Text

Forwards Evaluations to Document Completion of Confirmatory Actions Required to Support 35% Power Restriction During Environ Qualification Schedule Extension Period
	ML20138B512
Person / Time
Site:	Fort Saint Vrain
Issue date:	12/10/1985
From:	Walker R; PUBLIC SERVICE CO. OF COLORADO
To:	Berkow H; Office of Nuclear Reactor Regulation
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ML20138B522	List: ML20138B512; ML20138L407;
References
	P-85460, TAC-59787, NUDOCS 8512120319
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PUBLIG SERVIGE GOMPANY OF GOLORADO P. O. BOX 840 + DENVER, COLORADO 8020t R. F. WALKER December 10, 1985 Fort St. Vrain Unit No. 1 P-85460 Director of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, DC 20555 Attention: Mr. H.N. Berkow, Project Director Special and Standardization Project Directorate Docket No. 50-267

SUBJECT:

Confirmatory Actions in Support of 35 Percent Power Restriction During EQ Schedule Extension Period

REFERENCE:

1) PSC Letter dated 09/24/85, Walker to Chilk (P-85334)

2) PSC Letter dated 11/22/85, Lee to Chilk (P-85432)

Dear Mr. Berkow:

Reference 1) submitted the Public Service Company of Colorado's (PSC) request for a schedule extension for the Fort St. Vrain (FSV)

Environmental Qualification (EQ) Program under the provisions of paragraph (g) of 10CFR50.49. Reference 2) withdrew the schedule extension request in Reference 1) and submitted a revised schedule extension request based on restricting Fort St. Vrain operation to 35 percent of rated power during a schedule extension period which would expire May 31, 1986. Attachment 1 to Reference 2) contained a list of confirmatory actions for PSC to complete and document in support of the 35 percent power level restriction during the schedule extension period.

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, Attachment I to this letter contains the PSC evaluations to document the completion of the confirmatory actions required in Attachment 1 to Reference 2). The remaining attachments to 1,his letter contain supplemental information in support of interim plant operation with the 35 percent power level restriction.

The proposed interim operational restriction has been reviewed by the Fort St. Vrain Plant Operations Review Comittee and the Nuclear Facility Safety Comittee. These comittees have determined that interim operation with the 35 percent power restriction should be 3

reviewed by the NRC since a postulated high energy line break during this period could result in offsite radiological consequences, whereas the high energy line break accident previously evaluated in the Fort St. Vrain FSAR had no offsite radiological consequences.

Both comittees have concluded that the revised high energy line break accident scenario and postulated radiological consequences de not constitute an undue risk to the health and safety of the public.

Based on the justification submitted in Reference 2) and on the information contained in the attachments to this letter, NRC approval

to resume operation of Fort St. Vrain at power levels not to exceed 35 percent of rated thermal power during a schedule extension period which would expire May 31, 1986, is hereby requested.

If you have any questions concerning PSC's request or the attached information, please contact Mr. M.H. Holmes at (303) 480-6960.

Very truly yours, k.T. W O l R.F. Walker President RFW/MHH:jmt Attachments 4 cc: Director, IE, NRC

P-85460 4

LIST OF ATTACHMENTS AND

SUMMARY

DESCRIPTION

SUMMARY

DESCRIPTION FOR THE INTERIM MODE OF OPERATION AT 35 PERCENT POWER ATTACHMENT 1 - PSC CONFIRMATORY EVALUATIONS TO SUPPORT 35 PERCENT POWER OPERATION ATTACHMENT 2 - PSC RESPONSES TO NRC ADDITIONAL INFORMATION REQUEST ATTACHMENT 3 - FSV FUEL PERFORMANCE UNDER HIGH TEMPERATURE CONDITIONS ATTACHMENT 4 - ENGINEERING EVALUATION OF LINER COOLING WITH FIREWATER FOLLOWING A HIGH ENERGY LINE BREAK FROM 35 PERCENT POWER OPERATION - EE-EQ-0019 ATTACHMENT 5 - SIGNIFICANT HAZARDS CONSIDERATION FOR OPERATION AT 35 PERCENT POWER ATTACHMENT 6 - GA LETTER GP-2656, " FIRE WATER LINER COOLING DURING DESIGN BASIS ACCIDENT N0. 1", DATED OCTOBER 18, 1985 4

P-85460

SUMMARY

DESCRIPTION FOR THE INTERM MODE OF OPERATION AT 35 PERCENT POWER The following is a sumary of PSC's intentions with regard to the mode of operation of Fort St. Vrain (FSV) at 35 percent power until May 31, 1986. Supporting analyses and further details can be found in the attachments to this letter.

In the event. of a high energy line break (HELB) from 35 percent power, shutdown and cooling of the reactor will be accomplished by supplying fire water to the liner cooling system. For the worst case sc:nario, it has been shown by analysis that the PCRV does not have to be depressurized and liner cooling must be initiated within 29.4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months after the line break.

However, it is PSC's intent to comply with the FSV Technical Sp:cifications. Thus PCRV depressurization will be initiated at 2.3

_h:urs after the HELB if access to the bull' ding can be accomplished.

If this is not successful, PCRV depressurization will be initiated as soon as possible. By analysis, up to 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months is permitted to initiate PCRV depressurization. If PCRV depressurization initiation is accomplished, liner cooling will be initiated as soon afterwards as feasible. If depressurization is not initiated by 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months then tha PCRV must remain pressurized with liner cooling initiated before 29.4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months .

It has also been shown that actuation of the reserve shutdown (RSD) system is'not necessary for the interim operation HELB scenario.

However, to comply with Technical Specifications, manual actuation of the RSD system will be performed in conjunction with the manual initiation of depressurization and liner cooling.

The only other manual action that must be performed is manual actuation of the circulator brake and seal system if it fails to automatically actuate due to the HELB harsh environment. This will be performed as soon as possible, with I hour being the conservative time frame used in the analysis.

1 Fire water can also be supplied to cool the fuel storage facility.

The current heat load is only 4 percent of the design heat load.

Under current Technical Specifications up to 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months of no cooling is permitted before the fuel must be removed. In this time frame adequate _ cooling can be made available to the fuel storage facility.

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1 ATTACHMENT 1 PSC CONFIRMATORY EVALUATIONS TO SUPPORT 35 PERCENT POWER OPERATION k

i s Attachment 1 to P-85460 CONFIRMATORY ACTION 1: Complete an evaluation which confirms that PCRV liner cooling using fire water can be utilized to prevent significant damage to any of the fission product. barriers, including fuel particle coatings, in the event of a high energy line break at power levels up to the 35 percent power restriction.

PSC EVALUATION: To fully cover the above confirmatory action the following range of scenarios for the 35 percent power level were analyzed using the RECA code:

Case 1. The PCRV is manually depressurized beginning at 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months following a high energy line break. Liner cooling and the Reserve Shutdown System are initiated manually at 5 hours5.787037e-5 days 0.00139 hours 8.267196e-6 weeks 1.9025e-6 months . This is the preferred mode of operation which is in accordance with the FSV Technical Specifications (see Evaluation 8).

This is dependent upon personnel access to the reactor and turbine buildings in one hour following the event which created the harsh environment. The computer code, CONTEMPT-G, used to calculate the the temperature high energy profiles line following break (see Evaluation 3) conservatively estimates an access time of 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months with cool suits and Scott Air Paks.

PSC believes that access will be possible within one hour based on past utility experience with steam line breaks, the possibility of opening large bay doors and other openings (not modeled in CONTEMPT-G) and the much lower ambient outside air temperatures than the i

$ s Attachment 1 to P-85460 80 degrees Fahrenheit used by CONTEMPT-G.

However, since one hour access can not be demonstrated analytically for all cases, other scenarios were considered as noted below:

Case 2. The PCRV is manually depressurized beginning at 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months following a high energy line break. The 12 hour1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months time frame was chosen since this is the latest time at which depressurization can be initiated (see Evaluation 4). Liner cooling is initiated at 14 hours1.62037e-4 days 0.00389 hours 2.314815e-5 weeks 5.327e-6 months . Although this case does not meet the Technical Specification time limit of 2 1/4 hours (basis for LC0 4.2.18) to initiate depressurization for 35 percent power it does meet the basic intent of the specification to protect the helium purification train front end components (see Evaluation 4). It is in conformance with PSC's objective to depressurize, if possible, following LOFC from 35 percent power.

Case 3. The core remains pressurized and liner cooling is initiated at 29.4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months . This is the time when the top head cover plate reaches 1500 degrees Fahrenheit and is conservatively assumed to fail, taking with it all of the liner insulation. This is the worst case for liner temperatures and establishes an upper bound for the top head liner and adjacent concrete temperatures. It is also

. s Attachment 1 to P-85460 the situation which leads to the maximum liner heat flux and therefore poses the most stringent conditions on the liner cooling system with fire water.

Case 4. The core remains pressurized and liner cooling is initiated at 14 hours1.62037e-4 days 0.00389 hours 2.314815e-5 weeks 5.327e-6 months . FSAR Section D.2.1 concludes that the 1500 degrees Fahrenheit failure point assumed for the top head cover plate and Kaowool insulation is conservative.

This case was analyzed to determine

the maximum cover plate temperature without an arbitrary failure point being specified. The result is that the cover plate reaches a maximum temperature of about 1600 degrees Fahrenheit.

4 A summary of the results of the above calculations is presented in Table 1-1.

Although the depressurized cases result in peak fuel temperatures in excess of the 2900 degrees Fahrenheit fuel temperature limit, it is important to note that the 2900 degrees Fahrenheit value is a conservative lower limit for fuel failure. The actual temperature at which fuel failure takes place is several hundred degrees higher (see Attachment 3).

During the fuel manufacturing process the coated fuel particles within the rod are exposed to temperatures in the range 2900 degrees to 3340 degrees Fahrenheit for a period of one to two hours. In addition, tests have been performed which demonstrate that the 2900 degrees Fahrenheit limit is an unrealistically low value. The results of these tests were documented in an attachment

i s Attachment 1 to P-85460 1979 (P-79157),

to a letter Swart dated(NRC to Speis July).24, This attachment is included as Attachment 3 of this letter.

Also see FSAR Table A.2-9 for further fuel particle coating failure data as a function of higher temperatures.

