ML24176A124

NUREG/KM-0019, Technical and Regulatory Bases for Regulatory Guide 1.236, PWR Control Rod Ejection and BWR Control Rod Drop Accidents

ML24176A124
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Issue date:	06/30/2024
From:	Paul Clifford, Joseph Messina Office of Nuclear Reactor Regulation
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Malone, Tina; Sinclair, LaToya
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NUREG/KM-0019
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NUREG/KM-0019 Technical and Regulatory Bases for Regulatory Guide 1.236, PWR Control Rod Ejection and BWR Control Rod Drop Accidents Office of Nuclear Reactor Regulation

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NUREG/KM-0019 Technical and Regulatory Bases for Regulatory Guide 1.236, PWR Control Rod Ejection and BWR Control Rod Drop Accidents Office of Nuclear Reactor Regulation Manuscript Completed: June 2024 Date Published: June 2024 Prepared by:

P. Clifford (Retired)

J. Messina

iii ABSTRACT The purpose of this report is to document the technical and regulatory bases for the acceptance criteria and guidance provided in Regulatory Guide 1.236, Pressurized-Water Reactor Control Rod Ejection and Boiling-Water Reactor Control Rod Drop Accidents. Pressurized-water reactor control rod ejection and boiling-water reactor control rod drop accidents are low-probability events involving a sudden and rapid insertion of positive reactivity. These postulated accidents are safety significant because of their potential ability to cause significant core damage and challenge the integrity of the reactor coolant pressure boundary. This report identifies regulatory requirements, details the existing empirical database of fuel performance under rapid power excursion testing, and documents the technical bases of the analytical limits and guidance that satisfy the regulations.

v TABLE OF CONTENTS ABSTRACT............................................................................................................................... iii LIST OF FIGURES................................................................................................................... vii LIST OF TABLES...................................................................................................................... ix ABBREVIATIONS AND ACRONYMS....................................................................................... xi 1 INTRODUCTION.................................................................................................................1-1 2 REGULATORY REQUIREMENTS......................................................................................2-1 3 PREVIOUS GUIDANCE......................................................................................................3-1 4 TECHNICAL EVALUATION................................................................................................4-1 4.1 Empirical Database......................................................................................................4-1 4.2 Fuel Cladding Failure Mechanisms............................................................................. 4-11 4.2.1 High Temperature Cladding Threshold......................................................... 4-11 4.2.2 High-Temperature Cladding Failure Threshold Technical Basis................... 4-12 4.2.3 Applicability of High-Temperature Failure Threshold..................................... 4-18 4.2.4 Analytical Considerations for High-Temperature Failure Threshold.............. 4-19 4.3 Hydrogen-Enhanced PCMI Cladding Failure Threshold............................................. 4-20 4.3.1 Hydride Formation and Orientation............................................................... 4-20 4.3.2 Impact of Fuel Burnup.................................................................................. 4-24 4.3.3 Impact of Pulse Width................................................................................... 4-25 4.3.4 Cladding Temperature Effects...................................................................... 4-26 4.3.5 Mixed Oxide Fuel.......................................................................................... 4-27 4.3.6 RG 1.236 PCMI Cladding Failure Threshold................................................. 4-28 4.3.7 Hydrogen Enhanced PCMI Cladding Failure Thresholds Technical Bases........................................................................................................... 4-30 4.3.8 Applicability of PCMI Failure Thresholds....................................................... 4-38 4.3.9 Analytical Considerations............................................................................. 4-40 4.4 Molten Fuel Failure Threshold.................................................................................... 4-40 4.5 Allowable Limits on Damaged Core Coolability........................................................... 4-40 4.5.1 Regulatory Guide 1.236 Core Coolability Criteria.......................................... 4-40 4.5.2 Molten Fuel and Core Coolability Technical Basis........................................ 4-41 4.6 Allowable Limits on Radiological Consequences........................................................ 4-44 4.7 Allowable Limits on Reactor Coolant System Pressure.............................................. 4-44 4.8 Transient Fission Gas Release................................................................................... 4-44 5 CONCLUSIONS..................................................................................................................5-1 6 REFERENCES....................................................................................................................6-1

vii LIST OF FIGURES Figure 1-1 Fuel Response to a Rapid Reactivity Insertion.......................................................1-2 Figure 2-1 INEEL Safety Poster..............................................................................................2-1 Figure 3-1 PWR PCMI Cladding Failure Threshold.................................................................3-5 Figure 3-2 BWR PCMI Cladding Failure Threshold.................................................................3-5 Figure 4-1 Reported Peak and Failed Fuel Enthalpy versus Burnup.......................................4-3 Figure 4-2 Reported Peak and Failed Fuel Enthalpy versus Pulse Width................................4-4 Figure 4-3 Reported Peak and Failed Fuel Enthalpy versus Oxide/Wall Ratio........................4-5 Figure 4-4 Reported Peak and Failed Fuel Enthalpy Rise versus Burnup...............................4-6 Figure 4-5 Reported Peak and Failed Fuel Enthalpy Rise versus Cladding Hydrogen*...........4-7 Figure 4-6 Transient Fission Gas Release versus Burnup......................................................4-8 Figure 4-7 Transient Fission Gas Release versus Pulse Width...............................................4-9 Figure 4-8 Transient Fission Gas Release versus Total Fuel Enthalpy Rise......................... 4-10 Figure 4-9 High-Temperature Cladding Failure Threshold.................................................... 4-12 Figure 4-10 High Burnup Ballooning Failures........................................................................ 4-13 Figure 4-11 Non-PCMI Empirical DatabasePeak Fuel Enthalpy versus Burnup................. 4-13 Figure 4-12 Non-PCMI Empirical DatabasePeak Fuel Enthalpy versus Differential Pressure............................................................................................................ 4-14 Figure 4-13 High-Temperature Cladding Failure Threshold in Context.................................. 4-17 Figure 4-14 Hydride Orientation and Distribution in Various High Burnup Claddings............. 4-21 Figure 4-15 Comparison of EPRIs CSED(TE) correlations for Zircaloy-4 and Zircaloy-2 at CZP............................................................................................................... 4-22 Figure 4-16 NSRR RXA vs. SRA fuel enthalpy rise............................................................... 4-23 Figure 4-17 Comparison of EPRIs CSED(TE) and PNNLs Temperature Scaling Factors.............................................................................................................. 4-27 Figure 4-18 PCMI Cladding Failure ThresholdRXA Cladding at or Above 260°C (500°F).............................................................................................................. 4-28 Figure 4-19 PCMI Cladding Failure ThresholdSRA Cladding at or Above 260°C (500°F).............................................................................................................. 4-29 Figure 4-20 PCMI Cladding Failure ThresholdRXA Cladding Below 260°C (500°F).......... 4-29 Figure 4-21 PCMI cladding Failure ThresholdSRA Cladding Below 260°C (500°F)........... 4-30 Figure 4-22 Comparison of SRA Cladding HZP PCMI Failure Thresholds............................ 4-32 Figure 4-23 Comparison of SRA Cladding CZP PCMI Failure Thresholds............................ 4-32 Figure 4-24 Comparison of RXA Cladding HZP PCMI Failure Thresholds............................ 4-33 Figure 4-25 Comparison of RXA Cladding CZP PCMI Failure Thresholds............................ 4-33

viii Figure 4-26 Modified Exponential Function Curve Fit for RXA cladding at CZP..................... 4-35 Figure 4-27 EPRI CSED for Zircaloy-4 (SRA)....................................................................... 4-36 Figure 4-28 Data Adjustment for RXA Cladding with a Liner................................................. 4-38 Figure 4-29 Reported Fuel Dispersal During Prompt Power Excursions................................ 4-44 Figure 4-30 Transient Fission Gas Release Correlations...................................................... 4-46

ix LIST OF TABLES Table 4-1 Extent of Empirical Database..................................................................................4-2 Table 4-2 Data Set Results................................................................................................... 4-15 Table 4-3 Pulse Width Variability........................................................................................... 4-25

xi ABBREVIATIONS AND ACRONYMS ASME American Society of Mechanical Engineers AOO anticipated operational occurrence BIGR Fast Impulse Graphite Reactor (Russia)

BOC beginning of cycle BWR boiling-water reactor C

Celsius CABRI research reactor located in Cadarache, France cal/g calories per gram CFR Code of Federal Regulations CPR critical power ratio CRD control rod drop CRE control rod ejection CSED critical strain energy density CZP cold zero power DBA design-basis accident DG draft regulatory guide DNB departure from nucleate boiling DNBR departure from nucleate boiling ratio E110 cladding material used in VVER fuel rods (Zr-1.0Nb by wt%)

EPRI Electric Power Research Institute F

Fahrenheit FCI fuel-coolant interaction FGR fission gas release GDC general design criterion/criteria GNF Global Nuclear Fuel GWd gigawatt-day(s)

HZP hot zero power I

iodine IGR Impulse Graphite Ractor (Kazakhstan)

INEEL Idaho National Engineering and Environmental Laboratory JAEA Japan Atomic Energy Agency Kr krypton LOCA loss-of-coolant accident LSQ least squares LWR light-water reactor M5 cladding trademark of Framatome ANP (Zr-1.0Nb-0.13O by wt%)

MDA Mitsubishi Developed Alloy (Zr-0.8Sn-0.5Nb-0.32Fe-0.1Cr by wt%)

mm millimeter(s)

xii MOX mixed oxide (UO2/PuO2)

MPa megapascal(s) ms millisecond(s)

MTU metric ton(s) of uranium NDA New Developed Alloy (Zr-1.0Sn-0.27Fe-0.16Cr-0.1Nb-0.01Ni by wt%)

NEA Nuclear Energy Agency NRC U.S. Nuclear Regulatory Commission NSRR Nuclear Safety Research Reactor (Japan)

O.D.

outer diameter OECD Organisation for Economic Co-operation and Development PBF Power Burst Facility PCMI pellet-to-cladding mechanical interaction PNNL Pacific Northwest National Laboratory PuO2 plutonium dioxide PWR pressurized-water reactor RCS reactor coolant system RG regulatory guide RIA reacivity initiated accident RPS reactor protection system RTR research and test reactor RXA recrystallization annealed SBLOCA small-break loss-of-coolant accident SED strain energy density SPERT Special Power Excursion Test Reactor SRA stress relieved annealed SRP Standard Review Plan (NUREG-0800)

Sv sievert(s)

TE total elongation TEDE total effective dose equivalent TREAT Transient Reactor Test Facility UE uniform elongation UFSAR updated safety analysis report m

micrometer(s)

UO2 uranium dioxide VVER Russian-type pressurized-water reactor wppm weight parts per million wt%

weight percent Xe xenon ZIRLO cladding trademark of Westinghouse (Zr-1.0Nb-1.0Sn-0.1Fe by wt%)

1-1 1 INTRODUCTION The purpose of this report is to document the technical and regulatory bases for the acceptance criteria and guidance presented in Regulatory Guide (RG) 1.236, Pressurized-Water Reactor Control Rod Ejection and Boiling-Water Reactor Control Rod Drop Accidents, Revision 0, issued June 2020 (Ref. 1). The pressurized-water reactor (PWR) control rod ejection (CRE) and boiling-water reactor (BWR) control rod0F1 drop (CRD) accidents are low-probability events involving a sudden and rapid insertion of positive reactivity. These postulated accidents are safety significant because of their potential ability to cause significant core damage and challenge the integrity of the reactor coolant pressure boundary.

Many anticipated operational occurrences (AOOs) and postulated accidents involving an unintended positive reactivity insertion and subsequent reactor power excursion are analyzed in the licensees updated final safety analysis report (UFSAR). Unintended positive reactivity insertion may result from many abnormal operational situations, including an increase in reactor coolant system (RCS) pressure (e.g., BWR turbine trip), decrease in moderator (i.e., reactor coolant) temperature (e.g., PWR inadvertent opening of an atmospheric dump valve), control rod misoperation (e.g., BWR rod withdrawal error, PWR inadvertent control bank withdrawal),

and secondary piping failure (e.g., PWR main steamline break). For each of these AOO and postulated accident scenarios, the UFSAR documents the accident progression; response of safety-related systems, structures, and components; and compliance with applicable regulatory requirements.

More than the above reactivity insertion scenarios, the uncontrolled movement of a single control rod out of the core has the highest potential ability to cause significant core damage and challenge the integrity of the reactor coolant pressure boundary. As described below, the postulated PWR CRE and BWR CRD accidents are the most limiting events involving an uncontrolled movement of a single control rod out of the core. The following examples of event descriptions and sequence of events were derived from the UFSARs of existing plants:

PWR CRE:

The rod ejection accident is caused by a failure of a control rod drive mechanism housing, which allows a control rod to be rapidly ejected from the reactor by the reactor coolant system (RCS) pressure. The control rod is ejected in 0.15 seconds from the fully inserted position. A power excursion will result, and if the reactivity worth of the ejected control rod is large enough, the reactor will become prompt critical. The resulting power excursion will be limited by the fuel temperature feedback and the accident will be terminated when the reactor protection system (RPS) trips the reactor on high neutron flux or high RCS pressure. RCS pressure increases due to the core power excursion, and pressurizer spray, the pressurizer power operated relief valve, and the pressurizer code safety valves will respond to mitigate the pressure increase. If a rod ejection were to occur, the nuclear design of the reactor and limits on control rod insertion will limit any potential fuel damage to acceptable levels. Cladding failure can result from the core power excursion and the highly peaked core power distribution near the ejected rod location. The failure of the control rod drive mechanism housing also constitutes a 1.50-inch-diameter small-break loss-of-coolant accident (SBLOCA). The emergency 1

Commercial BWRs are designed with cruciform control blades that travel between four adjacent fuel bundles. These blade configurations are referred to as control rods.

1-2 core cooling system will actuate on low RCS pressure or high reactor building pressure and will maintain core cooling. This type of SBLOCA is bounded by the limiting SBLOCA analyses.

