole defect might evolve.

The finite element model consisted of a pressurized tube with a small circular through-wall hole.

Several models with increasing hole sizes were created, with hole diameters equal to 0.1 mm, 0.2 mm, and 0.5 mm. The pressure in the tube was set to 11.3685 MPa, the value of pressure theoretically resulting in a 90 MPa hoop stress. This resulted in a hoop stress of approximately 84.7 MPa in the finite element model, thus validating the accuracy of the model.

18

Figure 10, through Figure 14 respectively show the hoop stress and Von Mises stress at the location of the hole, for holes of diameter 0.1 mm, 0.2 mm, 0.5 mm, 1.0 mm, and 2.0 mm. In all cases, both the hoop stress and Von Mises stress map clearly shows that the highest stress region is at the ends of the hole that are aligned with the axial direction, and closer to the ID than the OD of the tube. This implies that if the stress near the pinhole was sufficient to cause the cladding to deform or locally fail, the failure would occur in a way that would cause the pinhole defect to evolve into an axial crack. The maximum predicted local stress in this study for a far field stress of 84.7 MPa was 347 MPa. For a far field stress of 98 MPa, assuming a linear scaling relationship, the maximum local stress is expected to be about 405 MPa. For high burnup spent fuel PWR cladding (cold-worked stress relieved Zircaloy-4 with 1.2x1026 n/cm2 fluence and 500 wt.ppm hydrogen), the yield stress is expected to exceed 520 MPa, which is significantly higher than the maximum expected local stress. As a result, even a 2.0 mm diameter hole is not expected to cause local cladding deformation or failure, nor result in any gross cladding ruptures or defects.

3.3.2 Stress Concentration Results and Assessment Table 7 and Figure 9 show the calculated stress concentration factors. Importantly, for the hoop stress, the stress concentration factors for the 0.1 mm, 0.2 mm, 0.5 mm, 1.0 mm, and 2.0 mm pinholes are in the range of 2 to 4, which is in the expected range and confirms that any existing pinhole defect would preferentially evolve into an axial crack. Axial cracks were analyzed in section 3.1 and it was concluded that axial cracks less than 20 mm in length would not pose a problem during drying operations. In fact, if one assumes a pinhole defect 1 mm in diameter and that 1 mm axial cracks have formed on each end of the pinhole in the axial direction, for a total axial crack length of 3 mm including the pinhole, the resulting stress intensity factor would be between 6 MPam and 17 MPam (see Table 3), which is not sufficient to further grow the defect during drying operations.

Table 7: Stress Concentration Factors for Several Modeled Pinhole Sizes ID Hoop Stress ID Von Mises Stress Pinhole Far-field Max Concentration Far-field Max Concentration Size (mm) (MPa) (MPa) Factor (MPa) (MPa) Factor 0.1 84.7696 179.725 2.12 83.1772 147.385 1.772 0.2 84.7625 171.944 2.029 83.1555 153.383 1.845 0.5 84.7625 213.99 2.525 83.1555 196.51 2.363 1.0 84.7625 257.026 3.032 83.1555 244.442 2.94 2.0 84.7625 346.932 4.093 83.1555 341.983 4.113 19

Figure 9: Stress concentration factor as a function of pinhole diameter, for ID hoop stress and ID Von Mises stress.

20

Figure 10: Hoop stress (top) and Von Mises stress (bottom) at the location of the pinhole defect for a 0.1 mm hole diameter 21

Figure 11: Hoop stress (top) and Von Mises stress (bottom) at the location of the pinhole defect for a 0.2 mm hole diameter 22

Figure 12: Hoop stress (top) and Von Mises stress (bottom) at the location of the pinhole defect for a 0.5 mm hole diameter 23

Figure 13: Hoop stress (top) and Von Mises stress (bottom) at the location of the pinhole defect for a 1.0 mm hole diameter 24

Figure 14: Hoop stress (top) and Von Mises stress (bottom) at the location of the pinhole defect for a 2.0 mm hole diameter 25

4 Summary and Conclusions In order to assess whether non-gross breaches in spent fuel cladding would evolve into gross breaches during drying operations, RES performed fracture mechanics analyses on axial cracks and circumferential cracks, and performed finite element stress analyses on pinhole defects.

Based on stress concentration analyses from finite element modeling results, if pinhole defects below the 1 mm gross defect size threshold were to evolve, they would evolve into axial cracks prior to causing any type of cladding failure. However, even pinholes as large as 2 mm are not predicted to cause sufficient local stresses to locally deform or fail the cladding, thus pinhole defects are not expected to evolve into gross ruptures or defects.

For hairline defects, it was shown that for axial crack lengths smaller than 15 or 18 mm and for circumferential crack lengths smaller than 12 or 13 mm, the defects are not expected to evolve during drying operations.

26

5 References

[1] D. J. Richmond and K. J. Geelhood, "PNNL-27418: FRAPCON Analysis of Cladding Performance during Dry Storage Operations," Pacific Northwest National Laboratory, Richland, WA, April 2018.

[2] B. E. Wells, N. R. Phillips and K. J. Geelhood, "PNNL-30430: Evaluation of Increased Peak Temperatures for Spent Fuel Cladding Performance during Dry Storage," Richland, WA, 2019.

[3] ASME, "API-579-1/ASME FFS-1: Firness-for-Service," American Petroleum Institute Publishing Services, Washington, DC, 2007.

[4] SIMULIA, "ABAQUS," SIMULIA, 19 June 2019. [Online]. Available: https://www.3ds.com/products-services/simulia/products/abaqus. [Accessed 19 June 2019].

[5] P. A. C. Raynaud, "Crack Growth Through the Thickness of Thin-Sheet Hydrided Zircaloy-4," The Pennsylvania State University, State College, PA, 2009.

[6] T. J. Walker, "Characterization of the Fracture Toughness of Zircaloy," Nuclear Technology, vol. 16, no. 3, pp. 509-520, 1972.

27