As stated previously, the objective is to depressurize the PCRV following a- postulated -

steam line break from 35 percent power which results in an LOFC. Although the depressurized condition does result in higher peak fuel temperatures (still, below predicted fuel particle coating failure temperatures) versus the pressurized condition, the depressurized case is favored because of the following advantages:

a) The driving force for leakage through the PCRV is reduced.

b) There is little or no predicted damage to the PCRV internals (see Evaluation 9).

c) The 30 day radiation dose at the low population zone boundary is reduced due to filtering of the circulating activity in the primary coolant by the helium purification system.

A significant factor in the 35 percent power case is that, for both pressurized and depressurized situations, if the equipment affected by the harsh environment can be repaired then forced circulation can be restarted at any time. In the 100 percent power LOFC case after five (5) hours forced

, circulation cannot be restarted. This is due to the upper plenum gas temperature exceeding

, s Attachment 1 to P-85460 a value of 2100 degrees Fahrenheit during the 100 percent power situation discussed in FSAR Section D.2.5. Initiation of forced circulation with gas temperatures exceeding 2100 degrees Fahrenheit could cause damage to the steam generator inlet ducts. In the 35 percent power case the upper or lower plenum gas temperatures never reach the 2100 degrees Fahrenheit temperature limit.

The conclusion of the above analysis is that PCRV liner cooling can be utilized to prevent significant damage to any of the fission

product barriers, including fuel particle coatings, in the event of a high energy line break from 35 percent power. This is true whether the PCRV is depressurized in the initial stages of the accident, or is maintained pressurized throughout the course of the accident.

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Attachment 1 to P-85460 TABLE 1-1

SUMMARY

OF FUEL & LINER TEMPERATURES FOR THE FOUR CASE STUDIES MAX PEAK FUEL PEAK AVG. FUEL TEMP. & TIME TEMP. & TIME t

TO REACH TEMP. PERCENT TO REACH TEMP.

OF FUEL t TEMP AB0VE TEMP (DEGREES TIME 2900 (DEGREES TIME CASE F) (HOURS) DEGREES F) (HOURS) 1 1 3140 107 5 1958 157 i

2 3160 125 5 2005 250 3 2150 29.4 0 1450 29.4 4 2350 150 0 1750 200 TIME MAX LINER TEMP. MAX IN CAL-

& TIME AT WHICH DEPTH OF CULATION TIME THIS IS ATTAINED CONCRETE AT WHICH '

LINER AB0VE LINER COVER TEMP 400 COOLING PLATE (DEGREES TIME DEGREES WAS IN-CASE FAILS F) (HOURS) (INCHES) ITIATED .

1 NEVER 340 5 0 5 hrs 2 NEVER 570 14 APPR0X. 14 hrs 4"

3 29.4 hrs 900 29.4 APPR0X. 29.4 hrs 1500 14" Degree F 1

4 ASSUMES 570 14 APPROX. 14 hrs N0 4" FAILURE

Attachment 1 to P-85460 CONFIRMATORY ACTION 2: Evaluate the leak tightness and structural integrity of the PCRV during the heatup which would occur following an extended loss of forced circulation cooling resulting from a high energy line break from 35 percent of rated power. Consider the cold reheat helium interspace leak, PCRV penetrations and seals, and other portions of the PCRV where leakage may be a concern. Actual PCRV leakage experience should be considered (e.g., the recent LER on the PCRV penetration cold reheat helium leak).

PSC EVALUATION: Leak Tightness PCRV penetration closures are designed to maintain the leak tight integrity of the PCRV during all normal operating conditions.

Each PCRV penetration includes at least two independent closures in series to provide primary and secondary containment.

The interspace between the primary and secondary closure of each penetration is normally pressurized with purified helium to a pressure slightly higher- than PCRV pressure, except for steam generator penetrations under circumstances which are discussed below. Therefore, leakage through either a primary or secondary closure is normally purified helium.

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l b Attachment 1 i to P-85460 The closures are designed and tested to assure extremely low leakage rates.

The expected initial total leakage from the PCRV, as described in FSAR Section 5.8.2.3, was:

LEAK RATE AT 700 PSI LEAK SOURCE PRESSURE DIFFERENTIAL TOTAL LEAKAGE Primary Closures 0.1 cc/sec 1.2 lb/yr e Secondary Closures 0.5 cc/sec 6.2 lb/yr An initial combined pressure test and leak test of the PCRV was performed in 1971. For the purposes of this -test, the allowable integrated leak rate was established as 14.4 percent per year of the design primary coolant inventory, by weight, which is based upon an annual release limit, including design release rate from the gas waste system, of one-tenth of 10CFR20 limits.

The test confirmed that the integrated PCRV leak rate was less than 14.4 percent per year (Reference 1).

Surveillance testing of PCRV primary and secondary closure leakage is performed once each quarter or as soon as practicable after an increase in pressurization gas flow is alarmed. The latest surveillance testing was performed on November 7-8, 1985 per SR 5.2.16.a-Q. The results of that testing indicated the following leakage rates:

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Attachment 1

, to P-85460 PRIMARY SECONDARY CLOSURE CLOSURE

? GROUP NO. PENETRATIONS LEAKAGE LEAKAGE

, I All penetrations in the 0 0 top head of the PCRV (37-control rod drive, 2-HTFA, and 1-top access)

II All 4nstrument penetra- 0 0

' tions (20), plus the

".3 gottomaccesspenetration III Total cf 6-Loop I Steam 0 0

. Generator Penetrations IV Total of 6-Loop II Steam 1.3 lb/hr*

, -1 Generator Penetrations V Helium Cir,culator 0 0 VI Hel,ium Circulator 0 0 VII Helium. Circulator 1.3 lb/hr 1.62xE-5 lb/hr VIII Helium Circulator 0 0 There are no previsions for measuring Group III or IV

- s secondary seal leakage, so only the total primary plus

_ ' secondary leakage plus leakage to the cold reheat (if any) can be_ determined. This leakage is assumed to be

. M61d rehest leakage based on~ the absence of activity in the interspace the presence of helium in the main condensor.and the leak tightness of all other secondary closures.

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Attachment 1 to P-85460 During operation, PCRV closure leakage is controlled by two Technical Specifications; LC0 4.2.7 - PCRV Pressurization, Limiting Conditions for Operation, and LC0 4.2.9 -

PCRV Closure Leakage, Limiting Conditions for Operation. LC0 4.2.7 requires that all interspaces between the primary and secondary closures (except for S/G penetrations as discussed below) be maintained at a pressure greater than primary system pressure with purified helium gas. This assures that any leakage across the primary closure is flow of purified helium into the PCRV and any leakage across the secondary closure is flow of purified helium into the reactor building.

It also minimizes the differential pressure across the primary closure, thereby limiting the flow of helium through any leak path that might be present in the primary closure. LCO 4.2.9 prohibits operation if helium leakage through any group of primary closure seals exceeds 400 lbs/ day at a differential pressure of 10 psi or if total leakage through all the secondary closure seals exceeds 400 lbs/ day at a differential pressure of 688 psi. These' limits provide g assurance that the maximum leak rate that

- would occur upon failure of either closure would not result in a release in excess of that evaluated 'for the Maximum Credible Accident (MCA). The maximum resultant dose rate, assuming failure of W. secondary closure and design primary cot.lant activity, is at least an order of magnitude less than 10CFR100 the exclusion area boundary EAB).

(guidelines _ at Steam Generator Penetration Interspace Helium Leakage to the Cold Reheat Steam System In 1980, a leak path developed between some of the steam generator penetration

Attachment 1 to P-85460 interspaces ,and the cold reheat steam piping internal to the penetrations,. allowing aurified helium to leak into the cold reheat leader. The maximum purified helium leak rate, with the interspace pressurized to above primary coolant pressure, was re to be 1312 pounds per day (Reference.2) atported 57 percent power. Due to the absence of primary closure leakage and the extreme difficulty of repair, the interspace pressurization system was modified and the Technical Specifications were revised to permit' the affected steam generator penetrations to be pressurized below primary coolant pressure but above cold reheat pressure. This modification reduces purified helium leakage to the reheat system, but also creates a potential flow path for primary, coolant to the cold reheat system.

When operating in! this mode, the Technical Specification requires monitoring of the interspaces. for primary coolant activity and prohibits' operation if the primary coolant leakage via the cold reheat system exceeds 1.4 curies per day, an amount which represents 10 percent of the plant's design objective for radioactive gas releases. At design circulating activity, the 1.4 curie per day limit equates to 0.6 pounds per day primary coolant. leakage, with a differential pressure of at least 50 psi, a limit that is far more restrictive than the original limit of 400 pounds per' day (for each group of 6 penetrat. ions) at a differential pressure of 10 psi. The off-site doses from this effluent pathway are extremely small; at'the exclusion area boundary, the whole body gamma dose is 0.14 millirad, and the annual thyroid p dose is 5.2xE-6 millirads (Reference 3.)

The recent LEP. d'ealing with steam generator interspaces with colc; reheat leaks (LER 85-013) dealt with the failure to pressurize the

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Attachment 1 to P-85460 interspaces in question to a pressure above cold reheat pressure. This condition is prohibited at PCRV pressures above 100 psia and, therefore, it was reported to the NRC in compliance with 10CFR50.72 and 10CFR73(a)(2)(1)(b). The leakage that occurred was that of steam from the cold reheat pipe into the penetration interspace.

No primary coolant leakage was detected.

Although activity monitor 2264 became inoperable due to water entrained in the sample line, primary coolant leakage would have been carried with the reheat steam to the air ejector activity monitor. The air

, ejector activity monitor has a lower level of detectability, at current activity levels, of 1.4 pounds per day. It did not detect any abnormal conditions during the period in question. In the event that a primary coolant leak were present, it would have pressurized the interspace with primary coolant and been limited by Technical Specification LC0 4.2.9 to 1.4 curies per day.