BWR CRD:

The accidents that result in releases of radioactive material from the fuel with the nuclear system process barrier, primary containment, and secondary containment initially intact are the results of various failures of the control rod drive system. Examples of such failures are collet finger failures in one control rod drive mechanism, a control drive system pressure regulator malfunction, and a control rod drive mechanism ball check valve failure. None of the single failures associated with the control rods or the control rod system results in a greater release of radioactive material from the fuel than the release that results when a single control rod drops out of the core after being disconnected from its drive and after the drive has been retracted to the fully withdrawn position. Thus, this control rod drop accident is established as the design-basis accident for the category of accidents resulting in radioactive material release from the fuel with all other barriers initially intact. A highly improbable combination of actual events would be required for the design basis control rod drop accident to occur. The actual events required are (1) failure of the rod-to-drive coupling, (2) sticking of the control rod in its fully inserted position as the drive is withdrawn, (3) full withdrawal of the control rod drive, (4) failure of the operator to notice the lack of response of neutron monitoring channels as the rod drive is withdrawn, and (5) failure of the operator to verify rod coupling.

Figure 1-1 illustrates the rapid increase in local core power following an uncontrolled withdrawal of a single control rod. Fuel temperature rapidly increases, prompting fuel pellet thermal expansion and pellet-to-cladding mechanical interaction (PCMI). An increase in cladding temperature follows. The reactivity excursion is initially mitigated by Doppler feedback and delayed neutron effects followed by a reactor trip.

Figure 1-1 Fuel Response to a Rapid Reactivity Insertion

2-1 2 REGULATORY REQUIREMENTS For stability and controllability, commercial reactor cores are designed to achieve and maintain criticality (i.e., sustain a chain reaction) on delayed neutrons. A vast majority (greater than 99 percent) of neutrons emitted from nuclear fission appear instantaneously and are referred to as prompt neutrons. The remaining delayed neutrons are emitted after a further decay step.

Designing the reactor core to be subcritical on prompt neutrons alone delays the rate-of-change of the reactor period and ensures that the reactor is more stable and controlled. The large insertion of positive reactivity associated with certain postulated PWR CRE and BWR CRD accident scenarios may produce a prompt critical condition in the reactor core. Under these conditions, the reactor core becomes completely unstable and uncontrollable. Hence, PWR CRE and BWR CRD are design-basis accidents (DBAs) with respect to reactor core design.1F2 On January 3, 1961, the U.S. Armys prototype modular reactor, Stationary Low Power Reactor (SL-1), experienced a prompt critical power excursion leading to a steam explosion. An improper central control rod withdrawal sequence initiated the violent explosion, which killed all three operators and released significant amounts of radiation. Figure 2-1 depicts an Idaho National Engineering and Environmental Laboratory (INEEL) safety poster showing the extent of damage to the reactor pressure boundary following the prompt critical power excursion and steam explosion.

Figure 2-1 INEEL Safety Poster (INEEL 81-3966) 2 Reactor core design refers to core configuration (i.e., fuel rod, assembly lattice, and core dimensions),

control rod design (e.g., dimensions, boron carbide loading), and fuel management (e.g., uranium-235 enrichment, reload loading pattern).

2-2 This nuclear accident was likely on the minds of the Atomic Energy Commission when it drafted reactivity-related general design criteria (GDC) in Appendix A, General Design Criteria for Nuclear Power Plants, to Title 10 of the Code of Federal Regulations (10 CFR) Part 50, Domestic Licensing of Production and Utilization Facilities (Ref. 2)

GDC 11, Reactor inherent protection, requires reactor core and associated coolant systems to be designed so that, in the power operating range, the net effect of the prompt inherent nuclear feedback characteristics tends to compensate for a rapid increase in reactivity. For certain postulated PWR CRE and BWR CRD accident scenarios, the rapid, prompt critical, rate of change in local fission power is beyond the response time of safety-related reactor protection systems. Hence, the reactor core must be designed to self-limit the local core power excursion.

As fuel pellets heat, the density of fissile material (e.g., uranium-235) decreases, introducing negative reactivity, which tends to limit the power excursion. Note that there may not be sufficient time for the heat flux to reach the moderator (i.e., reactor coolant) to credit negative reactivity coefficients associated with moderator density to limit core damage. Therefore, the reactor core must be designed to exhibit a negative fuel temperature coefficient sufficient to provide this inherent protection feature during postulated PWR CRE and BWR CRD accidents.

Hence, PWR CRE and BWR CRD are DBAs with respect to reactor core design and compliance with GDC 11.

GDC 28, Reactivity limits, requires the reactivity control system to be designed with appropriate limits on the potential amount and rate of reactivity increase to provide assurance that the effects of postulated reactivity accidents can do neither of the following:

(1) result in damage to the reactor coolant pressure boundary greater than limited local yielding nor (2) sufficiently disturb the core, its support structures or other reactor pressure vessel internals to impair significantly the capability to cool the core.

These postulated reactivity accidents shall include consideration of CRE, CRD, steamline rupture, changes in reactor coolant temperature and pressure, and cold-water addition. The degree of core damage must be quantified and shown to satisfy the above requirements.

Because the postulated PWR CRE and BWR CRD accidents are classified as low-probability Condition IV events, limited core damage (i.e., fuel rod failure) is allowed, provided offsite and control room radiological consequences remain within acceptable limits. The NRC established acceptable radiation dose limits in 10 CFR 100.11, Determination of exclusion area, low population zone, and population center distance (Ref. 3) and 10 CFR 50.67, Accident source term. For completeness, the regulatory limits from 10 CFR 50.67 follow:

(2) The NRC may issue the amendment only if the applicant's analysis demonstrates with reasonable assurance that:

(i) An individual located at any point on the boundary of the exclusion area for any 2-hour period following the onset of the postulated fission product release, would not receive a radiation dose in excess of 0.25 Sv [sievert] (25 rem) total effective dose equivalent (TEDE).

2-3 (ii) An individual located at any point on the outer boundary of the low population zone, who is exposed to the radioactive cloud resulting from the postulated fission product release (during the entire period of its passage), would not receive a radiation dose in excess of 0.25 Sv (25 rem) total effective dose equivalent (TEDE).

(iii) Adequate radiation protection is provided to permit access to and occupancy of the control room under accident conditions without personnel receiving radiation exposures in excess of 0.05 Sv (5 rem) total effective dose equivalent (TEDE) for the duration of the accident.

Footnote 2 of this regulation states that the 0.25 Sv (25 rem) TEDE limit is not intended as an acceptable limit for public exposure under DBA conditions but rather for the evaluation of proposed design-basis changes with respect to potential reactor accidents of exceedingly low probability of occurrence and low risk of public exposure to radiation. Based on this position, the allowable consequences for the PWR CRE and BWR CRD accidents have been limited to well within acceptance criteria. In practice, this has translated to 0.0625 Sv (6.25 rem) TEDE for members of the public. The allowable radiation exposures for workers have been maintained at 0.05 Sv (5 rem) TEDE. RG 1.183, Alternative Radiological Source Terms for Evaluating Design Basis Accidents at Nuclear Power Reactors (Ref. 4), and RG 1.195, Methods and Assumptions for Evaluating Radiological Consequences of Design Basis Accidents at Light-Water Nuclear Power Reactors (Ref. 5), provide guidance with respect to predicting radiological consequences for PWR CRE and BWR CRD accidents.

Based on the above, core damage must be limited and quantified to demonstrate compliance with (1) the GDC 28 requirement on reactor coolant pressure boundary local yielding, (2) the GDC 28 requirement to maintain a core geometry amenable to continued cooling, and (3) the limits in 10 CFR 100.11, 10 CFR 50.67, or both on radiological consequences. RG 1.236 defines analytical limits on maximum radial average fuel enthalpy and local fuel melting for demonstrating compliance with items i and ii above and define fuel rod cladding failure thresholds for demonstrating compliance with item iii.

3-1 3 PREVIOUS GUIDANCE In 1974, the NRC issued RG 1.77, Assumptions Used for Evaluating a Control Rod Ejection Accident for Pressurized Water Reactors (Ref. 6). This guidance identified acceptable analytical methods and assumptions as well as acceptance criteria to demonstrate compliance with GDC 28 and applicable dose requirements for the postulated PWR CRE accident. Key guidance included the following:

1.

A sufficient number of initial reactor states to completely bracket all possible operational conditions of interest should be analyzed to assure examination of upper bounds on ultimate damage.

2.

In areas of uncertainty, the appropriate minimum or maximum parameters relative to nominal or expected values should be used to assure a conservative evaluation.

3.

A calculated radial average energy density of 280 cal/g [calories per gram] at any axial fuel location in any fuel rod provides a conservative maximum limit to ensure that core damage will be minimal and that both short-term and long-term core cooling capability will not be impaired.

4.

Maximum reactor pressure limited to the value that will cause stresses to exceed the Emergency Condition (Service Level C) as defined in the ASME [American Society of Mechanical Engineers] Boiler and Pressure Vessel code (Ref. 7)

a.

No credit should be taken for the possible pressure reduction caused by the assumed failure of the control rod pressure housing.

5.

Offsite dose consequences limited to well within the guidelines in 10 CFR Part 100.

6.

The number of fuel rods experiencing clad failure should be calculated and used to obtain the amount of contained fission product inventory released to the reactor coolant system.

a.

It should be assumed that clad failure occurs if the heat flux equals or exceeds the value corresponding to the onset of the transition from nucleate to film boiling (DNB [departure from nucleate boiling]), or for other appropriate causes.

b.

A minimum DNBR [DNB ratio] should be determined from the evaluation of the experimental data to ensure a 95% probability with a 95% confidence level that DNB has not occurred for the fuel element being evaluated.

c.

Other DNB or clad failure correlations may be used if they are adequately justified by analytical methods and supported by sufficient experimental data.

3-2 In 1981, the NRC issued NUREG-0800, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition (SRP), Revision 1. Section 4.2, Fuel System Design, Revision 2, issued July 1981 (Ref. 8), provided the following updated guidance:

1.

Violent Expulsion of Fuel: In severe PWR CRE and BWR CRD accident scenarios, the large and rapid deposition of energy in the fuel can result in melting, fragmentation, and dispersal of fuel. The mechanical action associated with fuel dispersal can be sufficient to destroy the cladding and the rod-bundle geometry of the fuel and to produce pressure pulses in the primary system. To meet the guidelines of Regulatory Guide 1.77 as it relates to preventing widespread fragmentation and dispersal of the fuel and avoiding the generation of pressure pulses in the primary system of a PWR, a radially averaged enthalpy limit of 280 cal/g should be observed. This 280 cal/g limit should also be used for BWRs.

2.

For postulated accidents, the total number of fuel rods that exceed the criteria has been assumed to fail for radiological dose calculation purposes.

3.

For postulated accidents, the total number of rods that experience centerline melting should be assumed to fail for radiological dose calculation purposes.

4.

For severe BWR CRD scenarios at zero or low power, fuel failure is assumed to occur if the radially averaged fuel rod enthalpy is greater than 170 cal/g at any axial location. For full-power BWR CRD accidents and all PWR CRE accidents, the thermal margin criteria (DNBR and CPR [critical power ratio]) are used as fuel failure criteria. The 170 cal/g enthalpy criterion is primarily intended to address cladding overheating effects, but it also indirectly addresses pellet-to-cladding mechanical interactions (PCMI). Other criteria may be more appropriate, but continued approval of this enthalpy criterion and the thermal margin criteria may be given until generic studies yield improvements.

In March 2007, the NRC issued Revision 3 to SRP Section 4.2, Fuel System Design (Ref. 9).

Appendix B, Interim Acceptance Criteria and Guidance for the Reactivity Initiated Accidents, provided interim guidance and acceptance criteria for the postulated PWR CRE and BWR CRD accidents. By 2007, the state of knowledge of fuel rod performance under prompt power excursion conditions had increased significantly. This knowledge prompted the need for new guidance to build on the enhanced database drawn from operating experience and controlled experiments. The empirical database had expanded from the earlier Special Power Excursion Test Reactor (SPERT) and Transient Reactor Test Facility (TREAT) research programs (which formed the basis of the initial RG 1.77 analytical limits) to include test results from the Power Burst Facility (PBF), as well as significant contributions from international research programs at the CABRI research reactor (France), Nuclear Safety Research Reactor (NSRR) (Japan),

Impulse Graphite Reactor (IGR) (Russian Federation), and Fast Pulse Graphite Reactor (BIGR)

(Russian Federation).

With respect to core damage and compliance with GDC 28, the NRC provided the following updated acceptance criteria and guidance:

3-3

1.

Peak radial average fuel enthalpy must remain below 230 cal/g.

2.

Peak fuel temperature must remain below incipient fuel melting conditions.

3.

Mechanical energy generated as a result of (1) non-molten fuel-to-coolant interaction and (2) fuel rod burst must be addressed with respect to reactor pressure boundary, reactor internals, and fuel assembly structural integrity.

4.

No loss of coolable geometry due to (1) fuel pellet and cladding fragmentation and dispersal and (2) fuel rod ballooning.

With respect to estimating the number of fuel rods with failed cladding, the NRC provided the following updated acceptance criteria and guidance:

1.

The high cladding temperature failure criteria for zero power conditions is a peak radial average fuel enthalpy greater than 170 cal/g for fuel rods with an internal rod pressure at or below system pressure and 150 cal/g for fuel rods with an internal rod pressure exceeding system pressure. For intermediate (greater than 5% rated thermal power) and full power conditions, fuel cladding failure is presumed if local heat flux exceeds thermal design limits (e.g., DNBR and CPR).

2.

The PCMI failure criteria is a change in radial average fuel enthalpy greater than the corrosion-dependent limit depicted in Figure 3-1 (PWR) and Figure 3-2 (BWR).

3.