Accident Consequences Normal PCRV penetration interspace pressurization may not be maintained during the heatup which would occur following an extended loss of forced circulation cooling resulting from a high energy line break from 35 percent rated power. The supply of purified helium for pressurization is assumed l to be interrupted and the interspaces would I eventually bleed down to an -equilibrium pressure that would depend on the prevailing primary and/or secondary closure leaks and/or pressurization system leakage, if any. The normal differential pressures across the closures would not be maintained. The differential pressure across the primary 1

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Attachment 1 to P-85460 closures would change from as little as 10 psi to as much as 688 psi while the

differential pressure across the secondary closure could change from as much as 700 psi to as little as zero psi. Furthermore, the cold reheat pipe would depressurize to atmospheric pressure via the main condenser or the high energy line break. The table below summarizes the range of pressures and leak rates that could prevail

. s Attachment 1 to P-85460 TABLE 2-1 NORMAL CONDITIONS ALLOWABLE DIFFERENTIAL ACTUAL (1) MAXIMUM (2)

LEAK PATH PRESSURE LEAK RATE LEAK RATE Primary Closures

- All except S/G 10 psi 31.2 lb/ day 400 lb/ day penetrations with (Group VII) per Group CRH leak

- S/G penetrations with CRH leak

- at actual circu- 350 psi (3) 60 lb/ day lating activity maximum total

- at design circu- 350 psi (3) 0.6 lb/ day lating activity maximum total Secondary Closures 688 psi 3.89xE-4 400 lb/ day lb/ day total Cold Reheat Pipe 15 psi 31.2 lb/ day 700 lb/ day in Loop with Cold (4)

Reheat leak

Attachment 1 4

to P-85460

! TABLE 2-1 (continued)

ACCIDENT CONDITIONS LEAK RATE

' BASED ON ALLOWABLE DIFFERENTIAL ACTUAL MAXIMUM LEAK PATH PRESSURE LEAK RATE LEAK RATE Primary Closures

- All except S/G .

688 psi 10.8 lb/hr 1145 lb/hr penetrations with maximum (Group VII) per Group CRH leak

- S/G penetrations with CRH leak

- at actual circu- 688 psi 0 <100 lb/ day lating activity maximum total

- at design circu- 688 psi 0 <1.0 lb/ day lating activity maximum total i Secondary Closures 688 psi 3.89xE-4 400 lb/ day maximum ~lb/ day total Cold Reheat Pipe 688 psi 211 lb/ day 4740 lb/ day maximum T

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Attachment 1 to P-85460 NOTES TO TABLE Notes: (1) Reference FSV Surveillance Test Procedure SR 5.2.16a-Q, PCRV Closure Leakage Determination, performed November 7-8, 1985 (2) Maximum All'owable Leak Rate permitted by Technical Specifications, LC0 4.2.9 - PCRV Closure Leakage (3) S/G penetration primary closure leakage cannot be measured directly or indirectly. Leak tightness is infer, red from absence of activity in the interspace.

(4) Measured leakage in Group III is attributed to the CRH piping internal to the penetration. This is purified helium leakage at this time.

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Attachment 1 to P-85460 In addition to penetration closure leakage, there is only one other potential release path during the postulated event. That is leakage via the four circulator shafts due to failure of the static seals to automatically or remote-manually actuate during the first hour of the cooldown. This assumption leads to the conclusion that approximately 1250 pounds of primary coolant is released into the reactor building in the first hour of the event via the circulator shafts. This assumes worst case conditions with no bearing water or seal system, and with the seals set manually in one hour.

Based on the foregoing evaluation of actual and potential leak rates and consideration of the combinations of leak paths necessary to create a release of primary coolant, the worst case PCRV leakage can be determined.

That is, both primary and secondary closures must leak coincidentally or a primary closure and cold reheat pipe must leak coincidentally.

The hypothetical worst case primary coolant leakage consistent with Technical Specification limits and the circulator shaft leakage is 1250 lbs plus a continuing 401 lbs/ day (400 lbs/ day via secondary closure leakage plus I lb/ day via cold reheat). No credit is taken for primary closure leak tightness because of the relatively large leak rate allowed by Technical Specifications. It is highly unlikely that this leak rate would occur without other, non-mechanistic, failures because it would require coincidental maximum allowable leakage of independent, redundant pressure boundaries. There is no indication, based on experience to date, that such a set of circumstances would ever exist.

l Attachment 1 to P-85460 Nevertheless, assuming maximum design primary i coolant activity the off-site doses for this I hypothetical leak rate are shown in Table 2-2:

The measured primary coolant activity is presently approximately one percent of the maximum design activity allowed by Technical Specifications. If the expected primary coolant total activity value of about 300 curies rather than the maximum design of 30,000 curies was used to calculate the l

results in Table 2-2 then the doses would be about two orders of magnitude less.

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Attachment 1 to P-85460 t

TABLE 2-2 i

INTEGRATED DOSE fREM)

LOCATION TIME WHOLE BODY THYROID Exclusion 2 Hours 5.4 x E-3 2.3 x E-2 Area Boundary Low Popu- 30 Days 8.0 x E-4 5.9'x E-2 lation Zone Boundary 1

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Attachment 1 to P-85460 References

1. Barker, Rath, Tammadge,." Fort St.

Vrain Unit 1 PCRV Leak Test Report for Public Service Company of Colorado", GA-B10875

2. Reportable Occurrence 80-30, Issue 1, April 1, 1981, Report No. 50-267/80-30/03-X-1 Supplemental Report - Additional Information.

3. " Safety Evaluation by the Office of Nuclear reactor Regulation Related to Amendment No. 26 to Facility Operating License No. DPR-34",

Enclosure 2 to NRC letter G-82079 dated March 18, 1982 from G.

Kuzmycz to 0.R. Lee.

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Attachment 1 to P-85460 CONFIRMATORY ACTION 3: Provide high energy line break temperature profiles for accidents from reactor power levels up to 35 percent demonstrating access to plant areas where operators are required to take the necessary manual actions by specified times. Identify all assumptions made regarding operator response times to the high energy line break. Credit will be taken only for qualified systems. An evaluation will be made for the need for SLRDIS for 35 percent power operation.

PSC EVALUATION: Attached are 6 different building temperature profiles for 35 percent power. Each scenario is described below. Access to the buildings in reference to the resulting temperatures is addressed in Evaluation 12 and additional information on manual termination of a steam leak is addressed in Evaluation 13. Listed below are the assumptions used for all the scenarios:

- - For large and medium size line breaks, i.e., those corresponding to a leak rate greater than 40 lbs/sec or 18 percent of normal feedwater flow at 35 percent power, there is sufficient indication to ensure that the operator can take action after 10 minutes allowing one minute for each operator action required.

For smaller breaks, i.e. those with a leak rate smaller than 40 lbs/sec, the assumption is that the operator can initiate the required actions at 60 minutes.

Existing plant protective and control system actions act in their prescribed function for up to 20

Attachment 1

.- to P-85460 seconds after initiation of the event. After this time the conservative assumption is made that all components in these systems fail in the most deleterious positions (i.e., all valves in both the reactor and turbine buildings are assumed to fail in the open position regardless of. the harsh environment. The inventory of the entire secondary coolant system is thus dumped into the affected building).

No credit is taken for the existing SPRDS or the new SLRDIS.

Operation is at 35 percent power with the auxiliary boilers on line.

- No protection, control or power systems are environmentally qualified.

Reactor scram can occur either automatically or manually and the rods remain in the core.

No credit is taken for ventilation in the affected building.

The hot reheat line break was determined to be worst case in comparison to all other-line breaks.

The following is a description of each scenario presented (associated. temperature profiles are attached):

Attachment 1 to P-85460 HRH-9 Leak rate before manual actions equals 224 lbs/sec (normal feedwater flow at 35 percent power).

Turbine building . full offset rupture of the hot reheat line.

Reactor automatically scrams at 10 seconds due to low hot reheat pressure.

Auxiliary boiler is manually tripped at 11 minutes and is linearly ramped to zero flow in 1/2 hour for conservatism.

Electric boiler feed pump is manually tripped at 12 minutes.

Steam driven boiler feed pumps are assumed to run until the steam driving force is depleted.

The resulting turbine building temperature is 131 degrees Fahrenheit at 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

HRH-10 Reactor building full offset rupture of the hot reheat line.

Same operator actions and leak rate as HRH-9 The resulting building temperature is 121 degrees Fahrenheit at 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

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Attachment 1 to P-85460 HRH-13 Turbine building medium size break in the hot reheat line.

A rupture large enough to leak more than 140 lbs/sec will scram the reactor on low hot reheat pressure and will be enveloped by HRH-9.

Therefore, this scenario assumes a medium break that results in a 139 lbs/sec leak rate and will not cause a scram. The leak rate is still large enough for the operator to know right away a problem has occurred resulting in prompt action after 10 minutes.

The operator manually scrams the reactor at 11 minutes.

The auxiliary boilers are manually tripped at 14 minutes and the flow rate slows down to zero by 30 minutes.

The electric boiler feed pumps are manually tripped at 14 minutes.

The resulting turbine building temperature is 128 degrees Fahrenheit at 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

HRH-14 Reactor building medium break in the hot reheat line.

Same operator actions and leak rate as HRH-13. A leak rate greater than this medium break is enveloped by HRH-10.

Attachment 1 to P-85460 The resulting reactor building temperature is 118 degrees Fahrenheit at 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

HRH-15 Turbine building small break in the hot reheat line.

A rupture small enough to result in a 40 lbs/sec leak rate (approximately a 7" line) or less may not be instantly detected by the operator. Therefore, this scenario assumes a small break size that results in a leak rate of 40 lbs/see that is detected within 1 hour1.157407e-5 days 2.777778e-4 hours 1.653439e-6 weeks 3.805e-7 months . A rupture larger than 40 lbs/sec can be detected quicker and is enveloped by HRH-13.

The operator manually scrams the reactor at I hour.

The auxiliary boilers are manually tripped at 64 minutes and the secondary coolant system blows down to zero flow rate by 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months .

The electric boiler feed pumps are manually tripped at 64 minutes.

The resulting turbine building temperature is 140 degrees Fahrenheit at 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

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Attachment I to P-85460 HRH-16

- Reactor building small break in the hot reheat line.

- The same operator actions and leak rate as HRH-15. A leak greater than this is enveloped by HRH-14.

The resulting reactor building temperature is 126 degrees Fahrenheit at 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

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Attachment 1 to P-85460 CONFIRMATORY ACTION 4: Provide an evaluation of the estimated time and primary coolant temperature beyond which the PCRV should not be depressurized during the remainder of a PCRV liner cooldown following a postulated high energy line break-from the 35 percent power level.