Fuel cladding failure may occur almost instantaneously during the prompt fuel enthalpy rise (due to PCMI) or may occur as total fuel enthalpy (prompt + delayed), heat flux, and cladding temperature increase. For the purpose of calculating fuel enthalpy for assessing PCMI failures, the prompt fuel enthalpy rise is defined as the radial average fuel enthalpy rise at the time corresponding to one pulse width after the peak of the prompt pulse. For assessing high cladding temperature failures, the total radial average fuel enthalpy (prompt + delayed) should be used.

With respect to estimating the fission product release from each failed fuel rod, the NRC provided new guidance:

1.

The total fission-product gap fraction available for release would include the steady-state gap inventory (present prior to the event) plus any fission gas released during the event. The steady-state gap inventory would be consistent with the Non-LOCA [loss-of-coolant accident] gap fractions cited in RG 1.183 Revision 1 (Tables 3 and 4) and RG 1.195 (Table 2) and would be dependent on operating power history. Whereas fission gas release (into the rod plenum) during normal operation is governed by diffusion, pellet fracturing and grain boundary separation are the primary mechanisms for fission gas release during the transient.

3-4

2.

Based upon measured fission gas release (FGR) from several prompt critical test programs, the staff developed the following correlation between gas release and maximum fuel enthalpy increase:

Transient FGR = [(0.2286*H) - 7.1419]

Where:

FGR = Fission gas release, % (must be > 0)

H = Increase in fuel enthalpy, cal/g

3.

The transient release from each axial node which experiences the power pulse may be calculated separately and combined to yield the total transient FGR for a particular fuel rod. The combined steady-state gap inventory and transient FGR from every fuel rod predicted to experience cladding failure (all failure mechanisms) should be used in the dose assessment. Additional guidance is available within RG 1.183 and 1.195.

The NRC provided the bases of Section 4.2, Appendix B, in the memorandum, Technical and Regulatory Basis for the Reactivity-Initiated Accident Interim Acceptance Criteria and Guidance, dated January 19, 2007 (Ref. 10). The NRC issued the revised guidance as interim because, at that time, it recognized that several important test programs were being developed and planned that would help resolve widely debated issues such as scaling room temperature test results to higher temperature operating conditions. Anticipated test programs included the CABRI conversion (sodium-to-water loop) and NSRR hot capsule program.

In 2015, the staff evaluated newly published empirical data and analyses and identified further changes to guidance in the NRC memorandum, Technical and Regulatory Basis for the Reactivity-Initiated Accident Acceptance Criteria and Guidance, Revision 1, dated March 16, 2015 (Ref. 11).

In 2016, the NRC issued the revised guidance in the Federal Register (FR) for public comment as Draft Regulatory Guide (DG)-1327, Pressurized Water Reactor Control Rod Ejection and Boiling Water Reactor Control Rod Drop Accidents (Ref. 12). The NRC received a total of 124 comments from 12 stakeholders. It accepted over 100 comments and made subsequent changes to the draft guidance.

In 2019, the NRC published a revised version of DG-1327 in the Federal Register for a second round of public comments (Ref. 13). The NRC received a total of 54 comments from 7 stakeholders. It accepted over 30 comments and made subsequent changes to the draft guidance.

In 2020, the NRC issued RG 1.236, Pressurized Water Reactor Control Rod Ejection and Boiling Water Reactor Control Rod Drop Accidents. RG 1.236 contains state-of-the-art guidance that replaces the legacy guidance found in RG 1.77 and SRP Section 4.2, Appendix B. The NRC staff believed that RG 1.77 was no longer applicable or acceptable for contemporary analytical methods and fuel designs. In addition, RG 1.77 was only applicable to PWR CRE, whereas RG 1.236 addresses both PWR CRE and BWR CRD accidents. For those reasons, the NRC withdrew RG 1.77 concurrently with the issuance of RG 1.236. Although withdrawn, current licensees with RG 1.77 in their licensing basis may continue to use it, and its withdrawal does not affect any existing licenses or agreements.

3-5 Figure 3-1 PWR PCMI Cladding Failure Threshold Figure 3-2 BWR PCMI Cladding Failure Threshold

4-1 4 TECHNICAL EVALUATION This section details the empirical database of fuel performance under rapid power excursion testing and documents the technical bases of the RG 1.236 analytical limits and guidance that satisfy the regulatory requirements identified in Section 2.

4.1 Empirical Database Nuclear Energy Agency (NEA), Organisation for Economic Co-Operation and Development (OECD), State-of-the-Art Report, Nuclear Fuel Behaviour under Reactivity-initiated Accident (RIA) Conditions, issued 2010 (Ref. 14) describes the experimental test facilities and fuel rod test specimens, tabulation of experimental results, and key observations and findings, as well as the analysis of fuel performance under rapid power excursion test conditions. This report was updated in 2022 in Ref. 15. There were updates to the database in the 2022 report that demonstrate that the criteria detailed in RG 1.236 continue to be applicable. The 2022 data was used in the figures throughout this document. Since the issuance of RG 1.77 in 1974, the empirical database has grown significantly, not only in the sheer number of tests but also in the variety of fuel rod designs and materials tested, the extent of prior irradiation history of the test specimens (e.g., burnup, corrosion), the variety of test conditions (e.g., pulse width, temperature), and the quality and variety of experimental data reported. Table 4-1 lists the range of fuel rod test specimens and test conditions in the empirical database.

Portions of the empirical database used to develop earlier guidance have been subsequently reassessed and altered. These reported alterations include more detailed characterization of test specimens (e.g., cladding hydrogen content) as well as changes in test results (e.g., fuel enthalpy, transient fission gas release). The following figures depict the compiled, up-to-date empirical database employed to derive the damaged core coolability acceptance criteria and fuel rod cladding failure thresholds:

Figure 4-1 Reported peak and failed fuel enthalpy versus burnup Figure 4-2 Reported peak and failed fuel enthalpy versus pulse width Figure 4-3 Reported peak and failed fuel enthalpy versus oxide/wall ratio Figure 4-4 Reported peak and failed fuel enthalpy rise versus burnup Figure 4-5 Reported peak and failed fuel enthalpy rise versus cladding hydrogen In addition to fuel damage, measurements of fission gas released during the prompt-power testing have been reported for many of the experiments. The following figures depict the compiled, up-to-date empirical database employed to derive the transient FGR correlations:

Figure 4-6 Transient fission gas release versus burnup Figure 4-7 Transient fission gas release versus pulse width Figure 4-8 Transient fission gas release versus total fuel enthalpy rise

4-2 Table 4-1 Extent of Empirical Database Parameter Extent of Empirical Database Test Specimen Minimum Maximum Reactor Fuel Design PWR, BWR, VVER, RTR*

Cladding Alloy Zry-2 (liner, no liner), Zry-4, ZIRLO, Optimized ZIRLO, M5, E110, MDA Cladding O.D. (mm) 7.92 14.5 Cladding Thickness (µm) 495 915 Fill Pressure (MPa) 0.1 4.6 Cladding Oxide Thickness (µm) 0 110 Cladding Hydrogen Content (wppm) 0 800**

Segment Burnup (GWd/MTU) 0 80 Test Condition Minimum Maximum Pulse Width (ms) 2.5 950 Deposited Energy (cal/g) 51 695 Peak Fuel Enthalpy (cal/g) 37 350 Fuel Enthalpy Rise (cal/g) 14 335

• Fuel rods irradiated in research and test reactors

• • For several test specimens, including this maximum value, cladding hydrogen content was estimated based on alloy-specific hydrogen uptake models and measured cladding oxide thickness.

4-3 Figure 4-1 Reported Peak and Failed Fuel Enthalpy versus Burnup 0

50 100 150 200 250 300 0

10 20 30 40 50 60 70 80 Peak Radial Average Fuel Enthalpy (cal/g)

Fuel Burnup (GWd/MTU)

Reported Peak and Failed Fuel Enthalpy Versus Burnup Open Symbol = Non-failed peak fuel enthalpy Closed Symbol = Fuel enthalpy at failure SPERT-CDC PBF BIGR CABRI NSRR IGR

4-4 Figure 4-2 Reported Peak and Failed Fuel Enthalpy versus Pulse Width 0

50 100 150 200 250 300 1

10 100 1000 Peak Radial Average Fuel Enthalpy (cal/g)

Pulse Width (ms)

Reported Peak and Failed Fuel Enthalpy Versus Pulse Width Open Symbol = Non-failed peak fuel enthalpy Closed Symbol = Fuel enthalpy at failure SPERT-CDC PBF BIGR CABRI NSRR IGR

4-5 Figure 4-3 Reported Peak and Failed Fuel Enthalpy versus Oxide/Wall Ratio

4-6 Figure 4-4 Reported Peak and Failed Fuel Enthalpy Rise versus Burnup

4-7 Figure 4-5 Reported Peak and Failed Fuel Enthalpy Rise versus Cladding Hydrogen*

• For a portion of the data, cladding hydrogen content is calculated using alloy-specific corrosion and hydrogen uptake models based on measured oxide thickness and reported fuel burnup. An average value was used if a range of oxide thickness was reported. Several recent NSRR and CABRI tests reported cladding hydrogen content. Similar to oxide thickness, an average value was used if a range was presented.

4-8 Figure 4-6 Transient Fission Gas Release versus Burnup 0

5 10 15 20 25 30 35 40 0.0 10.0 20.0 30.0 40.0 50.0 60.0 70.0 80.0 Measured Transient FGR (%)

Fuel Burnup (GWd/MTU)

CABRI NSRR PWR NSRR BWR BIGR VVER IGR VVER CABRI MOX NSRR MOX Test Segment Exposure:

Purple: > 70 GWd/MTU Orange:60-69.9 GWd/MTU Red:

50 - 59.9 GWd/MTU Blue:

30 - 49.9 GWd/MTU Green: 10 - 29.9 GWd/MTU Open Symbol = Non-failed Closed Symbol = Failed

4-9 Figure 4-7 Transient Fission Gas Release versus Pulse Width 0

5 10 15 20 25 30 35 40 1

10 100 1000 Measured Transient FGR (%)

Pulse Width (ms)

CABRI NSRR PWR NSRR BWR BIGR VVER IGR VVER CABRI MOX NSRR MOX Test Segment Exposure:

Purple: > 70 GWd/MTU Orange:60-69.9 GWd/MTU Red:

50 - 59.9 GWd/MTU Blue:

30 - 49.9 GWd/MTU Green: 10 - 29.9 GWd/MTU Open Symbol = Non-failed Closed Symbol = Failed

4-10 Figure 4-8 Transient Fission Gas Release versus Total Fuel Enthalpy Rise 0

5 10 15 20 25 30 35 40 0

50 100 150 200 Measured Transient FGR (%)

Peak Enthalpy Increase (cal/g)

CABRI NSRR PWR NSRR BWR BIGR VVER IGR VVER CABRI MOX NSRR MOX Test Segment Exposure:

Purple: > 70 GWd/MTU Orange:60-69.9 GWd/MTU Red:

50 - 59.9 GWd/MTU Blue:

30 - 49.9 GWd/MTU Green: 10 - 29.9 GWd/MTU Open Symbol = Non-failed Closed Symbol = Failed

4-11 4.2 Fuel Cladding Failure Mechanisms To ensure that regulatory criteria associated with offsite and onsite radiological consequences are satisfied, the number of fuel rod failures must not be underestimated. Fuel cladding failure mechanisms associated with postulated CRE and CRD accidents include the following:

(1) brittle failure: high-temperature oxygen-induced embrittlement and fragmentation (2) ductile failure: high-temperature cladding creep (rod ballooning and burst)

(3)

PCMI: hydrogen-enhanced PCMI cladding failure (4) fuel melt: molten fuel-induced swelling PCMI cladding failure MacDonald et al. (1980) (Ref. 16) concluded that the mode of fuel rod failure is strongly dependent on previous irradiation history. Irradiation history would include power history, burnup, and inservice cladding corrosion (oxidation). Other important factors contributing to fuel rod failure include (1) the initial conditions of the fuel rod (e.g., initial fuel enthalpy, fuel-to-clad gap, rod internal pressure), (2) the initial conditions of the reactor coolant (e.g., temperature, pressure, mass flow), and (3) fuel design. Of course, the governing influence on the fuel rods response to the postulated transient is the amount and rate of reactivity insertion. The influence of each of these factors differs for each failure mechanism.

PCMI cladding failure is predicted to occur relatively early in the event before any significant increase in cladding temperature due to the prompt thermal expansion of the fuel pellet, which can be exacerbated at high burnups by gaseous fission product swelling, whereas failure mechanisms (1) and (2) involve high cladding temperature phenomena (e.g., reduced yield strength, oxygen diffusion into base metal) and will experience failure later in the event progression.

The following sections describe the technical bases for the fuel cladding failure thresholds in the guidance.

4.2.1 High Temperature Cladding Threshold The NEA RIA report (Ref. 15) shows the empirically based high-temperature cladding failure threshold. This composite failure threshold encompasses both the brittle and ductile failure mechanisms ((1) and (2) of Section 4.2, respectively) and should be applied for events with prompt critical excursions (i.e., ejected rod worth or drop rod worth greater than or equal to

$1.00). Because ductile failure depends on cladding temperature and differential pressure (i.e., rod internal pressure minus reactor pressure), the composite failure threshold is expressed in peak radial average fuel enthalpy (cal/g) versus fuel cladding differential pressure (megapascals (MPa)).

4-12 Figure 4-9 High-Temperature Cladding Failure Threshold For prompt critical scenarios that experience a prolonged power level following the prompt pulse, fuel cladding failure is presumed if local heat flux exceeds thermal design limits (e.g., DNB and CPRs).

For nonprompt critical excursions, fuel cladding failure is presumed if local heat flux exceeds thermal design limits.

4.2.2 High-Temperature Cladding Failure Threshold Technical Basis The high-temperature cladding failure threshold presented in RG 1.236 remains unchanged from that proposed in the 2015 NRC RIA memorandum (Ref. 11). As a result, much of this section and Section 4.2.2 has been reproduced from the 2015 memorandum.