PSC EVALUATION: The initiation of depressurization is governed by the upper plenum gas temperature (see FSAR Section D.2.3). Initiation of depressurization prior to the upper plenum gas temperature exceeding 1350 degrees Fahrenheit ensures that overheating of the HTFA and helium purification cooler heat exchangers is avoided during the high flow depressurization transient.

i Following a high energy line break at 35 percent power, With neither forced circulation cooling nor liner cooling, the upper plenum gas temperature increases to the

1350 degrees Fahrenheit limit in about 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

. The time beyond which PCRV depressurization should not be initiated following a 4

postulated high energy line break from 35 percent power is thus 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

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' CONFIRMATORY ACTION 5: Provide an estimate of the maximum reactor power level at which it would be safe to perform a pressurized PCRV liner cooldown while protecting the integrity of the PCRV liner and liner cooling system.

PSC EVALUATION: The maximum reactor power level at which it would be safe to perform a pressurized PCRV liner cooldown while protecting the integrity of the PCRV liner and liner cooling system is in excess of 50 percent power. This conclusion is based on the results of a pressurized RECA analysis at 50 percent power. The results of this calculation are shown in Table 5-1. The pressurized case at 35 percent power is shown for comparison purposes. This shows that the maximum heat flux for 50 percent power does not increase significantly from the 35 percent power case.

The key parameter that did change appreciably is the time when liner cooling must be initiated, which decreased from 29.4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months for the 35 percent power case to 17.6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months for the 50 percent power case.

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Attachment 1 to P-85460 TABLE 5-1 COMPARISON OF CONSEQUENCES OF A HIGH ENERGY LINE BREAK FROM 50 PERCENT AND 35 PERCENT WITH THE PCRV PRESSURIZED TIME TOP HEAD-MAX PEAK FUEL MAX AVG. FUEL COVER PLATE FAILS POWER TEMPERATURE TEMPERATURE (i.e. TEMP EQUALS LEVEL DEGREES F DEGREES F 1500 DEGREES F) 50 percent 2420 1500 17.6 HOURS 35 percent 2150 1450 29.4 HOURS MAX LINER DEPTH OF CONCRETE POWER TEMPERATURE AB0VE 400 TIME LINER LEVEL DEGREES F DEGREES F COOLING ACTUATED 50 percent 750 APPR0X. 13" 17.6 HOURS 35 percent 900 APPR0X. 14" 29.4 HOURS

Attachment 1 to P-85460 l CONFIRMATORY ACTION 6: Submit a plan and perform inservice inspections on several critical areas of Fort St. Vrain's high energy piping to verify the integrity of this piping, prior to returning the plant to operation.

PSC EVALUATION: An ins,pection program to verify the integrity of the Main Steam, Hot Reheat, Cold Reheat, and Feedwater Systems will be accomplished prior to startup. This verification will be accomplished by either reviewing recent maintenance work documentation or by performing non-destructive test inspections.

A review of the plant maintenance and/or modification documentation has revealed that the Feedwater System in the area of Heaters 5 and 6, and the Cold Reheat System in the area of the Helium Circulator discharge piping has been opened in the recent past. These areas are considered to be representative of the Feedwater and Cold Reheat Systems respectively. During any maintenance activities, routine visual examinations are performed, and if degraded conditions are identified as a result of this routine examination, then Non-Conformance Reports are issued. Therefore, a review of this documentation will allow a determination of whether or not a substantial degraded condition may exist in these systems.

For the Hot Reheat and Main Steam Systems, four inspections locations have been identified for volumetric inspection to assess the current condition of the systems.

The locations for inspection are as follows:

1. Main Steam butt weld of 8" pipe to 8" x 16" Latrolet on PSC Drawing ISO 14A-98, item 16.

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',; . 3'. Main Steam 8" elbow mid-point on PSC Drawing ISO 14A-90, item 756.

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4. Hot Reheat butt weld of 20" pipe to 20" x 25" adapter to Steam Chest on

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% <he high stress and turbulent zones of these systems.

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Inspection methods s and analyses of the

, results will be based upon the. original FSV piping construction code, ANSI B31.1. The f'c review and evaluation of the results of these inspections will be considered indicative of the existing high energy piping system condition.

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. to P-85460 CONFIRMATORY ACTION 7: Evaluate the need for actuating the Fort St.

Vrain reserve shutdown system for postulated high energy line breaks from pcder levels up to 35 percent. Determine the latest time by which manual actuation must be accomplished, if needed, and evaluate the feasibility of taking the required manual actions by this time.

PSC EVALUATION: Actuation of the Fort St. Vrain reserve shutdown system is required following a LOFC at 100. percent power. This insertion is necessary to provide negative reactivity insertion to counteract the large amounts of fission product poisons being driven out of the core due to fuel particle coating failure. For all of the cases analyzed in Evaluation 1 the peak fuel temperatures are several hundred degrees below the fuel temperature where fuel particle coatings exhibit significant failure. Thus, neither the pressurized nor the depressurized situations require the reserve shutdown balls to be inserted.

However, in accordance with the defense in depth concept, the procedures contain instructions to manually insert the reserve shutdown systems together with the manual initiation of depressurization and liner cooling with fire water.

There are two factors which determine the time by which the reserve shutdown- balls must be inserted. These two factors are:

1. The time when the temperature of the control rod guide tubes exceed 2075 degrees Fahrenheit. Tests have been conducted on control rod tubes at this temperature and the results showed only minor

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deformation which posed no threat to the safety function of the guide tubes. For the pressurized cases the control rod guide. tubes never

( reach this temperature. For the depressurized cases the reserve shutdown system will be inserted in the same time frame (12-15 hours) ast initiation of depressurization and liner cooling. This is well before the control rod guide tube temperature reaches 2075 degrees Fahrenheit.

2. The time when the top head cover and liner insulation fails, which could . interfere with the insertion of the ' reserve shutdown balls.

This time is conservatively assumed to be the time when the top h'ead

. cover plate reaches a temperature of 1500 degrees Fahrenheit. The top head cover plate for the g- depressurized cases peaked at 1250 degrees Fahrenheit and thus'without any possible failure. For the worst pressurized situation, case 3 in Evaluation 1, the top head cover e plate 'was calculated to reach a temperature of 1500 degrees Fahrenheit by 29.4 hours.

Therefore, for this pressurized case the reserve shutdown balls must be inserted prior to 29.4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months into the event, t

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to P-85460 CONFIRMATORY ACTION 8: Confirm that procedures for taking the necessary operator actions in response to a high energy line break from 35 percent power will be in place prior to resuming Fort St.

Vrain plant operation. Evaluate the need for any changes to the Fort St. Vrain Technical Specifications to accommodate Fort St. Vrain operation up to the 35 percent power level restriction, including the possible use of a pressurized PCRV liner cooldown in the event of a high energy line break.

4 PSC EVALUATION: Normal and abnormal operating procedures were in place to support decay heat removal with the liner cooling system, utilizing fire water as a cooling medium prior to the environmental issue. For purposes of providing a concise and explicit procedure to manually lineup the associated systems, parts of the System Operating Procedures and Safe Shutdown Cooling Under Highly Degraded Conditions procedures were combined and issued as Operations Order No. 85-17. The procedure contained in this Operations Order assumes that no AC power, DC power,

. instrument air supply, or instrumentation is available to assist in the system lineups and continued operation. Operations personnel, including licensed operators and Technical Advisors, have completed training on this mode of operation.

This procedure will be changed to accommodate the modifications to permit manual actuation of the circulator brake and seal system and to permit operator action to shutdown the electric feedwater pump and auxiliary boilers. These changes will be in place and personnel trained prior to plant restart to achieve the 35 percent power level. Also operations personnel will be trained in the use of cool suits and Scott Air-Paks.

Attachment 1 to P-85460 There is no need for any changes to the Fort St. Vrain Technical Specifications to accommodate plant operation at 35 percent power. Although the consequences of a pressurized PCRV liner cooldown with fire water have been analyzed, this was performed to demonstrate that, even in the unlikely event that personnel access is not possible within 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months , the plant could be safely shutdown using a pressurized cooldown with liner cooling following a high energy line break. The intent is to operate the plant at 35 percent power level in accordance with the present Technical Specifications.

No Technical Specification changes for valve surveillances are necessary since all valve lineups are manual operations. Technical Specification limiting conditions for operation and surveillance tests of fire water and circulating water makeup pumps assures that this equipment will be ready and operable. The two fire water pumps were surveillance tested in June, 1985 for their rated flow and discharge head.

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Attachment 1 to P-85460 CONFIRMATORY ACTION 9: Document the extent of PCRV damage expected during a liner cooldown from 35 percent power following a high energy line break.

PSC EVALUATION: The following discussion lists the damage to the PCRV and internals in each of the four situations at 35 percent power analyzed in response to Confirmatory Action 1.

Case l. Following the steam line break that results in an LOFC, depressurization and liner cooling are initiated at 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months and 5 hours5.787037e-5 days 0.00139 hours 8.267196e-6 weeks 1.9025e-6 months respectively. In this case there is no damage to the PCRV.

The top cover plate temperature stays below 1500 degrees Fahrenheit, the liner temperature never exceeds 340 degree Fahrenheit and none of the concrete exceeds 400 degrees Fahrenheit. The temperature allowables for the liner and concrete are 1500 and 400 degrees Fahrenheit, respectively.

Case 2. Depressurization is initiated in 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months and liner cooling is started at 14 hours1.62037e-4 days 0.00389 hours 2.314815e-5 weeks 5.327e-6 months . The only difference between this and the first case is that the first 4 inches of the concrete exceeds the allowable temperature limit of 400 degrees Fahrenheit.

Case 3. The PCRV is never depressurized and liner cooling is initiated at 29.4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months just as the top head cover plate and all of the insulation is assumed to fail due to the top head cover plate temperature reaching 1500 degrees Fahrenheit. For this situation the liner retains its

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Attachment I to P-85460 integrity since the maximum temperature it experiences is 900 degrees Fahrenheit compared to the allowable of 1500 degrees Fahrenheit. The first 14 inches of concrete experience temperatures above the 400 degrees Fahrenheit allowable.

Case 4. The fourth case is also a pressurized situation with liner cooling initiated at 14 hours1.62037e-4 days 0.00389 hours 2.314815e-5 weeks 5.327e-6 months . In this case the top head cover plate was assumed to retain its integrity through the accident. The maximum temperature reached by the. top head cover plate was calculated at 1600 degrees Fahrenheit. The amount of damage to the PCRV was essentially the same as Case 2 since the calculations showed that the top head cover plate peaked at about 1600 degrees Fahrenheit.

u-Attachment 1 to P-85460 CONFIRMATORY ACTION 10: Evaluate the effect of a PCRV liner cooldown from 35 percent power level on the previously analyzed PCRV hot spots.