Figure 4-10 presents the most recent high-temperature ballooning failure database. Note that this figure is similar to the data discussed in the NEA RIA report (see Figure 54 of Ref. 15) and Electric Power Research Institute (EPRI) Report 1021036, Fuel Reliability Program: Proposed RIA Acceptance Criteria, issued December 2010 (see Figure 3-4 of Ref. 17), but updated to incorporate any modifications to the database. These figures show peak fuel enthalpy from various BIGR and IGR tests on irradiated fuel rod segments and NSRR tests on pressurized, unirradiated fuel rod segments. Figure 4-11 combines this high-temperature ballooning failure database with the entire RIA database and illustrates peak fuel enthalpy as a function of fuel burnup. Figure 4-12 provides the same information but as a function of rod fill pressure. Test results from high burnup and heavily corroded fuel rods that failed due to PCMI have been removed. All CABRI test results were removed due to nonprototypical cladding temperatures experienced in the sodium loop (relative to water). In addition, test results from failed Japanese

4-13 Material Test Reactor rodlets were removed. As described in Section A.5.4 of the NEA RIA report, these rodlets and their results are not representative of light-water reactor (LWR) fuel.

Figure 4-10 High Burnup Ballooning Failures Figure 4-11 Non-PCMI Empirical DatabasePeak Fuel Enthalpy versus Burnup

4-14 Figure 4-12 Non-PCMI Empirical DatabasePeak Fuel Enthalpy versus Differential Pressure Before discussing the above figures and the development of a failure threshold, it is important to emphasize that the data are presented as peak fuel enthalpy. Unlike the PCMI failure plots, the fuel enthalpy at the time of cladding failure (i.e., failure enthalpy) is not reported. Hence, cladding failure may have occurred at a lower enthalpy than the reported peak fuel enthalpy.

Under these circumstances, more data sets for each variable are necessary to establish the boundary between failure and nonfailure.

Examination of Figure 4-11 does not reveal a significant burnup effect on failure threshold.

Based on data from fresh and irradiated fuel, the boundary between failure and nonfailure remains constant at approximately 160 cal/g. A closer look at the data is necessary to confirm the NEAs observation regarding the impact of burnup. Table 4-2 compares select IGR/BIGR data sets. A comparison of RT-8, RT-9, and RT-10 suggests that a transient FGR in these high burnup rods dominated the initial fill pressure. This observation is consistent with the NEA RIA report. Since the fuel enthalpy at failure is unknown, the effect of any difference in transient FGR between RT-8 (60 gigawatt-days per metric ton of uranium (GWd/MTU)) and RT-10 (47 GWd/MTU) is difficult to quantify. A similar comparison of nonfailed specimens, RT-6 and RT-12, confirms the minimal impact of initial pressure on the failure threshold.

4-15 Table 4-2 Data Set Results Test Rod Burnup (GWd/MTU)

Fill Pressure (MPa)

Pulse Width (ms)

Failure (F/NF)

Peak Fuel Enthalpy (cal/g)

RT-8 60 2.0 2.6 F

164 RT-9 60 0.1 2.7 F

165 RT-10 47 2.0 2.6 F

164

- - - -Failure Boundary- - - -

RT-6 48 2.1 2.6 NF 153 RT-12 48 0.1 2.8 NF 155 H1T 49 1.7 750 NF 151 The tight grouping of narrow pulse BIGR and wide pulse IGR test results in Figure 4-12, as well as the comparison of RT-6 and H1T in Table 4-1, suggests that pulse width effects are minimal on failure threshold.

In contrast, Figure 4-12 illustrates a clear trend of decreasing failure threshold with increasing cladding differential pressure. The pressurized, unirradiated NSRR test results suggest a linear relationship between failure enthalpy and cladding differential pressure dropping from 190 cal/g at 0 MPa with a slope of 20 cal/g/MPa. As shown in Figure 4-12, the IGR and BIGR test results on irradiated fuel support this y-intersect and slope. Incorporating a cladding differential pressure dependent failure threshold is consistent with NEA observations, and the slope of this relationship is consistent with an earlier EPRI presentation (see Figure 3.1-2 of Ref. 10). Based upon expected fuel rod conditions, including FGR, rod power history, and restricted axial gas flow, the EPRI report concludes that failure by ballooning and burst in high burnup, overpressured fuel rods is unlikely. The EPRI report also states that low to intermediate burnup fuel rods have internal gas pressure below system pressure, and therefore the driving forces are insufficient to produce ballooning deformations. As a result of these observations, the industry proposed a peak radial average fuel enthalpy failure threshold of 170 cal/g. While these observations have some merit, fuel rod designs and operating conditions continue to evolve, and it is prudent to define failure thresholds based on limiting conditions.

In Section 3.2.1 of EPRI Report 1021036 (Ref. 17), EPRI concludes the following:

A comparison of cladding temperature from RIA-simulation tests with the failure boundary from Figure 3-1 indicates that cladding failure by oxidation-induced embrittlement following an RIA event is unlikely at fuel enthalpy levels below 170 cal/gm.

As described above, the EPRI report documents an industry position with respect to the cladding failure thresholds for (1) oxidation-induced embrittlement and (2) ballooning and burst.

Based upon experimental data, including measured cladding surface temperatures from NSRR experiments with post-DNB operation, the EPRI report concludes the following limiting high-temperature failure threshold:

4-16 At maximum radial average fuel enthalpy levels below 150 cal/g, the cladding temperatures will remain well below the conditions to produce failure by oxidation-induced embrittlement.

It is likely this position was based on the historical practice of truncating PCMI failure for low corrosion rods at an enthalpy rise of 150 calories per gram (cal/g). Within the main body of its report, EPRI concludes that fuel failure from either oxidation-induced embrittlement or ballooning and burst is unlikely below 170 cal/g peak fuel enthalpy.

An examination of the empirical database depicted in Figure 4-12 reveals three failed rods below the red dotted line. BIGR test RT-9 consisted of a 60 GWd/MTU fuel rod segment clad in E110 alloy and failed at a reported peak enthalpy of 165 cal/g. As described in the NEA RIA report, it is likely that RT-9 experienced transient FGR during the power pulse, which increased rod internal pressure. Additionally, RT-16 and RT-19 were below the line. RT-16 was also clad in E110 alloy and possessed a burnup of 75 GWd/MTU. RT-16 failed at peak fuel enthalpy of 145 cal/g. Moreover, RT-19 had a differing fuel design, a modern VVER fuel design. This modern VVER fuel design has a solid pellet as opposed to conventional VVER annular pellets and a thinner cladding with a larger diameter based on sponge zirconium. This rod was irradiated to 49 GWd/MTU and failed with a peak fuel enthalpy of 140 cal/g. RT-16 and RT-19 results were not available at the time of development of the 2015 interim criteria (Ref. 11). The RT-16 and RT-19 datapoints lie just below the red dotted line at a cladding differential pressure of 2 MPa.

Both differential pressure and elevated temperature are necessary to achieve cladding failure due to balloon rupture. Hence, at some point, the relationship between differential pressure and cladding failure (negative slope of 20 cal/g per MPa) is no longer valid, since the cladding temperature remains low enough to preclude balloon rupture. Figure 29 of the NEA RIA report provides measured cladding surface temperatures under RIA simulation tests in NSRR.

Examination of this figure reveals that cladding temperatures remain below 427 degrees Celsius

(°C) (800 degrees Fahrenheit (°F)) during tests where the peak radial average fuel enthalpy was less than 80 cal/g. Figure 6 of Udagawa et al. (2014) (Ref. 18) provides a measured cladding surface temperature based on the revised fuel enthalpies for the Japan Atomic Energy Agency (JAEA) NSRR test program. Consideration of the revised data increases the 227°C (800°F) threshold to above 100 cal/g peak radial average fuel enthalpy. Note that only the irradiated test results were considered since high rod internal pressure (above system pressure) may only occur in high exposure fuel rods.

NUREG-0630, Cladding Swelling and Rupture Models for LOCA Analysis, issued April 1980 (Ref. 19), provides cladding rupture models based upon mechanical testing of zirconium alloy tubes. An examination of NUREG-0630 reveals that cladding rupture at 427°C (800°F) is unlikely at end-of-life rod internal pressure design limits (proprietary limits established to avoid cladding liftoff). Note that the + phase transition temperature (approximately 727°C (1,340°F), depending on alloy composition) is well above 427°C (800°F). Based on this information, the staff has set a lower threshold for cladding failure of 100 cal/g.

Based upon a review of the entire empirical database and an analytical requirement to consider transient FGR, the staff has maintained the upper high-temperature failure threshold of 170 cal/g. This failure threshold considers both oxidation-induced embrittlement and balloon and burst failure modes. Combining this upper threshold with the cladding differential pressure relationship yields a composite high-temperature cladding failure threshold. The green line in

4-17 Figure 4-13 depicts this composite failure threshold in the midst of the empirical database, which matches that in the NEA RIA report.

Figure 4-13 High-Temperature Cladding Failure Threshold in Context The composite failure threshold is represented by the following equation:

High-Temperature Cladding Failure Threshold:

Cladding differential pressure < 1.0 MPa, Peak radial average fuel enthalpy = 170 cal/g Cladding differential pressure > 1.0 MPa, < 4.5 MPa Peak radial average fuel enthalpy = 170 - ((P - 1.0)*20) cal/g Cladding differential pressure > 4.5 MPa, Peak radial average fuel enthalpy = 100 cal/g Following a brief discussion on high-temperature oxidation tests, the NEA RIA report concludes as follows:

since it cannot be ruled out that the film-boiling phase in some scenarios for RIA in light water reactors may have longer duration, and that embrittlement-induced clad failure thus is possible at lower cladding temperatures and at lower fuel enthalpies than the typical pulse reactor threshold value of 1 050 J(gUO2)-1, acceptance criteria usually postulate that fuel rod failure should be assumed when film-boiling is predicted to occur under the reactivity-initiated accident.

4-18 For many reactor designs, technical specification limits on allowable rod insertion will preclude a prompt critical power excursion for at-power RIAs. Unlike narrow pulse prompt excursions, the broad power pulse experienced in these scenarios allows for cladding temperatures to rise and transfer heat to the coolant. As the rate and amount of reactivity insertion decreases, these scenarios begin to behave more like traditional overpower events (e.g., excess load, bank control element assembly withdrawal, rod withdrawal error). Since the empirical database used above to define the high cladding temperature failure does not encompass all power operating conditions (e.g., coolant conditions, DNB/CPR thermal margin), the staff maintains that, for at-power conditions (greater than 5 percent rated thermal power), any fuel rod predicted to exceed thermal design limits (DNB and CPR) must be assumed to fail and must be accounted for in dose calculations.

RG 1.77 established the presumption of cladding failure at the onset of DNB. However, RG 1.77 also included the following provision:

Other DNB or clad failure correlations may be used if they are adequately justified by analytical methods and supported by sufficient experimental data.

Alternative cladding failure criteria will be addressed on a case-by-case basis.

4.2.3 Applicability of High-Temperature Failure Threshold As described below in Figure 4-11, the RIA database encompasses a wide array of fuel rod designs, cladding alloys, and experimental conditions (e.g., pulse width, temperature). Fuel specimens were fabricated from BWR, PWR, and VVER commercial rods and research reactor fuel rods and included several commercial-grade cladding alloys: Zircaloy-2, Zircaloy-4, low tin Zircaloy-4, ZIRLO, M5, E110, New Developed Alloy (NDA), and Mitsubishi Developed Alloy (MDA). BWR Zircaloy-2 fuel rod samples include both liner (e.g., natural or low alloy Zircaloy-2 bonded to cladding inner diameter) and nonliner configurations. Based upon this comprehensive database, it is judged that the proposed non-PCMI, high-cladding-temperature fuel cladding failure criteria conservatively predict cladding failure (i.e., balloon/burst, post-DNB oxidation/embrittlement) at zero power conditions for both BWRs and PWRs.

Thermal-hydraulic conditions in the test capsules for the empirical database depicted in Figure 4-10 through Figure 4-13 (e.g., stagnant water at 1 atmosphere and room temperature) more closely resemble BWR cold start-up (also referred to as cold zero power (CZP)) conditions than PWR hot zero power (HZP) conditions. However, due to the rapid temperature excursion experienced by the cladding, the initial cladding temperature (and associated mechanical properties) is not as important as in the prompt PCMI-type failure mode. Given the relatively large coolant volumes, mass flow rate, and fluid conditions of PWR coolant systems, application of the empirically derived cladding failure threshold to PWR HZP conditions is conservative.

Section 6.2.5.2 of the NEA RIA report also describes the potential impact of annular fuel pellets on reported test results. Based upon observations of clad ballooning and burst in rods with annular pellets, whereas comparable rods with solid pellets survived, the NEA RIA report postulates that the central hole provides a channel for axial gas and a high gas pressure can therefore be maintained in regions where clad ballooning occurs (Ref. 15). This improved gas communication, relative to high burnup, closed gap solid pellet rods, makes fuel rods with annular pellets more susceptible to clad ballooning. The NEA RIA report also notes that annular pellets may affect PCMI, since a portion of the fuel thermal expansion will be accommodated by the central hole.

4-19 Figure 4-12 shows that the annular and nonannular test results support the same relationship between failure threshold and cladding differential pressure. Application of the high-temperature failure threshold (empirically derived based on both solid and annular pellet tests) to solid pellet designs is conservative. Further, modern PWR integral fuel burnable absorber fuel rod designs often include axial zones with annular fuel pellets. Thus, any attempt to divide the limited database into solid and annular subsets may not prove useful. In addition, modern BWR fuel assembly designs include part-length fuel rods. As a result, the distance between the rod plenum region and peak fuel enthalpy may be reduced, allowing for improved axial gas communication even at high burnup conditions.

In summary, the range of applicability is limited to the following conditions:

all PWR and BWR uranium dioxide (UO2) fuel rod designs with zirconium-based cladding, including barrier designs and fuel rods with both annular and solid pellets.