PSC EVALUATION: The PCRV liner cooling system, in conjuction with the PCRV liner thermal barrier, is designed to limit heat transfer.to the PCRV concrete and, thereby, limit local concrete temperatures, thermal stresses, and degradation. During normal operation, the concrete t.emperature is maintained less than 150 degrees F except for seven localized areas, referred to as "PCRV Hot Spots" at the following locations:

Top Head Refueling Penetrations Core Outlet Thermocouple Penetrations Core Barrel Seal / Core Support Floor Area

Peripheral Seal

Loop Divider Baffle Steam Generator Penetrations at the Bottom Head HTFA Cross-Over Pipe In each of these areas operating data revealed concrete temperatures above 150 degrees F in localized areas. Subsequent analysis by GA Technologies and Los Alamos National Laboratory showed that, although the design temperatures for the concrete were not being met, the local structural loads on the PCRV. internal components, liner, and concrete were within those allowed by applicable design codes.

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Attachment 1 to P-85460 Compliance with the design codes implies an acceptable level of performance under abnormal and extreme environmental conditions, although these were not explicitly evaluated.

The consequences of a permanent loss of forced circulation from the 100 percent power level have been evaluated in the Final Safety Analysis Report as Design Basis Accident No.

1 (DBA No. 1). In supplementary analysis of DBA No. 1, described in Section D.2.3 of the FSAR, the additional loss of all PCRV and core support liner cooling water systems is assuraed to occur simultaneously with the loss of forced circulation and last for 30 hours3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months .

Under these circumstances, PCRV internal temperatures increase to above 1500 degrees F where upon portions of the thermal barrier fail, resulting in higher heat flux to the liner. Concrete temperatures exceed normal values in all areas adjacent to the liner which renders the previously analyzed PCRV hot spots inconsequential. For this accident scenario, concrete temperatures exceed 400 degrees F to a depth of about 10 inches into the concrete and 150 degrees F as far as about 35 inches into the concrete. However, since. " concrete suffers no significant strength reduction up to a temperature of 400 degrees F" and the unaffected concrete provides adequate strength to resist the applied loads, the FSAR concludes that "the PCRV is safe for this type of condition".

Likewise, during the postulated cooldown from 35 percent power following a high energy line break, it is assumed that there is a permanent loss of forced circulation and that cooling water flow to the liner cooling system is interrupted for up to 29.4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months .

The differences between the two postulated 1

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Attachment 1 to P-85460 j accidents are 1) the power level of interest now is 35 percent power, while the DBA No. 1 power level was 100 percent power and 2) the current 35 percent power level accident scenario does not require PCRV depressurization while the FSAR scenario did.

The liner / concrete temperature profiles associated with the depressurized 35 percent power level accident are significantly less severe than those for the 100 percent power DBA No. 1. Therefore, for the depressurized cases less concrete will be affected and the FSAR conclusion about the integrity of the PCRV in this type of an accident remains valid including the hot spots.

For the worst pressurized situation, Case 3 in Evaluation 1, up to 14 inches of the concrete adjacent to the liner experiences or temperature in excess of 400 degrees Fahrenheit. Although this is slightly more than that analyzed for the 100 percent CBA-1 accident in the FSAR it is well within the limit established in FSAR Section D.2.3.3.

Attachment 1 to P-85460 CONFIRMATORY ACTION 11: Evaluate the impacts on the integrity of the PCRV liner cooling system of re-establishing liner cooling after prolonged periods of core heating without liner cooling or forced circulation cooling. Verify that the impacts of re-establishing liner cooling following a postulated high energy line break from 35 percent power with a pressurized liner cooldown would be no worse than those for the depressurized liner cooldown after 30 hours3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months from full power previously evaluated in the Fort St. Vrain FSAR. Compare the associated heat fluxes to the maximum acceptable heat flux to the PCRV liner cooling system assuming single loop operation.

PSC EVALUATION: The pressurized liner cooldown case from 35 percent power after a 29.4 hour4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months interruption of liner cooling (Case 3) is nearly identical to the depressurized liner cooldown from 100 percent power after a 30 hour3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months interruption of liner cooling during a permanent loss of forced circulation (see Table 11-1). The 30 hour3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months interruption of liner cooling from 100 percent power is described in FSAR Section D.2.3.2. A detailed analysis of the effects of re-establishing liner flow after 30 hour3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months for the 100 percent case is included in this letter as Attachment 6.

Since the worst case top head heat fluxes for the 100 percent depressurized case and the 35 percent pressurized case are nearly identical, the effects on tube integrity of re-establishing liner flow would be essential identical. The effects of re-establishing liner . cooling water flow for the 100 percent case are discussed in FSAR Section D.2.3 and Attachment 6 to this letter.

Both of the top head heat fluxes reported in Table 11-1 are well below the 24,000 Btu /hr-Attachment 1 to P-85460 ft2 ~ maximum heat flux at which boiling will occur (see Section 5.3.3 of EE-EQ-0019, Attachment 4). Except for the boiling which will occur during the approximately-12 minute period required to re-establish liner cooling flow, no further boiling will occur during liner cooldown from 35 percent power with a pressurized PCRV after a 29.4 hour interruption of forced circulation and liner cooling. This represents the worst case condition based on the 35 percent power

restriction.

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TABLE'11-1 PCRV LINER COOLING HEAT FLUX COMPARIS0N 100% POWER DBA-1 35% POWER CASE 3 DEPRESSURIZED PRESSURIZED MAX. TOPHEAD 1500 Degrees 1500 Degrees COVER PLATE Fahrenheit Fahrenheit TEMPERATURE AT FAILURE TIME OF COVER 28 Hours 29.4 Hours PLATE FAILURE TOP HEAD HEAT 14,200 Btu 14,460 Btu FLUX hr-ft2 hr-ft2 REFERENCE -FSAR - D.2.1.1 EE-EQ-0019 D.2.3.2 5.3.3 0.2.3.3 r

Attachment 1 to P.85460 CONFIRMATORY ACTION 12: If Fort St. Vrain operators will be required to take manual actions in environments whose temperatures exceed normal power plant operating temperatures to respond to high energy line breaks from 35 percent power, provide the NRC with information on the ability of operators to work in these higher temperature environments and the need for operators to utilize cool suits.

PSC EVALUATION: As shown by the temperature profiles included in Confirmatory Action 3, temperatures for the various scenarios associated with 35 percent power level show that temperatures are at 180 degrees Fahrenheit or less by four hours and at 135 degrees Fahrenheit or less by 15 hours1.736111e-4 days 0.00417 hours 2.480159e-5 weeks 5.7075e-6 months .

Access into areas where the temperatures are in the 180 to 135 degrees Fahrenheit range can be accomplished with the use of cool suits. The cool suits selected by PSC consist of an ice vest with approximately 9 pounds of ice and an insulating jacket. The use of these cool suits in conjunction with Scot Air-Paks will allow the performance of required manual actions in hot (180 degrees Fahrenheit), moist environmental conditions.

This conclusion is supported by Dr. Thomas Bernard, Westinghouse Electric Corporation, who has been working under EPRI contract for the past two years on heat stress management.

Access and the associated manual actions can be accomplished without protective clothing at temperatures of 135 degrees Fahrenheit and less. This position was previously discussed with the NRC on October 29, 1985 and is supported by EPRI Report NP2868 and informal surveys.

44

Attachment 1 l to P-85460 It can be concluded that initiation of liner cooling at 29.4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months following the loss of forced circulation can be easily accomplished without relying on protective clothing.

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Attachment 1 to P-85460

' CONFIRMATORY ACTION 13: Provide a summary list of which systems and their associated equipment items are considered qualified for 35 percent power operation and a confirmation that no unqualified equipment items are needed.

PSC EVALUATION: The following systems / equipment items are required for responding to a high energy line break at 35 percent power:

A. Helium Circulator Brake and Seal System B. PCRV Liner Cooling C. Reserve Shutdown System D. PCRV Depressurization Equipment E. Fuel Storage Facility Cooling F. Equipment to Mitigate Steam Production DISCUSSION A. Helium Circulator Brake & Seal System The need for primary coolant boundary isolation by this system is discussed in Section 5.2.1 of EE-EQ-0019. This isolation can be accomplished manually without the need for electrical components.

B. PCRV Liner Cooling This system and its associated equipment items are discussed in Section 5 of EE-EQ-0019. Tables 1, 2 and 3 summarize the equipment / manual actions required for liner cooling.

The electrical items relied upon are listed in Table 1.

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Attachment 1 to P-85460 The basis for taking credit for operation of this equipment is as follows:

1.) PSC does not expect that there would be a loss of outside power coincident with a HELB.

The reasons for this are as follows:

a) At 35 percent power Fort St. Vrain does not represent a large percentage of the total PSC grid, b) Fort St. Vrain has tripped several times in the past from power levels well above 35 percent without causing a loss of outside power.

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i c) The Fort St. Vrain switchyard, in addition to interconnections with the PSC system, is also directly interconnected with an outside power system (Platte River Power Authority).

2.) There are multiple circuits and power sources to the required electrical equipment as follows:

i a) With outside power available P4501, P4118 and P4118S can be powered via either the normal cables (some of which are l

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1 ss o Attach' ment 1 to P-85460 in the turbine building HELB environment or via an ACM backfeed (cables are all in mild environments).

b) If offsite power were lost, power can be supplied via the ACM Diesel Generator.

(Cables and other items are in mild environments.) If an HELB were to occur in the reactor building the emergency diesel generators (cables in turbine building) would be available to supply power via cables in a mild environment.

3.) P4501S is diesel driven and independent of offsite power.

Cables in the auto start circuitry are in the harsh environment. However, these cables cannot prevent local manual start of this pump.

C. Reserve Shutdown System This system is detailed in Appendix C to EE-EQ-0019 and may be actuated manually without the need for electrical components.

D. PCRV Depressurization Equipment This equipment is detailed in Appendix B to EE-EQ-0019 and may be

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Attachment 1 to P-85460 1 actuated manually without the need for electrical components.

E. Fuel Storage Facility Cooling System Equipment This equipment is detailed in Appendix D to EE-EQ-0019 and may be actuated manually without the need for electrical components.