BWR cold startup conditions up through PWR hot zero conditions 4.2.4 Analytical Considerations for High-Temperature Failure Threshold In addition to the above observations, Section A.4 of the NEA RIA report further discusses transient FGR and its impact on gas pressure loading. The observed failure enthalpy for the irradiated rods, 650-700 J(gUO2)-1 or 155-167 cal/g, is similar to the RG 1.236 failure thresholds (i.e., 150-170 cal/g). However, the guidance does not explicitly account for transient FGR. Instead, the guidance provides failure thresholds based on pretransient rod internal pressure. Ignoring the potential contribution of transient FGR may lead to a nonconservative application relative to the empirical database.

In contrast, Section 3.2.2 of the EPRI report concludes that the NSRR test results demonstrate that transient FGR will not enhance cladding ballooning and burst. Based upon expected fuel rod conditions, including fission gas release, rod power history, and restricted axial gas flow, the EPRI report concludes that failure by ballooning and burst in high burnup, overpressure fuel rods is unlikely. The EPRI report also states that low to intermediate burnup fuel rods have internal gas pressure below system pressure, and therefore the driving forces are insufficient to produce ballooning deformations. While these observations have some merit, fuel rod designs and operating conditions continue to evolve, and it is prudent to define failure thresholds based on limiting conditions. For example, integral fuel burnable absorber fuel designs may promote higher rod internal pressure at a given burnup due to (1) higher FGR for an equivalent rod power and burnup level (e.g., gadolinium, erbium) and (2) potential helium generation (e.g., boron coating).

RG 1.236 states that the cladding differential pressure should include both the initial, pretransient rod internal gas pressure and any increase associated with transient FGR. An approved fuel rod thermal-mechanical performance code should be used to predict the initial, pretransient rod internal conditions (e.g., moles of fission gas, void volume, FGR, rod internal pressure). The amount of transient FGR may be calculated using the burnup-dependent correlations in Appendix B to RG 1.236.

4-20 4.3 Hydrogen-Enhanced PCMI Cladding Failure Threshold 4.3.1 Hydride Formation and Orientation During operation, the zirconium cladding will react with the coolant to form an oxide layer in the corrosion process, as characterized by the reaction below:

+ 222 + 22 As can be seen above, hydrogen is produced in the corrosion reaction. A fraction of the hydrogen produced from the corrosion process will be absorbed by the oxide layer and the base cladding metal, while the remainder is released into the coolant. The fraction of hydrogen produced in the corrosion process that is absorbed by the base cladding metal is often referred to as the hydrogen pickup fraction. This is an alloy-specific property that is often represented as a constant value or a function of burnup (Ref. 20). The hydrogen in the cladding will be in solution until the concentration of hydrogen reaches the solubility limit. The solubility of hydrogen in zirconium alloys is expressed in the following equation:

= 1.2 x 105 exp (

8550 1.985887) (Ref. 21)

Where is hydrogen solubility in weight parts per million (wppm)

T is temperature in Kelvin At PWR hot operating conditions, hydrogen solubility ranges from 50 wppm (279°C (535°F)) to 150 wppm (371°C (700°F)). At BWR cold startup conditions, hydrogen solubility is less than 1.0 wppm (21°C (70°F)). Any hydrogen above solubility precipitates as zirconium hydride platelets. In this report, the term excess hydrogen refers to the amount of hydrogen above solubility (i.e., precipitated hydrogen).

Hydrogen content and hydride distribution and orientation have a first-order impact on the PCMI resistance of fuel rod cladding. The amount of hydrogen absorbed during normal operation depends on many factors, including the corrosion resistance of the zirconium alloy, time at elevated temperature, and hydrogen pickup fraction of the zirconium alloy (which may also depend on fluence). The orientation of the hydrides is affected by the thermal and mechanical treatment of cladding tubes during manufacturing and by the stress state prevailing under hydride precipitation (Ref. 15).

Figure 4-14 shows the distribution and orientation of hydrides in high burnup fuel rod cladding segments. Figure (a) illustrates the characteristic distribution and random orientation (including radial) of hydrides in a high burnup M5 recrystallization annealed (RXA) cladding. Figure (b) illustrates the characteristic distribution with a pronounced rim and circumferential orientation of the hydrides in a high burnup Zircaloy-4 and ZIRLO stress relieved annealed (SRA) cladding.

The accumulation of hydrides on the outer rim of the cladding is a result of the thermal gradient across the cladding. Hydrogen will tend to migrate toward the colder and high-stress regions of the cladding.

4-21 Figure 4-14 Hydride Orientation and Distribution in Various High Burnup Claddings Hydrides will typically precipitate circumferentially, as shown in the figure, but under high hoop stresses, they may reorient radially.

Section 6.2.6.3 of the NEA RIA report (Ref. 15) concludes the following with respect to the effect of hydride distribution and orientation:

For the same average hydrogen content, materials with uniformly distributed hydrides are more ductile than those having local concentrations of hydrides in certain regions.

Ductility decreases rapidly with increasing thickness of the hydride rim at the clad outer surface. This embrittlement seems to saturate beyond a rim thickness of 100 m [micrometers] and it is not clear whether the effects of oxide thickness are additive.

4-22 In BWR Zircaloy-2 cladding with an inner liner of zirconium (e.g., barrier cladding), local concentrations of hydrides were observed not only on the clad outer surface, but within the liner zirconium. It seems that the liner material, due to its lower solubility for hydrogen, is more sensitive to hydrogen-induced embrittlement than the Zircaloy-2 base material.

The orientation of zirconium hydrides platelets strongly influences embrittlement.

The more deleterious radial hydrides are more common in RXA materials; however, [they] may appear in any material as a result of applied tensile stress or residual stress (e.g., rod internal pressure). Radial hydrides may also form as a result of power cycling or overpower transients; where existing circumferential hydrides may dissolve by the temperature excursion, and then re-precipitate as radial hydrides, if the tensile hoop stress is large enough.

As described above, the EPRI critical strain energy density (CSED) is based on separate effects mechanical testing of irradiated cladding specimens. Figure 4-17 compares the predicted CSED (total elongation (TE)) for Zircaloy-4 (SRA) and Zircaloy-2 (RXA) test results described in EPRI Report 3002005540, Fuel Reliability Program: Proposed Reactivity Insertion Accident (RIA)

Acceptance Criteria, Revision 1, issued November 2015 (Ref. 22). Examination of this figure confirms that the RXA material is more susceptible to hydrogen-induced embrittlement under a variety of loading conditions.

Figure 4-15 Comparison of EPRIs CSED(TE) correlations for Zircaloy-4 and Zircaloy-2 at CZP

4-23 Research into hydride reorientation during prestorage drying-transfer operations and early stage storage provides further evidence of the detrimental role of radial hydrides on cladding ductility.

Researchers varied drying conditions (i.e., temperature and loading) to study their influence on hydride reorientation. Mechanical testing was subsequently performed on these test specimens to quantify the impact of varying degrees of hydride reorientation on the mechanical properties of these high burnup fuel cladding segments. Examination of the test results documented in Aomi et al. (2008) (Ref. 24) and Billone et al. (2013) (Ref. 25) shows a decrease in cladding ductility with increasing proportion of radial hydrides. While the ring compression tests performed in these studies are not prototypical of the loading expected during an RIA, the results provide further evidence of the increased sensitivity of cladding to hydrogen-induced embrittlement with the introduction of radial hydrides versus the sensitivity of the same material with the same amount of circumferential hydrides.

While the above separate effects mechanical testing demonstrates the increased sensitivity of cladding containing radial hydrides to hydrogen-induced embrittlement, the true effect of radial hydrides on hydrogen-enhanced PCMI under RIA loading conditions may be best observed by studying a subset of the RIA empirical database. Figure 4-16 shows a plot of fuel enthalpy versus excess hydrogen content for the cold NSRR pulse reactor RIA tests. The data are divided between SRA and RXA cladding material. Examination of this figure confirms that the RXA material is more susceptible to hydrogen-induced embrittlement under RIA loading conditions.

Figure 4-16 NSRR RXA vs. SRA fuel enthalpy rise In recognition of the first-order effect that zirconium hydride distribution and orientation have on mechanical properties, RG 1.236 segregates the RIA empirical database into SRA and RXA

4-24 cladding types, and unique PCMI failure thresholds as a function of excess hydrogen were developed for each type of cladding microstructure.

4.3.2 Impact of Fuel Burnup During irradiation, each fuel rod undergoes several evolutions that alter its behavior under RIA conditions. These include, but are not limited to, the following:

Irradiation-induced fuel swelling and cladding creep reduce the size of the as-fabricated pellet-to-cladding gap. The rate and timing of gap closure depends on fuel design and power history. Once gap closure is complete and hard contact predicted, the fuel rod is more susceptible to PCMI failure under RIA conditions.

Self-shielding promotes a higher exposure, plutonium rich region toward the pellet periphery. As a result, the pellet radial power distribution is skewed toward the pellet periphery during RIA power pulses.

Key core physics parameters change with core depletion.

Control rod worth changes with exposure.

Reactor kinetics change with exposure. For example, the fraction of delayed neutrons,, decreases with fuel exposure. This change in influences the dynamic behavior of the reactor under transients.

The fuel temperature reactivity coefficient, Doppler, decreases (becomes more negative) with fuel exposure.

Cladding corrosion increases with time-at-temperature.

Fuel thermal conductivity and solidus temperature decreases as a function of exposure.

Sections 3.5 and 3.8 of Pacific Northwest National Laboratory (PNNL) Report 22549, Pellet-Cladding Mechanical Interaction Failure Threshold for Reactivity Initiated Accidents for Pressurized Water Reactors and Boiling Water Reactors, issued June 2013 (Ref. 26), describe the potential effects of fuel burnup and initial gap size on PCMI susceptibility. Based on an examination of the RIA empirical database, PNNL concludes that only a weak burnup dependence is observed. FRAPCON-3.4 and FRAPTRAN-1.4 calculations were used to explore the dependency on initial gap size and timing of gap closure. An analytical approach was necessary since all of the PCMI failed rods in the empirical database were beyond 33 GWd/MTU and likely exhibited gap closure and hard contact (before the RIA test). PNNLs calculations were based on Zircaloy-4 cladding, which is conservative with respect to time-dependent PCMI susceptibility relative to modern alloys. Based upon the predicted timing of gap closure for two different PWR fuel rod designs, PNNL concludes that the initial as-fabricated gap has little or no effect on the PCMI failure threshold and that low burnup fuel rods with an existing gap will likely fail by ballooning and rupture.

The EPRI report and the NEA RIA report support the position that low burnup, low-corrosion fuel rods will likely fail by high-temperature modes before PCMI.

4-25 In conclusion, the timing of pellet-to-cladding gap closure, which is strongly dependent on fuel rod design and power history, is such that low burnup fuel rods with an existing gap will likely fail by high-temperature modes before PCMI. Nevertheless, application of an empirically based PCMI failure threshold based on medium to high burnup rods (with closed gaps and hard contact) is conservative over the entire range of burnup, especially for low to medium burnup ranges where an existing gap may accommodate a portion of the pellet thermal expansion.

4.3.3 Impact of Pulse Width Table 4-3 lists the minimum and maximum pulse widths from the various RIA test facilities, along with predicted pulse widths for commercial PWRs and BWRs. Examination of this table reveals that many of the NSRR and CABRI tests experienced shorter pulse widths than would be expected during postulated accidents.

Table 4-3 Pulse Width Variability Scenario/Test Facility Pulse Width (ms)

Minimum Maximum PWR HZP CRE 25 65 PWR HFP CRE 400 4500 BWR CZP CRD 45 75 BWR HZP CRD 45 140 SPERT-CDC 13 51 PBF 11 16 IGR 750 950 BIGR 2.5 5.8 NSRR 4.3 9.0 CABRI 8.8 75 Notes:

(1) Estimated PWR and BWR pulse widths are based on realistic and moderately conservative computer analyses. Data are from Section 2.2 of the NEA RIA report (Ref. 15).

(2) Ranges of pulse widths are from RIA test facilities in an empirical RIA database.

Section 3.7 of PNNL Report 22549 investigated pulse width effects on the proposed PCMI cladding failure thresholds. This investigation concludes that no obvious bias with pulse width between 4.4 and 76 millisecond (ms) exists in the NSRR and CABRI failure data set. Further investigation using the FRAPTRAN transient fuel rod performance code revealed that the predicted cladding temperature will increase with pulse widths greater than 20 ms, and the higher cladding temperatures should increase the failure threshold. For example, the predicted failure enthalpy for a standard 17x17 PWR rod with Zircaloy-4 cold work SRA cladding at 40 GWd/MTU was 100 cal/g with a 5 ms pulse width. The predicted failure enthalpy increased to 145 cal/g with a 30 ms pulse width. Recognizing analytical limitations and experimental observations of brittle failure in the hydride rim at the outer diameter, the PNNL report concludes that either the ductility (fracture toughness) of the hydride rim is not sensitive to temperature, or the rim temperature does not increase sufficiently before the crack initiates in a 75 ms pulse to increase the ductility (fracture toughness) of the hydride rim. Based on the above studies, the PNNL report concludes that the PCMI cladding failure empirical correlation should be used with no adjustment for pulse width.

4-26 While both reports recognize the importance of pulse width on the overall fuel performance, neither the EPRI report nor the NEA RIA report attempts to scale the empirical data to account for differences between the pulse reactor and commercial reactor characteristic pulse width.

Based on the above discussion, the staff decided not to scale the data for the potential impact of pulse width.

4.3.4 Cladding Temperature Effects Differing opinions on the magnitude of scaling for the cold NSRR test results to PWR HZP operating conditions were an important contributor to a staff decision to await the NSRR hot capsule tests before finalizing the RIA criteria and guidance. Temperature affects the cladding yield strength, ductility, and hydrogen solubility (which also affects ductility).

Section 6.2.6.2 of the NEA RIA report concludes the following with respect to the effect of cladding temperature:

The ductile-to-brittle temperature for irradiated cladding increases with increasing hydrogen.

The ductile-to-brittle temperature for a given hydrogen content increases with increasing strain rate.

Experimental results demonstrate that the time necessary for hydride dissolution exceeds the timing of PCMI loading during the early stage of an RIA.

Section 3.3 of PNNL Report 22549 investigated cladding temperature effects on the proposed PCMI cladding failure thresholds. Based upon comparisons of similar hot and cold RIA test results, PNNL concluded that an adjustment of 18-20 cal/g on cold NSRR failure enthalpy is appropriate to account for HZP operating conditions. Furthermore, based on results from NSRR RIA tests FK10 and FK12, PNNL concluded that the ductile-to-brittle transition temperature for irradiated Zircaloy-2 cladding is beyond 85°C (185°F). Hence, the available data do not support the proposed scaling of the BWR data and resulting PCMI failure threshold. Figure 4-17 shows the magnitude of PNNLs proposed scaling. Note that, since PNNLs failure curves are expressed in terms of excess hydrogen, the effective scaling factor would also include the worth of the increase in hydrogen solubility at the higher temperature.

As described above, EPRI used the FALCON fuel rod thermal-mechanical model to scale several of the cold NSRR test results (i.e., measured failure enthalpy) to PWR hot operating temperatures based upon CSED (uniform elongation (UE)), CSED(TE), and a lower bound CSED(TE). Figure 4-17 shows the magnitude of scaling (cal/g) based on EPRIs CSED(TE) and PNNLs proposed scaling. EPRIs CSED(TE) data were approximated from Figure 4-13 in EPRI Report 3002005540 (Ref. 22). Note that since PNNLs failure curves are expressed in terms of excess hydrogen, the effective scaling factor would also include the worth of the increase in hydrogen solubility at the higher temperature.

4-27 Figure 4-17 Comparison of EPRIs CSED(TE) and PNNLs Temperature Scaling Factors Examination of Figure 4-17 reveals that the PNNL scaling factor lies below EPRIs CSED(TE) scaling factor. Since the data scaled using CSED(UE) were not illustrated in EPRI Report 3002005540 as they were in Revision 0 of EPRI Report 1021036, they were not included in Figure 4-17. The CSED(UE) scaling factor is expected to be below the PNNL scaling factor, as was shown in Figure 3.2.2-15 of the NRC memorandum of March 6, 2015 (Ref. 11).

Based upon the above discussion, the staff decided to apply the PNNL scaling factor to adjust test results from BWR cold startup conditions to PWR hot operating conditions. Since the NSRR test conditions are prototypical of BWR cold startup conditions, no adjustment is being applied.

4.3.5 Mixed Oxide Fuel Based upon the reported failure enthalpies between standard UO2 and mixed oxide (MOX) fuel rods, PNNL concludes that there is no conclusive evidence of an inherent MOX effect. In Section 5 of EPRI Reports 1021036 and 3002005540 (Refs. 17 and 22), EPRI concludes that gaseous swelling-enhanced PCMI is an important contributor to the mechanical behavior of MOX fuel rods under RIA conditions, especially at operating temperatures. The NEA RIA report notes that MOX effects have been observed and this issue needs further investigation. The staff agrees that MOX effects require further investigation.

Based on the above discussion, the staff decided to limit the applicability of the PCMI failure thresholds in RG 1.236 to UO2 fuel rods.

4-28 4.3.6 RG 1.236 PCMI Cladding Failure Threshold Figure 4-18 through Figure 4-21 show the empirically based PCMI cladding failure thresholds.

Because fuel cladding ductility is sensitive to hydrogen content, zirconium hydride orientation, and initial temperature, separate PCMI failure curves are provided for RXA and SRA cladding types at both low initial cladding temperature conditions (i.e., below 260°C (500°F) down to BWR cold startup) and high initial cladding temperature conditions (i.e., at or above 260°C (500°F)). The RXA cladding failure threshold is further refined for cladding designs with and without a barrier liner (e.g., sponge or low-alloy cladding inside-diameter liner). The SRA cladding failure threshold is applicable regardless of the presence of a barrier liner. The PCMI cladding failure threshold is expressed in peak radial average fuel enthalpy rise (cal/g) versus excess cladding hydrogen content (wppm). Excess cladding hydrogen content refers to the portion of total hydrogen content in the form of zirconium hydrides (i.e., it does not include hydrogen in solution).

Figure 4-18 PCMI Cladding Failure ThresholdRXA Cladding at or Above 260°C (500°F)

4-29 Figure 4-19 PCMI Cladding Failure ThresholdSRA Cladding at or Above 260°C (500°F)

Figure 4-20 PCMI Cladding Failure ThresholdRXA Cladding Below 260°C (500°F)

4-30 Figure 4-21 PCMI cladding Failure ThresholdSRA Cladding Below 260°C (500°F) 4.3.7 Hydrogen Enhanced PCMI Cladding Failure Thresholds Technical Bases Guidance before 2007 did not identify PCMI as a fuel cladding failure mechanism. As a result, the number of failed fuel rods may have been underestimated. The 2007 NRC RIA memorandum (Ref. 10) and SRP Section 4.2, Revision 3, defined interim PCMI failure thresholds by two curvesone for PWRs and one for BWRs. In the years following the 2007 NRC RIA memorandum, new data led to a revision of the PCMI failure thresholds and the establishment of more comprehensive failure curves in 2015 (Ref. 11). The updated interim guidance included four unique failure thresholdsone each for RXA and SRA cladding at BWR cold startup conditions and PWR operating conditions. The information published after 2007 that helped lead to the development of the 2015 interim criteria is summarized below from the 2015 NRC memorandum (Ref. 11):

OECD NEA State-of-the-art Report, Nuclear Fuel Behaviour under Reactivity-initiated Accident (RIA) Conditions, 2010 (Ref. 15)

The influence of temperature, hydride distribution, and hydride orientation on hydrogen-enhanced PCMI cladding failure thresholds is described.

EPRI Report 1021036, Fuel Reliability Program: Proposed RIA Acceptance Criteria, December 2010 (Ref. 17)

Proposed PCMI cladding failure thresholds are provided for PWR SRA and BWR RXA cladding alloys. Failure thresholds are expressed as a function of radial average fuel enthalpy increase versus cladding hydrogen.

4-31 The thresholds are based on a CSED cladding failure function derived from separate-effects mechanical testing and FALCON predicted strain energy density (SED) under RIA conditions.

PNNL Report 22549, Pellet-Cladding Mechanical Interaction Failure Threshold for Reactivity Initiated Accidents for Pressurized-Water Reactors and Boiling-Water Reactors, June 2013 (Ref. 26)

Proposed PCMI cladding failure thresholds are provided for cold RXA, hot RXA, cold SRA, and hot SRA cladding alloys. Failure thresholds are expressed as a function of radial average fuel enthalpy increase versus cladding excess hydrogen.

The thresholds are based on an empirical database of failed rods under RIA test conditions.

published results from NSRR Hot Capsule RIA Test VA3, VA4, RH2, BZ3, and LS2 (see Appendix A, Ref. 26).

The test results from the NSRR hot capsule program provide valuable insights on the effect of cladding temperature on failure enthalpy. This information was used to scale the earlier NSRR RIA test results performed at room temperature to PWR hot conditions.

JAEA-published revised fuel enthalpy predictions for 43 previous NSRR test specimens (Ref. 18)

JAEA data were relied upon to develop the empirically based cladding failure criteria; hence, the revised data had a direct impact.

The proposed PNNL and EPRI PCMI failure thresholds detailed in the above references were derived before the publication of the revised NSRR data. Section 3.2.2.1 of the 2015 NRC RIA memorandum (Ref. 11) describes the impact of the revised NSRR data (Ref. 18) as of 2015.

After release of the 2015 NRC RIA memorandum, EPRI revised its proposed RIA acceptance criteria in EPRI Report 3002005540. This revision considered the revised NSRR data listed in Udagawa et al. (2014) (Ref. 18). As stated in Section 4.1 of this report, since 2015, further modifications and additions have been made to the NSRR, CABRI, and BIGR RIA data. These changes are reflected in data presented in the graphs throughout this report. The impact of the updates to the RIA database since 2015 will be examined later in this section.

Figure 4-22 compares the RG 1.236 PCMI failure threshold for SRA zirconium cladding at operating conditions with several different thresholds: (1) the NRCs 2015 interim PCMI failure threshold, (2) EPRIs proposed threshold in 2015, and (3) a revised failure threshold using PNNLs methods on the updated RIA data since 2015. Figure 4-23 compares the SRA cladding failure thresholds at BWR cold startup conditions. Figure 4-24 and Figure 4-25 provide similar comparison plots for RXA cladding. Note that the EPRI failure thresholds are based on SRA cladding and therefore may not be directly applicable to RXA cladding.

4-32 Figure 4-22 Comparison of SRA Cladding HZP PCMI Failure Thresholds Figure 4-23 Comparison of SRA Cladding CZP PCMI Failure Thresholds

4-33 Figure 4-24 Comparison of RXA Cladding HZP PCMI Failure Thresholds Figure 4-25 Comparison of RXA Cladding CZP PCMI Failure Thresholds

4-34 The 2015 NRC interim PCMI failure thresholds for SRA cladding at CZP and HZP conditions were based on the revised NSRR data as of 2015 and used the method presented in the PNNL report, which was a natural logarithm least squares (LSQ) fit to the data. The revised failure threshold using PNNLs methods, denoted as Revised PNNL in Figure 4-22 and Figure 4-23, used this same methodology on the most recent RIA empirical database at the time of the publication of this report. Comparison of the 2015 NRC interim PCMI failure thresholds and the revised PNNL threshold for SRA cladding at CZP and HZP conditions reveals that the modifications and additions to the empirical RIA database since 2015 do not have a significant effect.

For RXA cladding in CZP and HZP conditions, the data for failed rods that is available lies at excess hydrogen concentrations greater than 150 wppm. Due to the lack of failed rod data below 150 wppm excess hydrogen, the extrapolation of the natural logarithm LSQ fit curve to lower hydrogen concentrations is not justified. As a result, the NRC adopted a different approach for the 2015 interim failure thresholds for RXA cladding, a broken linear relationship.

In DG-1327, which was later revised and finalized as RG 1.236, the NRC proposed a curve with the form of a

• Hb for the RXA failure curves. In DG-1327, the SRA failure curves continued to be based on a natural logarithmic LSQ fit.

The failure curves in RG 1.236 were not developed in the same manner as the DG-1327 interim thresholds (i.e., using a natural logarithmic LSQ fit to the data for the SRA cladding and a

• Hb for the RXA cladding). When soliciting a second round of public comments on DG-1327 (Ref. 27), Global Nuclear Fuel (GNF) (Ref. 28) and the Nuclear Energy Institute (Ref. 29) argued that the NRCs draft curves were inaccurate between 55 to 100 wppm hydrogen, especially for the CZP RXA cladding failure curve. To develop a more accurate failure threshold at low hydrogen levels, where there are fewer failure data in the empirical RIA database, the data were fit to an alternate exponential function (i.e., a

• Hb + c). This equation incorporates the nonfailed data and hence provides a more precise threshold in the region of interest. The exponential function also improves the curve by better representing (1) the rapid loss in RXA cladding ductility as zirconium hydrides form and (2) the saturation effect at higher concentrations of zirconium hydrides. The commenters provided both best fit and lower bound coefficients. Figure 4-26 compares the RXA CZP DG-1327 curve with the RG 1.236 curve.

4-35 Figure 4-26 Modified Exponential Function Curve Fit for RXA cladding at CZP Figure 4-26 shows that the proposed 2019 DG-1327 failure curve had several data points between 55 and 100 wppm excess hydrogen that were misrepresented by being above the curve, wrongly indicating failure. The primary reason for the inaccuracy of the 2019 DG-1327 curve at low hydrogen levels was because the 2019 DG-1327 RXA curves were developed using a curve fit to the failed data rather than a broken linear relationship, as was done in the 2015 NRC interim guidance and the original DG-1327 RXA curves. A curve fit was used to expand the RXA curves above 300 wppm excess hydrogen, incorporating the VA-6 test point that was newly reported at the time and is located above 700 wppm excess hydrogen. As shown in Figure 4-26, below 150 wppm excess hydrogen, there are no RXA cladding failures in the database. To address this lack of data, the highest enthalpy point was treated as a failure point for the curve fit, which resulted in several nonfailed points at low hydrogen levels being situated above the failure threshold (Ref. 30). The RG 1.236 curves clearly better capture the data points below 150 wppm excess hydrogen. A lower bound curve was adopted to capture the lower of the two datapoints above 300 wppm excess hydrogen (i.e., NSRR VA-6), the lower bound curve intersects, while a best estimate curve would be above the NSRR VA-6 test datapoint.

The NRC revised all four PCMI cladding failure threshold curves using the modified form of the equation. Coefficients were selected to better represent the nonfailed data at low hydrogen levels and bound much of the failed data at higher hydrogen levels. The final curves presented in RG 1.236 reflect these revisions.

4.3.7.1 EPRI CSED(TE) and CSED(UE)

As previously stated, EPRI originally proposed PCMI failure curves using CSED in EPRI Report 1021036 from 2010 and revised the curves in Revision 1 to the report in 2015.

0 25 50 75 100 125 150 175 200 0

50 100 150 200 250 300 350 400 450 500 550 600 650 700 750 800 Peak Radial Average Fuel Enthalpy Rise ( cal/g)

Excess Cladding Hydrogen (wppm)

RXA Cladding - Adjusted CZP Fuel Enthalpy Rise Versus Hydrogen Open Symbol = Non-failed peak fuel enthalpy Closed Symbol = Fuel enthalpy at failure CABRI NSRR DG-1327 (2019)

RG 1.236 Rev. 0

4-36 Revision 1 of the report was published after the 2015 RIA memorandum, so the memorandum described EPRIs 2010 curves.