F. Equipment to Mitigate Steam 4

Production In order to generate a bounding set of temperature profiles, very conservative assumptions were utilized in regards to termination of the high energy line break (HELB). That is, all valves regardless of location (reactor building or turbine building) were assumed to fail open. In reality, with the HELB occurring in only one or the other, many building (even though unqualified circuits at this time) that are in the building that did not have the HELB would be expected to be functional.

The conservative method utilized for terminating the HELB in the profile scenarios is to trip various pumps and allow the entire steam system inventory to blow down. The basis for tripping these pumps is as follows:

1.) Equipment items, electric power cables and electric control cables are all located in the harsh environment. It

o- o Attachment 1 to P-85460 is highly unlikely that a discriminatory common mode failure would occur. That is, it is not reasonable to expect that electrical equipment items and their associated power cables survive the HELB while at the same time all control cables fail.

Therefore any equipment that is still operating following a HELB is assumed to be controllable from a trip standpoint.

2.) Multiple methods exist for tripping equipment, such as tripping the equipment breaker, a 480V bus, or a 4160V bus.

3.) In a worst case scenario, all 4Kv buses can be de-energized from controls, cables and circuit breakers not in a harsh environment.

4 1

l l

l o-o Attachment 1 to P-85460

. CONFIRMATORY ACTION 14: Submit an engineering evaluation that describes the systems and equipment which will be utilized to respond to a high energy line break from 35 percent power.

PSC EVALUATION: Engineering Evaluation EE-EQ-0019 has been prepared and enclosed with this letter to describe these systems and equipment as well as the associated manual actions.

4

\

ATTACHMENT 2 PSC RESPONSE TO NRC ADDITIONAL INFORMATION REQUEST

p [

t Attachment 2 to P-85460

\

RESPONSE TO NRC REQUEST FOR ADDITIONAL INFORMATION TO SUPPORT 35 PERCENT POWER OPERATION

-QUESTION 1 "LCS"- When the ' Firewater System is providing cooling water, how many loops of the LCS would be running and how much flow would be supplied to each loop? _ What approximate pressure would be maintained in the LCS tubes? Would the operators be able to affect the re-routing of flow .from cooler areas of the PCRV to the upper side walls and top head; if so, what would be the flow per LCS tube?"

ANSWER If available, both loops would be utilized for liner cooling and each loop would provide approximately 1510 gpm. However, credit has only been taken for one loop operating. The worst case pressure in a cooling tube- occurs at the outlet of the highest elevation tube, and is-approximately 10 psig (or 22 psia). As specified by existing plant ,

procedures, flow would be redistributed to bias more than normal flow to the upper side walls and top head. The resulting redistributed flow per tube would be approximately 10 gpm. A more complete sununary of the results of the flow analysis is attached.

QUESTION 2 t "HTFA"- Exchanger type (counter-current, etc.) and calculated performance data (helium and water flows, inlet and outlet temps) for at least one operating point for the helium cooler. - How much water would the Firewater System supply to the HTFA- helium cooler under LOFC accident conditions?"

ANSWER-The heat exchanger in the HTFA is in a configuration such that the helium ' flows over a spiral tube configuration carrying the cooling water which enters and flows in a counter-current direction. The tubes then make a second pass carrying the cooling water in a parallel flow direction to the outlet.

.As far as performance is concerned, PSC previously submitted detailed information on the helium purification system performance during depressurization in a letter P-77250 dated Dec. 22, 1977 from J. K.

Fuller to Mr. Richard P. Denise. To summarize, the letter states that the helium exiting the HTFA 6

would not exceed 800*F based on the maximum heat load of 2.7 X 10 Btu /hr at a- helium flow of 3340 lbs/hr with a cooling water flow of 46 GPM ' at 80*F. This occurs at about I hour and 15 minutes into depressurization.

The new flow analysis performed by Proto-Power Corp. shows that approximately 48.5 gpm of firewater would be supplied to the operating HTFA. However, based on the recent GA Technologies analysis, depressur12ation will not be required following a total and permanent LOFC from 35 percent power operation.

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Below is a single performance point calculated by GA Technologies for the HTFA. '

Single Performance Point For HTFA Temp. In Temp. Out Flow Helium 1350*F 475'F lbs 1225 hr.'

Water 100*F 145'F 60 gpm i

1

_ QUESTION 3

" Helium Purification Cooler: SD-23 provides calculated performance data, but we don't know the exchanger type. Hcw much water flow would be supplied to this cooler during LOFC accident conditions?

ANSWER 4

The heat ' exchanger in the HPC is in a spiral tube configuration. The helium is carried through concentric spirally-wound coils which are connected to a central manifold. Water in the shell side flows in ,

a counter-current direction.

The flow analysis shows that approximately 84 gpm of firewater would be supplied to the operating HPC. However, based on the recent GA i

Technologies analysis, depressurization will not be required following a total and permanent LOFC from 35 percent power operation.

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Core Support Floor 429 <100 101 23 260

} Bottaa Head E 223 < 50 30E 23 260

Penetrations I Top Penetrations 208 >200 203 10 233 Lipper Barrel & 267 >850 850 13 240*

4 Boad i smer Barrel 282 < 50 50 23 260

Total Flow to Liner = 1510 Get l Maximous tymperatare rise for an indivhial tube occurs in the top head and i equals 93 F, corresponding to a tube outlet temperature of 178 F. This j temperatuoe rise conservatively -- the peak heat flux occurs over the entire top head ap.a. The maximun bulk temperature rice across the PCW is

i expected to be 25 F.

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b ATTACHMENT 3 FSV FUEL PERFORMANCE UNDER HIGH TEMPERATURE CONDITIONS i

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1.0 SMDfARY AND CONCLUSIONS The behavior of FSV initial core production fuel was studied under conditions predicted for FSV Design Basis Accident #1 (DBA #1). The purpose of the experimental program was to evaluate models developed to account for effects of microporosity induced OPyC failure on total coating failure during DBA #1. Test samples had been irradiated to average fast neutron exposures and kernel burnups expected upon removal of a 6 yr old fuel segment from th,e core. Total coating failure fractions were determined from Kr-85 release.

fractions measured while heating test samples having" intact or failed OPyC layers in the temperature range 900 to 2500 C. Observations relative to FSVT fuel performance include

, 1. Only fertile fuel with failed OPyC layers exhibited total

, ' coating failure at temperatures less than 1850 C.

2. Fertile A and B performance is virtually identical..

3. . . . . -

Total coating failure fractions remain constant as tempera-

tures are increased from 1100 C to temperatures exceeding

,, 1750 C.

4.

" - ~ ~ ~ ~ Failure fractions for samples with intact or missing OPyC

'. . layers did not ' Exceed 0.10 until t'est temperatures exceeded 2050 C.

~

, , 5. Performance above 2050 C is independent of the presence or

.abisen~cTof t' h'ii OPyC layer. . ..

The following conclusions were drawn relative to fuel performance models used for DBA_f1-analyses.

1. The basis for pressure vessel failure predictions used to evaluate fuel performance during DBA #1 is conservative.

2. Observed total coating failure fractions are much lower than values predicted for DBA #1 analyses. -

3. The margin associated with the heat load on the LTA during

-~

the depressurization stage of DBA #1 is not affected by microporbsity induced OPyC failure.

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2.0 INTRODUCTION

1 Analyses of irradiation data collected on FSV and LHTGR fuels sug-

gest that failure of the OPyC layer on TRISO fuels is related to OPyC micro-por.osity (Ref. 1). Recent evaluation of FSV fuel OPyC microporosity data sug-gests a potential for 5.6% failure of the OPyC layers on fissile fuel and 34% .

failure of the OPyC layers on fertile fuel in the first seven FSV fuel seg-ments (Ref. 2). The degree of OPyC failure is expected to increase linearly with fast neutron exposure from 0 at beginning of life to the values indicated above at 2.0 x 1025 ,7 ,2 (E)29 fJ " "# I * "#* #** " " " * ""

HTGR

remain constant as fast neutron exposures increase. This potential OPyC failure does not translate directly to total coating failure and fission pro-duct relcase; however, it does increase the probability for pressure vessel coating failure and fission product release. In o'rder to evaluate the poten-tial for increased fuel failure, pressure vessel performance models (Ref. 2)

[ were developed for FSV fissile and fertile fu.els having intact or failed OPyC

~

i layers. The models were based on kernel and coating dimensions and densities measured during FSV segment 7 fuel production. These models can be used to evaluate the' impact of OPyC f ailure on cotal coating failure and fission pro-duct release during normal reactor operation or during the hypothetical acci-dents that are considered for reactor licensing and siting applications.

Concern was expressed that the increased probability for total coat-ing failure that is associated with OPyC failure' could represent a substantial safety hazard as defined by 10CFR21. A committee was consequently convened to evaluate the situation. The key item identified and analyzed for presentation i to the committee was the heat load on the low temperature absorber (LTA) following depressurization during FSV DBA #1. The following steps were followed in this analysis.

1. Develop predictions of core average total coating failure fraction vs temperature for DBA #1 conditions that (n)- assume no OPyC failure and (b) account for expected levels of OPyC failure (Ref. 3, 4, 5). .

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. 2. Utilize performance models from step i to predict fuel failure fractions, fission product release, and the heat >

load on the LTA as a function of time during DBA #1 (Ref. 4,5).

3. Chapare the predicted' heat load with projections made using the original.FSV FSAR fuci performance models (Ref. 6).

These efforts' led to the conclusion that the margin between the LTA heat load estimated in Refs. 5 and 6 is not reduced significantly if one accounts for The 10CFR21 coc:mittee agreed possible' microporosity induced OPyC failure.

with this conclusion but requested that testing of FSV fuel with failed and intact OPyC layers be donc under DBA #1 conditions to confirm the fuel per-

formance assumptions. The results of the fuel test program are summarized in this report.

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3.0 TEST PROGRMt PilILOSOPld fuel The primary objectivle of the testf 4) program are conservative.

was to show that performance models used in the DBA #1 analysis I i f (Re . pre-analyses, This experimental verification would support the conclus onI seated to the 10CFR21 commit ee. The key steps utilized while -

developed using assumptions outlined in Ref. 3. I developing the models included: i calculation of Sic layer stress distributions for fuel with asintact a funct on (1) ty ,

.of irradiation exposure and temperature and failed OPyC layers using kernel and coa rs estimate total coating failure fractions for temperatu b e (2) <1725 C assuming the failure fraction equals the pro a-bility that calculated sic stresses are more positive than -2800 psi, l

assume that coating f ailure fractions increase linear C y (3)

with temperature from the pressure vessel value at 1725 to 1.0 at 2000 C, develop failure models for fissile and fertile fuel that (4) (a) assume no OPyC failure

~

' nd and failed(b) OPyC account layers, for expecte

, populations of fuel with intact a

- (5) develop core average failure models for tion exposure conditions projected for FSV, and 5 at 900 C assume (a) an initial failed fuel fraction of 0.0 (6) and (b) that failure f ractions increase with step 5, d ls would A specific experimental verification of the DBA fractions,#1 (2) performanc seg-require that test samples have (1) expected OPyC failured (3) the ment 7 kernel and coating property distributions, an It was not of irradiation exposures expected in an equilibrium i tics.

FSV core. .

possible to obtain test samples having these character s

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The approach utilized was to test the pressure vessel model by com-paring failure predictions for specific sampics of initial core production fuel with performance data collected under simulated DBA #1 conditions. The test samples were irradiated in the FSV fuel proof test (Capsule F-30, Ref. 7).

Samples were tested with intact and failed

OPyC layers. Failure predictions for temperature <1725*C were dade from Sic layer stress distributions that I

were calculated using kernel ,and coating property distributions and F-30 irradiation conditions for ea'ch individual sample. It was assumed that the failure fraction would equal the probability that calculated stresses were more positive than -2800 psi. Total coating failure fractions were assumed to increase linearly with temperature from the pressure vessel value at 1725 C to 1.0 at 2000 C. Comparison of experimental results with model predictions at this point results in a verification of models provided through the third step of the six step development process outlined above. Since the first three steps represent the basis for the model, verification through step 3 is equivalent to verification of the total model.

OpyC failure was simulated by removing OPyC layers during a 2-hr anneal in air at 900 C. '

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4.0 EXPERIMENTAL t .

4.1 SAMPLE DESCRIPTION i I .

Samples chosen for'this series of tests were from initial core FSV production fuel batches hich were irradiated in capsule F-30 (Re'f. 7).

Fissile A, fertile A, and fe tile B samples were tested with and without outer pyrocarbon layers; kernel and coating properties are given in Table 1.

A comparison of the test particle vioperties with those of segment 7 fuel (see Table 2) shows that the kernel and coating properties of the test samples are consistent with those of fuel currently in the FSV reactor. Irradiation conditions experienced by the test samples are given in Table 3; predicted FSV conditions are shown for comparison. Test sample fast neutron exposures exceed maximum values expected after 6 yrs of residence in FSV; kernel burn-ups are equivalent to average values expected for a segment removed after 6 yrs of residence in FSV.

4.2 TEST SETHODS

."' ~

FSV production fuel was tested using methods (Ref. 8,9) developed on the DOE sponsored core heatup simulation test (CHST) program. Key stages of the test method are itemized in Table 4. Test conditions were chosen to simulate predicted FSV DBA #1 conditions. Total coating failure fractions were d,etermined from Kr-85 release fractions measured (1) af ter removal of the OPyC layer and (2) continuously as all samples were heated from approx.

1100 to approx. 2500 C. A detailed description of the test method is provided in Appendix A. ,

4.3 CRST CONDITIONS Each of the three samples identified in Tables 2 and 4 was tested with intact and missing (failed) OPyC layers. A summary of the CllST condi-tions for each of the six tests conducted is given in Table 5. The rates at which temperatures were increased with time are representative of maximum l heating rates shown for DBA #1 in the FSV FSAR (Ref. 10).

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5.0 RESULTS __, ,

Krypton 85 release ractions measured during 25-27 hr CHST's of PSV fertile fuel are shown as a function of temperature on Figs. 2 and 3 I

(Ref. 11-14). Krypton release fractions detected following removal of the OPyC layers from fertile A and B samples at 900 C were, respectively, 0.025 -

i and 0.011. Radiographic examinations following OPyC removal suggested total _

coating failure fractions of 0.02 for the fertile A fuel and 0.01 for the fertile B fuel. The good agreement between total coating failure and Kr-85 release fractions observed after OPyC removal suggests that the Kr-85 release fraction per failed particle was one at this stage. This is not surprising since the particles were heated in air and any ThC2kernels exposed by coat-ing failure would be expected to oxidize and release their total Kr-85 inven-tory. The inventory of Kr-85 released at t.his stage was added to Kr-85 inven-tories released during subsequent CHSTing to obtain the total Kr-SS released as a function of temperature.

~

All test samples were held at approx. 1100 C for 10 to 20 hrs prior to CHSTing. Krypton 85 release measurements were made within 1 hr of reaching 1100 C and after the 10-20 hr hold period was completed. The only samples showing Kr-85 release at this point were fertile A and B particles with miss-ing OPyC layers. Krypton 85 release was detected from both samples within I hr of reaching approx. 1100 C. Although additional Kr-85 release was de-tected during the 10-20 he hold, the increases were negligible. The time l

period over which most of the Kr-85 was released at 1100 C (<1 hr) suggests l that the Sic and IPyC layers failed by a pressure vessel mechanism.

Release of Kr-85 would occur in two stages following failure at 1100*C. Krypton 85, that had been released from the kernel to the buffer during irradiation, would be released rapidly following total coating failure l

(stage 1). Stage 1 release from fertile A and B fuels with missing OPyC layers was ' observed during the 10-20 hr hold at 1100 C. Krypton 85 remaining l

7

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, in the kerncis of failed particles would then be released slowly via a diffusive mechanism (stage 2). This was also observed during the hold at 1100*C. Recent studies conducted on laser failed TRISO UC2 fu 1 have shown that Kr-85 release from carbide kerncis would be complete at approx.1750 C under CHST conditions (Ref. 15). If no additional coating failure occurred in the temper'ture a range 1100-1750 C, the Kr-85 release fraction would increase smoothly with temperature. If additional total coating failure occurred'in the temperature range 1100-1750 C, rapid increases in Kr-85 release (stage 1) would be superimposed over the smooth increase assdeiated with stage 2 release. Careful examination of the Kr-85 release data from tests of fertile fuel with missing OPyC layers showed a smooth increase with tempera-

~

ture in the range 1100-1750 C suggesting that no additional failure occurred.

The Kr-85 release fraction at 1750 C can therefore be assumed equivalent to the failed fuel fraction at 1100 C. ,

Results obtained from fertile fuels with intact OPyC layers suggest no Kr-85 release (coating failure) until CHST temperatures approach 2050 C.

Because of the high temperatures reached before ob, serving Kr-85 release, it is reasonable to assume.for these fuels, that release fractions are equivalent to total coating failure fractions.

A general examination of results from the four fertile fuel CHST's (Figs. 2 and 3) leads to two additional conclusions. The first is that fertile A and B performance is virtually identical. The second is that fertile fuel performance at temperatures exceeding 2050 C is independent of the presence or absence of the OPyC layer.

Krypto'n 85 release fractions measured during CHSTing of FSV fissile A fuel are shown as a function of temperature on Fig. 4. No Kr-85 release (coating failure) was detected in samples with or without OPyC's until tem-paratures exceeded 1850 C. Krypton 85 release fractions observed from fissile fuel with missing OPyc layers were higher than observed from fuel with intact OPyc layers at temperatures >1850 C. Fuel with missing OPyC layers was tested

, in a 26.5 hr CHST; fuel with intact OPyC's was tested in a'9.0 hr CHST. Past i

! ' tests have shown that the release fraction at any given temperature will in-irease*as the length of a CHST increases. This result coupled with the results

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A 3 .

obtained on FSV fertile fuel leads to the potentially conservative assump-tion that Kr-85 release from fissile fuel at temperatures exceeding 2050 C would be independent of the presence or absence of the OPyC layer.-

- Civen the assumption indicated above, one reaches .the final conclu-sion that Kr-85 release fractions at temperatures exceeding 2'050 C are in-dependent of OPyC condition (failed or intact), fuel size (A or B), and fuel type- (fissile or fertile). ,

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6.0 DISCUSSION ,

6.1 PREDICTED VS OBSERVED' FAILURE FRACTIONS i -

The primary purpose of this test sccies was to evaluate fuel failure assumptions utilized to predict the heat loal on the LTA during FSV DBA #1.

To accomplish this, the experimentally measured Kr-85 release fractions must .

be related to total coating failure fractions. Based on discussions presented in Section 5.0, the following assumptions were made to define failed fuel fractions from the measured Kr-85 release fractions.

Failure mode: all gas release occurred after failure of the IPyC, sic, and OPyC layers if the OPyC layer was present (i.e., no Kr-85 was released by diffusion or permeation through intact Sic or PyC layers).

OPyC removal stage: Total coating failure that occurred while OPyC layers were removed by heating in air at 900'C resulted in oxidation of ex-posed kernels and 100% release of stored fission gases. Kr-85 release frac-tions observed at this stage therefore equal total coating failure . fractions.

Temperatures >1750 C: Tests conducted on laser-failed TRISO'UC2 fuel under CHST conditions (Ref.15) suggest 100% release of.3r from carbide kernels at temperatures exceeding 1750 C. Krypton 85 release fractions ob-served at temperatures exceeding 1750 C are therefore equivalent to failed fuel fractions., ,

Temperatures in the range 1100-1750 C: Fuel failing at 1100 C would release that fraction of Kr-85 released to the buffer during irradiation.

The Kr-85 remaining in the kernel would be released slowly with increasing temperature between 1100 and 1750*C. Since Kr-85 data indicated gradual in-creases in release between 1100 and 1750 C but no " bursts" of Kr-85, which would be expected following failure of additional fuel, it will be assumed 10 t

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that failed fuel fractions at 1100 C are the same as indicated by Kr-85 release fractions at 1750 C.

Heasured Kr-85 release fractions and the corresponding failed fuel fractions are shown for each test in Table 6. Total coating failure frac-tions predicted for each testf sample (Ref.16) are compared dith failure frac-I tions determined experimentally on Figs. 5-10. Predicted failure fractions I

for test samples having intact OPyC layers are, although somewhat conserva-I tive, in reasonable agreemen ; with experimental values at temperatures

<1725 C. Predictions are ex,:remely conservative for temperatures 2,1725 C. '

Predicted failure fractions 'for samples having missing OPyC layers are ex-tremely conservative for temperatures 2,900 C. .