As described in Section 2.3.2 of Revision 1 of the EPRI report (Ref. 22), EPRIs CSED is developed from material property tests as a function of material conditions, including temperature, fast fluence, outer surface corrosion, hydrogen concentration, and hydride morphology. The database of mechanical property tests on irradiated cladding material used to develop the CSED relations contains a variety of cladding designs, irradiation conditions, corrosion (oxide thickness and hydrogen concentration), testing conditions (e.g., temperature, stress state, strain rate), and anticipated cladding damage mechanisms (e.g., hydride lenses, hydride rim, spalled oxide, cracks). The overall goal of this section is to define the threshold for cladding failure; specifically, the point at which a through wall crack (i.e., breach) in the cladding may allow the release of fission gas. The goal is not to define gross cladding failure, as that is addressed in separate criteria related to coolable geometry. For the purpose of defining the threshold of cladding failure, measured UE is a more appropriate and repeatable characterization of the fuel rod claddings resistance to PCMI than TE.

The PCMI cladding failure thresholds proposed in the EPRI report are based on TE data (i.e., CSED(TE)). However, in response to staff concerns, the EPRI report also includes CSEDs based on UE data. Figure 4-27 shows EPRIs Zircaloy-4 (SRA) CSED(TE) and CSED(UE) as a function of cladding hydrogen content, as reported in EPRI Report 3002005540. As expected, limiting the failure data to UE results in a significant reduction in CSED. Differences between CSED(TE) and CSED(UE) are strongly influenced by temperature. Cladding hydrogen content and hydride distribution and orientation are also likely to influence this relationship.

Figure 4-27 EPRI CSED for Zircaloy-4 (SRA) 0.00 5.00 10.00 15.00 20.00 25.00 30.00 35.00 40.00 0

100 200 300 400 500 600 CSED (MJ/m^3)

Cladding Hydrogen (ppm)

EPRI CSED: Zircaloy-4 (SRA)

Hot, TE, 5 ms pulse Hot, TE, 10 ms pulse Hot, UE Cold, TE Cold, UE

4-37 Using its FALCON fuel rod thermal-mechanical model, EPRI scaled several of the cold NSRR test results (i.e., measured failure enthalpy) to PWR hot operating temperatures based upon CSED(TE) with a 10 ms pulse width. Figure 4-13 of EPRI Report 3002005540 (Ref. 22), plots the scaled failure data along with the proposed Zircaloy-4 (SRA) HZP failure threshold. Scaling the data with CSED(UE) would result in a significant reduction in the temperature scaling factor relative to the CSED(TE). EPRI states that, if the data are scaled using the CSED(UE) failure threshold, the NRC interim criteria would be reproduced. Figure 4-22 depicts the proposed EPRI failure threshold for Zircaloy-4 as a function of excess hydrogen.

Similarly, as shown in Figure 4-25, EPRIs Zircaloy-2 CSED(TE) functions were used to develop an updated proposed curve for RXA cladding. EPRI proposed a flat limit of 150 cal/g for hydrogen levels up to approximately 600 wppm.

EPRI proposed its 2015 CSED-based failure thresholds in EPRI Report 3002005540 and formally as a comment to the NRC in the first public comment period for DG-1327 (Ref. 31). In response to EPRIs comments, the NRC stated that it elected to develop cladding failure thresholds based on in-pile testing rather than the EPRI methodology, which is based on separate effects tests (Ref. 30).

4.3.7.2 Effect of Liner on RXA Cladding During the public comment period for DG-1327, industry comments proposed the development and incorporation of PCMI failure limit curves for nonliner RXA cladding by adjusting the hydrogen content of the liner RXA cladding tests to account for the influence of the liner on the hydrogen distribution in the bulk cladding material (Refs. 27, 32). Many BWR RXA claddings possess liners to help prevent pellet-cladding interaction stress-corrosion cracking failures.

BWR liner (i.e., barrier) fuel typically has a natural or low alloy zirconium liner that acts as a sponge for hydrogen. The liner depletes the base metal of hydrides and their detrimental effect.

In other words, there is a preferential accumulation of hydrogen in the liner at the interface with the bulk cladding. The liner will remain ductile even with a high concentration of hydrides present.

Attempts to quantify the soaking effect that the liner has on the bulk cladding hydrogen concentration demonstrated that, on average, the bulk hydrogen content is reduced by approximately 21 percent relative to the total hydrogen value, with an upper bound of less than 30 percent (Ref. 32). To develop a PCMI failure curve for unlined RXA cladding, an upper bound adjustment of 30 percent was performed (i.e., 30 percent of total hydrogen content resides in the liner). In other words, the excess hydrogen content reported for NSRR tests with a lined cladding (i.e., FK series) was reduced by 30 percent to account for the liner (Ref. 27). This data adjustment for the liner is depicted in Figure 4-28 as well as the RG 1.236 failure curves for lined and unlined cladding.

4-38 Figure 4-28 Data Adjustment for RXA Cladding with a Liner Based upon the reported failure enthalpies between standard UO2 and MOX fuel rods, PNNL concluded that there is no conclusive evidence of an inherent MOX effect. In Section 5 of EPRI Reports 1021036 and 3002005540 (Refs. 17 and 22), EPRI concludes that gaseous swelling-enhanced PCMI is an important contributor to the mechanical behavior of MOX fuel rods under RIA conditions, especially at operating temperatures. The NEA RIA report notes that MOX effects have been observed and this issue needs further investigation. The staff agrees that MOX effects require further investigation.

Based on the above discussion, the staff decided to limit the applicability of the PCMI failure thresholds in RG 1.236 to UO2 fuel rods.

4.3.8 Applicability of PCMI Failure Thresholds PWR technical specification Power Dependent Insertion Limits restricts control bank insertion (as well as alignment and overlap), which limits the worth of any individual control rod. Core neutronics calculations, especially for intermediate and full power operating conditions, may predict maximum ejected rod worth below $1 reactivity (/ < 1.0), indicating a nonprompt power excursion.

The above discussion relates to the potential pulse width effects on experimental data. The RIA empirical database used to derive the PCMI cladding failure thresholds consists of prompt critical, narrow-pulse power excursion experiments. Cladding failures reported in the broad pulse IGR facility were due to high-temperature failure mechanisms. Therefore, these PCMI cladding failure thresholds may not be directly applicable to (1) nonprompt RIA scenarios (e.g., ejection of partially inserted control rod or low-worth control rod), (2) nonprompt accident

4-39 overpower scenarios (e.g., PWR main steamline break), or (3) nonprompt AOO overpower scenarios (e.g., PWR control rod bank withdrawal, BWR turbine trip). Relative to a prompt-critical, narrow-pulse power excursion, the broader power excursion exhibited in these scenarios allows additional time for the cladding temperature (and ductility) to increase such that brittle PCMI failure is less likely.

As shown in Figure 4-2, no experimental data exist between 76 ms and approximately 700 ms pulse width. Hence, it is difficult to determine a threshold for PCMI brittle failure susceptibility.

As such, the NRC staff recommends applying the PCMI cladding failure thresholds (in addition to high-temperature thresholds) to all PWR CRE and BWR CRD acident scenarios. For the purpose of calculating fuel enthalpy for assessing PCMI failures, the fuel enthalpy rise is defined as the radial average fuel enthalpy rise at the time corresponding to one pulse width after the peak of the initial pulse.

With the exception of FK-10 (80°C (176°F)) and FK-12 (85°C (185°F)), the databases supporting the BWR cladding failure criteria were all developed from an initial temperature of 20°C (68°F), consistent with cold startup BWR conditions. Due to temperature and hydrogen solubility effects, application of the BWR cladding failure criteria to higher operating temperatures is conservative. Future evaluation of hydrogen solubility and temperature effects may be pursued to refine the BWR PCMI failure criteria for application to higher operating temperatures. Another potential conservatism in the proposed criteria is the short NSRR pulse width relative to the broader pulse width of operating BWRs.

Based upon the effect of hydride distribution and orientation, separate PCMI failure curves were developed for SRA and RXA cladding materials. Ideally, each applicant should provide evidence of the hydride distribution and orientation in the irradiated condition to demonstrate applicability of the PCMI failure curves to its cladding alloy. As described in Udagawa et al. (2014) and the NEA RIA report, hydride reorientation from the circumferential direction to the radial direction is possible when the fuel rod is heated and subsequently cooled under an applied tensile load (e.g., high rod internal pressure). RG 1.236 states that each applicant should address the possibility of hydride reorientation.

In summary, the range of applicability is limited to the following conditions:

Approved LWR fuel rod designs comprise slightly enriched UO2 ceramic pellets (up to 5.0 weight percent uranium-235) within cylindrical zirconium-based cladding, including designs with or without barrier lined cladding, an integral fuel burnable absorber (e.g., gadolinium), or a pellet central annulus irradiated up to a maximum rod average burnup of 68 GWd/MTU.

These are BWR cold startup conditions up through PWR hot full power operating conditions.

The RXA PCMI cladding failure threshold curves apply to cladding that has undergone final thermal treatment that produces an RXA metallurgical state, while the SRA PCMI cladding failure threshold curves apply to cladding that has undergone final thermal treatment that produces an SRA metallurgical state. For any other metallurgical condition (e.g., partially RXA), the licensee or applicant should justify its similarity to either the SRA or RXA metallurgical condition.

4-40 4.3.9 Analytical Considerations In the application of the PCMI cladding failure thresholds, an NRC-approved alloy-specific cladding corrosion and hydrogen uptake model should be used to predict the initial, pretransient cladding hydrogen content. These approved models should account for the influence of (1) time at temperature (e.g., residence time, operating temperatures, steaming rate), (2) cladding fluence (e.g., dissolution of second-phase precipitates), (3) enhanced hydrogen uptake mechanisms (e.g., shadow corrosion, proximity to dissimilar metal), and (4) crud deposition, either directly or implicitly through the supporting database. As an alternative, Appendix C of RG 1.236 presents acceptable alloy-specific hydrogen uptake models to estimate pretransient cladding hydrogen content.

The approved application methodology should address the uncertainties in the analytical methods being used to demonstrate compliance. An approved core physics model that accounts for biases and uncertainties should be used to provide high-confidence predictions (e.g., control rod worth, local power peaking, Doppler feedback, fuel temperature).

When applying the hydrogen-dependent PCMI cladding failure curves, the cladding average (e.g., mid wall) temperature at the start of the transient should be used to define the excess hydrogen in the cladding. Use of the Kearns solubility correlation (Ref. 21) is acceptable. To calculate peak fuel enthalpy rise for CZP conditions, zero fuel enthalpy is defined at 20°C (68°F).

Accident analyses should consider the full range of cycle operation from beginning of cycle to end of cycle. At HZP, analyses should encompass both (1) beginning of cycle following core reload and (2) restart following recent power operation.

Fuel enthalpy calculations should account for burnup-related effects on reactor kinetics (e.g., eff, l*, rod worth, Doppler effect) and fuel performance (e.g., pellet radial power distribution, fuel thermal conductivity, fuel-clad gap conductivity, fuel melting temperature).

Many of the primary analytical considerations for applying the RG 1.236 PCMI were discussed above. Section 2 of RG 1.236 includes a complete list of the analytical considerations.

4.4 Molten Fuel Failure Threshold The melting of the fuel can lead to molten fuel-induced swelling PCMI (item (4) in Section 4.2 )

as well as molten fuel-coolant interaction (FCI). As described in RG 1.236, fuel cladding failure is presumed if the predicted fuel temperature anywhere in the pellet exceeds incipient fuel melting conditions. Section 4.5 describes the history and basis for this molten fuel failure threshold together, as well as the criteria and basis regarding FCI and coolability, because fuel melting is also addressed in the core coolability guidance of RG 1.236.

4.5 Allowable Limits on Damaged Core Coolability 4.5.1 Regulatory Guide 1.236 Core Coolability Criteria Limiting peak radial average fuel enthalpy to prevent catastrophic fuel rod failure and avoiding molten FCI is an acceptable metric to demonstrate that there is limited damage to core geometry and that the core remains amenable to cooling. The following restrictions should be met:

4-41 Peak radial average fuel enthalpy should remain below 230 cal/g.

A limited amount of fuel melting is acceptable, provided that it is less than 10 percent of fuel volume. If fuel melting occurs, the peak fuel temperature in the outer 90 percent of the fuel volume should remain below incipient fuel melting conditions.

For fresh and low burnup fuel rods, the peak radial average fuel enthalpy restriction will likely be more limiting than the limited fuel melt restriction. However, because of the effects of edge-peaked pellet radial power distribution and lower solidus temperature, medium to high burnup fuel rods are more likely to experience fuel melting in the pellet periphery under prompt power excursion conditions. For these medium to high burnup rods, fuel melting outside the centerline region should be precluded, and this restriction will likely be more limiting than the peak radial average fuel enthalpy restriction.

4.5.2 Molten Fuel and Core Coolability Technical Basis The molten fuel cladding failure threshold described in Section 4.4 and the coolability criteria described Section 4.5.1 remain unchanged from the guidance established in the 2015 NRC RIA memorandum (Ref. 11). This replaced the guidance established in the 2007 NRC RIA memorandum (Ref. 10) that was reproduced in the interim RIA criteria in Appendix B to SRP Section 4.2, Revision 3. The 2007 interim guidance is reproduced below:

Coolability Criteria:

Fuel rod thermal-mechanical calculations, employed to demonstrate compliance with criteria #1 and #2 below, must be based upon design-specific information accounting for manufacturing tolerances and modeling uncertainties using NRC-approved methods, including burnup-enhanced effects on pellet power distribution, fuel thermal conductivity, and fuel melting temperature.

1.

Peak radial average fuel enthalpy must remain below 230 cal/g.

2.

Peak fuel temperature must remain below incipient fuel melting conditions.

3.

Mechanical energy generated as a result of (1) nonmolten fuel-to-coolant interaction and (2) fuel rod burst must be addressed with respect to reactor pressure boundary, reactor internals, and fuel assembly structural integrity.

4.

No loss of coolable geometry due to (1) fuel pellet and cladding fragmentation and dispersal and (2) fuel rod ballooning.

As shown in Figure 28 of the NEA RIA report (Ref. 15), the maximum allowable peak fuel enthalpy to preclude fuel melting decreases with exposure and will likely become more limiting than the 230 cal/g coolability limit at approximately 30 GWd/MTU pellet burnup.

The above criteria were intended to maintain the fuel rod array and avoid the energetic reaction associated with molten FCI. The no-melt criterion also precludes cladding failure due to molten fuel swelling PCMI.

4-42 Over the years, many PWRs have adopted refined acceptance criteria relative to those described in RG 1.77. The following text was derived from a typical Westinghouse plants UFSAR.

Average fuel pellet enthalpy at the hot spot is below 225 cal/g for unirradiated fuel and 200 cal/g for irradiated fuel.

Fuel melting will be limited to less than the innermost 10 percent of the fuel volume at the hot spot even if the average fuel pellet enthalpy is below the limits of criterion above.

The allowable hot spot fuel enthalpy for fresh fuel and decrease due to irradiation effects are similar to the interim guidance. However, unlike the 2007 interim guidance, which avoids fuel melting, the Westinghouse criteria allow a limited volume of molten fuel.

As shown below, regulatory guidance documents provide an acceptable method for addressing fuel melting during the postulated BWR CRD accident and PWR CRE accident. However, this guidance pertains only to the radiological source term and does not address the thermal-mechanical reaction of molten FCI (e.g., pressure surge):

BWR CRD accident, RG 1.183, Alternative Radiological Source Terms for Evaluating Design Basis Accidents at Nuclear Power Reactors, Revision 1, Appendix C (Ref. 4):

The release attributed to fuel melting is based on the fraction of the fuel that reaches or exceeds the initiation temperature for fuel melting and on the assumption that 100 percent of the noble gases and 50 percent of the iodines contained in that fraction are released to the reactor coolant.

PWR CRE, RG 1.183, Revision 1, Appendix H (Ref. 4): The release attributed to fuel melting is based on the fraction of the fuel that reaches or exceeds the initiation temperature for fuel melting and the assumption that 100 percent of the noble gases and 25 percent of the iodines contained in that fraction are available for release from containment. For the secondary system release pathway, 100 percent of the noble gases and 50 percent of the iodines in that fraction are released to the reactor coolant.

Employing the above radiological guidance, several licensees have presumed a small fraction of molten fuel within their onsite and offsite dose calculations. In 2015, a survey of UFSARs revealed that, at the time, 33 of 35 BWRs and 32 of 65 PWRs included a small fraction of molten fuel in their dose calculations. For example, several licensees assume 0.25 percent of the core inventory is molten fuel (based on 10 percent pellet volume within 50 percent of axial height of 5 percent of the fuel rods). Fuel rods with a fuel temperature above incipient melting conditions are assumed to experience cladding failure. The licensee combines this melt source term with the nonmelt source term from fuel rods predicted (or assumed) to experience DNB to achieve the total RCS activity level.

Recognizing past precedence, the NRC staff proposed a revision to the 2007 interim criteria and guidance in 2015 and adopted it into RG 1.236. As described above, the no-fuel-melting criterion was intended to avoid molten FCI. A limited amount of fuel melting is permissible, provided the applicant demonstrates that this intent is satisfied. Limiting fuel melting to (1) the fuel centerline region and (2) a small fraction of the pellet volume (i.e., 10 percent) has been judged acceptable in past applications.

4-43 The above discussion addresses molten fuel guidance (i.e., the molten fuel cladding failure threshold and the allowable amount of melted fuel in a pellet). Concerning coolability, historically, limiting peak radial average fuel enthalpy and avoiding molten FCI were acceptable metrics to demonstrate coolable core geometry. Past applications did not specifically address items (3) and (4) above. Regulatory guidance does not include an acceptable method for predicting the population of burst rods and the amount of fuel fragmentation and dispersal, or for converting the thermal-to-mechanical energy.

Section 7.2.2 of the NEA RIA report (Ref. 15) describes the empirical database associated with fuel dispersal and FCI in RIA simulation tests. The NEA RIA report concludes that fuel dispersal occurs in connection with PCMI-type cladding failure and that balloon/burst failure does not lead to significant fuel dispersal. In contrast, recent LOCA simulation tests have exhibited fuel fragmentation and dispersal under balloon/burst conditions (e.g., as documented in NRC Research Information Letter 2021-13, Interpretation of Research on Fuel Fragmentation, Relocation, and Dispersal at High Burnup, issued December 2021 (Ref. 33)).

During PWR at-power scenarios with relatively low ejected rod worth, rod failure predictions will be more dominated by DNB than PCMI. In these scenarios, fuel rods operating with elevated rod internal pressure may balloon and burst. Since RIAs may experience multiple failure modes, and each failure mode may have its own sensitivities with respect to fuel fragmentation and dispersal, more work is needed to develop guidance for addressing these complex phenomena.

Figure 4-29 illustrates reported fuel dispersal as a function of local burnup and peak fuel enthalpy during prompt power excursion tests performed at PBF, BIGR, IGR, CABRI, and NSRR. The solid blue diamonds represent specimens that exhibited an unacceptable level of fuel dispersal (i.e., greater than10 percent). The data suggest that higher burnup fuel rods are more susceptible to fuel dispersal. However, until regulatory guidance exists to address items (3) and (4) above, applicants need only demonstrate compliance with coolability criteria 1 and 2.

Items (3) and (4) were removed in the interim guidance established in 2015 and are not present in RG 1.236.

4-44 Figure 4-29 Reported Fuel Dispersal During Prompt Power Excursions 4.6 Allowable Limits on Radiological Consequences RG 1.236 does not include accident dose radiological consequences criteria for CRD and CRE accidents, so this document will not discuss such limits. Previously, RG 1.77 included guidance for evaluating radiological consequences of PWR CRE accidents, but the updated criteria were not included in RG 1.236, as the NRC has included them in Revision 1 of RG 1.183 (Ref. 4).

4.7 Allowable Limits on Reactor Coolant System Pressure For new license applications, the maximum RCS pressure should be limited to the value that will prevent stresses from exceeding emergency condition (Service Level C), as defined in section III of the ASME Boiler and Pressure Vessel Code. For existing plants, the allowable limits for the reactor pressure boundary specified in the plants UFSAR should be maintained.

4.8 Transient Fission Gas Release For non-LOCA DBAs involving a rapid increase in fuel rod power, such as the BWR CRD accident and PWR CRE accident, additional fission product releases may occur due to pellet fracturing and grain boundary separation. This transient FGR increases the amount of activity available for release into the RCS for fuel rods that experience cladding breach. In other words, the total fission-product source term during an RIA includes both (1) the steady-state gap inventory present from normal operation before the RIA and (2) any transient fission gas released during the event.

0 50 100 150 200 250 300 0

10 20 30 40 50 60 70 80 Peak Fuel Enthalpy (cal/g)

Burnup (GWd/MTU)

Open Diamond < 10% Fuel Loss Blue Diamond > 10% Fuel Loss Red Diamond - Loss not Quantified Green Diamond - End Weld Failure Open Circles - No Fuel Loss

4-45 Fission gas released during the experiments was measured for many of the prompt power excursion tests that comprise the RIA empirical database. This transient FGR database was originally documented in PNNL Report PNNL-18212, Revision 1, Update of Gap Release Fractions for Non-LOCA Events Utilizing the Revised ANS 5.4 Standard, issued June 2011 (Ref. 34). The database was revised, expanded, and captured in Appendix B to RG 1.236 and the 2015 technical bases document (Ref. 11).

Figure 4-6 through Figure 4-9 capture the compiled empirical database of fission gas released during prompt-power testing. Examination of the database reveals a prominent distinction between high burnup (greater than 50 GWd/MTU) and lower burnup data. Higher burnup rods tend to exhibit higher transient FGR (percent). The differences in transient FGR for high and low burnup rods can be attributed to the mechanism of FGR. Whereas FGR (into the rod plenum) during normal operation is governed by diffusion, pellet fracturing and grain boundary separation are the primary mechanisms for FGR during an RIA. Hence, the amount of release is dependent on local burnup (fission gas accumulation along grain boundaries and within the porous rim region) and a local power increase. A combination of porous rim region with a high concentration of fission gas bubbles and edge peaked power excursion would tend to maximize FGR in high burnup fuel.

Based on this database, the following transient FGR correlations were derived. Figure 4-30 illustrates these correlations in the context of the empirical database.

pellet burnup < 50 GWd/MTU TFGR = maximum [ (0.26

• H) - 13) / 100, 0 ]

pellet burnup > 50 GWd/MTU TFGR = maximum [ (0.26

• H) - 5) / 100, 0 ]

where:

TFGR = transient fission gas release fraction H = increase in radial average fuel enthalpy, calories per gram

4-46 Figure 4-30 Transient Fission Gas Release Correlations PNNL-18212, Revision 1, investigated the effect of differences in diffusion coefficients and radioactive decay on fission product transient release. This investigation concluded that adjustments to the above empirically based correlations are needed for different radionuclides.

For stable, long-lived noble gases (e.g., krypton (Kr)-85) and alkali metals (e.g., cesium-137),

the transient fission product release is equivalent to the above burnup-dependent correlations.

For volatile, short-lived radioactive isotopes such as halogens (e.g., iodine (I)-131, I-132, I-133, I-135) and xenon (Xe) and krypton noble gases except Kr-85 (e.g., Xe-133, Xe-135, Kr-85m, Kr-87, Kr-88), the transient fission product release correlations should be multiplied by a factor of 0.333.

Due to the lack of data, it is difficult and likely overly conservative to attempt a 95/95 upper tolerance correlation. Instead, the staff elected to capture the leading edge of the two burnup intervals (greater than 50, less than 50 GWd/MTU). While a few data points exceed the correlations, the NRC staff felt that, given the spread in the data, the proposed, burnup-dependent correlations provided a reasonable level of conservatism.

These transient FGR correlations are presented in Appendix B to RG 1.236, as well as in Revision 1 of RG 1.183 (Ref. 4).

0 5

10 15 20 25 30 35 40 0

50 100 150 200 Measured Transient FGR (%)

Peak Enthalpy Increase (cal/g)

CABRI NSRR PWR NSRR BWR BIGR VVER IGR VVER CABRI MOX NSRR MOX Test Segment Exposure:

Purple: > 70 GWd/MTU Orange:60-69.9 GWd/MTU Red:

50 - 59.9 GWd/MTU Blue:

30 - 49.9 GWd/MTU Green: 10 - 29.9 GWd/MTU BU < 50 GWd/MTU: Transient FGR (%) = [(0.26

• H) - 13]

BU 50 GWd/MTU: Transient FGR (%) = [(0.26

• H) - 5]

Open Symbol = Non-failed Closed Symbol = Failed

5-1 5 CONCLUSIONS The purpose of this report is to examine the technical and regulatory bases for the acceptance criteria and guidance for RIAs described in RG 1.236. Additionally, this report aims to document the history and evolution of NRC RIA acceptance criteria and guidance. RIAs consist of postulated accidents that involve a rapid insertion of positive reactivity. These accident scenarios include CRE in PWRs and CRD for BWRs.

RG 1.236 is the second RG published by the NRC that addresses RIAs. It is a cumulation of significant research and the expansion of knowledge since the first, RG 1.77, was published in 1974. Research has shown that the legacy RG 1.77 is not adequate for ensuring that the number of failed fuel rods due to an RIA is not underestimated. The NRC provided interim RIA acceptance criteria and guidance in 2007 and revised them in 2015 in NRC memoranda before it published RG 1.236 in 2020. RG 1.236 takes into consideration modern and updated RIA test data published as of 2020.

RG 1.236 formalizes many of the guidance and criteria issued as interim guidance by the NRC in 2015 and updates others. The area that had the most significant changes was the hydrogen enhanced PCMI failure limit curves. RG 1.236 updated the PCMI failure curves primarily based on more recent data and industry comments on the draft of RG 1.236, DG-1327. Specific changes included modification of the functional form of the PCMI failure curves (i.e., all PCMI curves were updated to an exponential function) and the addition of failure thresholds for RXA cladding without a liner (i.e., nonbarrier RXA cladding).

In the development of the guidance in RG 1.236, consideration of regulatory stability was accounted for. For example, when developing the PCMI failure curves, a lower bound fit was employed rather than a best estimate fit, to ensure that the limits do not have to be repeatedly reevaluated or are invalidated by new RIA test results.

In summary, RG 1.236 is a state-of-the-art regulatory document for addressing RIAs, considering the various fuel failure modes during such events, and is the result of substantial research and several iterations of interim guidance.

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6-4

NUREG/KM-0019 Paul Clifford (retired)

Joseph Messina Division of Safety Systems Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, DC 20555-0001 Joseph Messina Same as above The purpose of this report is to document the technical and regulatory bases for the acceptance criteria and guidance provided in Regulatory Guide (RG) 1.236, Pressurized-Water Reactor Control Rod Ejection and Boiling-Water Reactor Control Rod Drop Accidents. The pressurized-water reactor control rod ejection and boiling-water reactor control rod drop accidents are low-probability events involving a sudden and rapid insertion of positive reactivity. These postulated accidents are safety significant because of their potential ability to cause significant core damage and challenge the integrity of the reactor coolant pressure boundary. This report identifies regulatory requirements, details the existing empirical database of fuel performance under rapid power excursion testing, and documents the technical bases of the analytical limits and guidance that satisfy the regulations.

Reactivity Initiated Accident Control Rod Drop Accident Control Rod Ejection Accident Pellet-Cladding Mechanical Interaction June 2024 Technical Technical and Regulatory Bases for Regulatory Guide 1.236, PWR Control Rod Drop and BWR Control Rod Ejection Accidents

NUREG/KM-0019 Technical and Regulatory Bases for Regulatory Guide 1.236, PWR Control Rod Drop and BWR Control Rod Ejection Accidents June 2024