The predictions shown on Figs. 5-10 were made using the same failure criterion used to develop the core average performance models utilized to evaluate FSV fuel performance during DBA #1 (Refs. 3,4,5). Since the pre-dictions have been shown to be extremely conservative for individual batches of initial core production fuel it is concluded that FSV core average per-formance predictions are also conservative.

6.2 FSV CORE PERFOMMNCE DURING DBA #1 .

The basis for models used to describe the impact of OPyC microporo-sity on fuel failure during FSV DBA #1 has been shown to be conservative. The data used to draw this conclusion were developed from CHST's conducted on FSV production fissile and fertile fuels that were irradiated to average conditions expected for a fuel segment removed af ter 6 yrs residence in FSV.

The data collected on this high exposure fuel can be used to develop an en-i pirical model for FSV fuel behavior under DBA #1 conditions. If one assumes that all fuel in FSV<will perform like the test fuel and then compares the empirical model with failure assumptions made in Refs. 4 and 5, an e~ valuation

,. of the assumed core average failure model can be made. The steps used to develop the empirical model are described below.

Fissile Fuel with Failed OPyC Layers: The variation in expected

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failure fraction with temperature was developed from the test conducted on i

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fissile A fuel. Ninety percent confidence bounds for the expected failure j fractions were defined on the basis of sample size. Although no fissile B fuel was tested, it is assumed that fissile A and B fuel behavior is identi-cal based on fertile A and B fuel test results.

Fertile Fuel with Failed OPyC Layers: Results fror$ both tests con-ducted on fertile fuel with failed OPyC layer's were combined to define ex-

' p.ected and 90% confidence bounds for expected failure fractions.

-Fissile and Fertile Fuel with Intact OPyC Layers: Fissile fuel

~

with intact OPyC layers was tested in a 9.0 hr CHST. All other data were collected from 24.75-27.0 hr C11ST's. This one fissile test suggested failure fractions less than those observed in all other tests. Data collected from both CHST's conducted on fertile fuel with inta'ct OPyC's were combined and assumed to represent expected and 90% confidence bounds for expected failure fractions of fissile and fertile fuel having intact OPyC layers'. -' -

. Expected values and the range for expected failure fractions for fissile and fertile fuels with missing or intact OPyC layers are given as a function of temperacare in Table 7. Values shown for fissile and fertile fuel with intact OPyC layers represent expected behavior of a FSV core that experiences no OPyC failure. The data in Table 7 were combined to determine fissile and fertile total coating failure fractions, that account for expected OPyC failure fractions, using. *

"^

f g~= (1-FOPyC II)

~

, 1+ OpyC 2 where ,

fg= total fissile or fertile fuel failure fraction ,

OM = fertile OPyC failure fraction (0.056 for fissile fuel and 0.34-' for -

F fuel., Ref. 2) fy= total coating failure probability for fuel with intact OPyC layers ,

f = total coatifig failure probability for fuel with failed OPyC

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. layers -- -- -

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The fissile and fertile fuel failure fractions were then combined to yield

. an " average" failed fuel fraction (favg) using

'avs = 0.451 ffg, + 0.549 f f ,,g .

(2) where .

, fgg, = fissile fuel failure fraction f ,,g=

f fertile fuel failure fractica The values 0.451 and 0.549 are, respectively, the fraction of fissions in fissile and fertile fuel in a 6 yr old FSV fuel segment. Expected and high confidence values for ffg,, ff , g and f, are given in Table 8. Comparison of the range of values for f, (Table 8) with the range of failure values for fuel with no OPyC failures (Table 7) shows, as expected, that total coat-ing failure fractions will b'c increased slightly because of OPyC failure at normal reactor operating temperatures. Failure models assumed in Refs. 4 and 5 are compared with the high confidence bound empirical models (Tables 7,8) on Figs. 11 and 12. Figure 11 assumes no OPyC failure; Figure 12 assumes ex-pected fissile (0.056) and fertile (0.34) fuel OPyC failure fractions. In each case, the failure fraction based upon experimental observations is less than assumed in DBA #1 analysis.

Although the presence of OPyC failure. increases the expected total coating failure fraction during normal reactor operation, the most significant observ'ation is that the increase in failure fraction, as temperatures are increased to simulate DBA #1 conditions, is essentially independent of the

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presence or absence of OPyC failure during normal reactor operation. [The.,,"~~

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fuel performance models used in Refs. 4,5 assumed an initial total coating failure fraction of 0.05 at 900 C to be consistent with the design circulating I activity allowed by FSV technical specifications (assumes no hydrolysis).

l If the empirical models based upon the results of this test program (Tables 7,

8) are adjusted to a failure fraction of 0.05 at 900 C, using the same method used to develop models for Refs. 4 and 5, the resulting failure frac-didns are 0.55 at 900-1750 C, 0.10 at 2050 C, approx. 0.80 at 2175 C, and 1.0 e

13

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a't 2300 C independent of the presence or absence of fuel with failed OPyC layers. *This model for fuel failure is compared with original FSV fuel per-formance assumptions made when analyzing DBA #1 (Ref. 6) and assumptions made in Refs. 4 and 5 on Fig.13. On the basis of this comparison one would not !

l expect the margin between the heat load on the LTA estimated in Ref. 6 and the expected heat load to be reduced by the presence of microporosity in-duced OPyC failure. .

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l REFERENCES

1. Memo, W. J. Kovacs to Distribution, " Justification of a Proposed OPyC !!icroporosity and Coating Rate Specification for FSV Fuel (C. N. No. 004512)," WJK:001:FMB:79, Jan. 9, 1979 Rev. 26 Jan.

1979. -

2. W. J. Kovacs, " Pressure Vessel Performance Models for FSV TRISO (Th/U)C2 and ThC2 Fuel", Doc. No.184-45, Feb. 7,1979.

3. Memo, C. L. Smith to F. A. Silady; " Fuel Failure Assumptions for FSV Hypothetical Accident Analyses," FCB:090:CLS:78, Dec. 7, 1978.

4. Memo F. A. Silady/F. S. Dombek to D. Alberstein, "FSV Technical Memo: Effeet of Failed Outer PyC Coatings on LTA Analyses,"

SAM:299:FS/.FD:78, Dec. 7, 1978.

5. Memo, F. S. Dombeck/F. A. Silady to Distribution, "FSV, Summary of OPyC Effect on DBA #1", SAM:309:FD/FS:78, December 19, 1978.

6. ,

J.K. Fuller, "PCRV Depressurization - Ioss of Forced Circulation Accidents", PSC letter to NRC, P-77250, December 22, 1977.

7. C. B. Scott and D. P. Harmon, " Post Irradiation Examination of Capsule F-30," GA-A13208, April 1, 1975.

,8 . "ETGR Fuels and Core Development Program Quarterly Progress Report For the Period'Ending' November 30,1976," GA- A14180, December 27, 1976.

9. " Procedure for Burning Back Outer Pyrocarbons of Irradiated Fuel Particles", Document No. 903878, Issue A, in review.

10. Fort St. Vrain Nuclear Generating Station Final Safety Analysis Report, Docket No. 50-267, Public Service Company of Colorado, Appendix D.

11. GA Notebook #7284.

12. GA Notebook #7828

13. CA Notebook #7945

14. GA Notebook #7946

. 15

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REFERENCES (continued) .

15. B. F. Myers and R. E. Morrissey, "The Measurement and Modeling of Post Irradiation Fission Product Release From HIGR Fuel Particles Under Accident Conditions," CA- A15018, December 1978. :

1

16. Meno, W. J. Kovacs to C. L. Smith, " Fuel Failure Predictions for CHST Samples," WJK:048:DIB:78, 29 Dec. 1978.

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q OPyC Failure f L.

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4 I' h *

/*No OPyC Failure 9

___--------s 9 . , , , ,

800 1000 1200 1400 1600 1800 2000 .

TEMPERATURE ('C) .

. Figure 1. Total Coating Failure Fraction as a Function of Temperature '

Assumed for Evaluation of the Impact of Microporosity Induced e OPyC Failure on the LTA Head Load During FSV DBA f1 l

l

l.0 . . . . . .

dG CT6A 2399 (Fertile A) Oo O OPyC Missing (26.0 hr 'CHST) 0.8 - O OPyC Intact (24.75 hr CHST) ,

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900 1000 1300 1500 1700 1900 2l00 2300 2500 l

1 -

1 TEMPERATURE (*C)

Figure 2. Kr-85 Release Fractions Observed as a Function of Temperature

during 30 hem CHST's conducted on Fertile A FSV Production Fuel With and Without Outer Pyrocarbon Coatings , ,

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1

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1.0 , , , , . .

g.

CT6B 932 (Fertile B)

Q f0PyCMissing (27.0 hr CHST)

}

O lOPyCIntact ( (26.5 hr CHST) "

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1900 S-2l00 2500 2500 I 900 1000 1300 1500 1700 4

TEMPERATURE (*C) l l

Figure 3. Kr-85 Release Fractions Observed as a Function of Temperature l During 30 hour3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months CHST's Conducted on Fertile B FSV Production Fuel l With and Without Outer Pyrocarbon Coatings ,

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1 l.O . . . . . .

. gv ;

CU6A 6328 (Fissile A) .

G l g Q OPyC Missing (27 hour3.125e-4 days 0.0075 hours 4.464286e-5 weeks 1.02735e-5 months CHST) ,

. O O.8 -

O OPyC Intact ( 9 hour1.041667e-4 days 0.0025 hours 1.488095e-5 weeks 3.4245e-6 months CHST)

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900 1000 lb 500 h00 00 lbO 2NOO 2500

. TEMPERATURE (*C)

Figure 4., Kr-85 Release Fractions Observed as a Function of Temperature .

During CHST's Conducted on Fissile A FSV Production Fuel With and Without Outer Pyrocarbon Coatings l

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. 2 Fissile A -

o i Intact OPyC

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cr 9.0 hr CIIST *

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1 Observed

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900 110 0 1300 1500 1700 1900 2l00 2300 I

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TEMPERATURE ('C) i Figure 5. Comparison of observed and Predicted Total Coating Failure Fractions for FSV Core I Production Fuel Tested Under ,l Hypothetical DDA il Conditions

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4 ct , 24.75 hr CHST f .

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i I Predicted I l '

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900 110 0 1500 1500 1700 1900 2l00 2300 k

. TEMPERATURE (*C) -

Figure 6.l Comparison of Observed and Predicted Total Coating Failure

, Fractions for FSV Core I Production Fuel Tested Under Hypothetical DBA #1 Conditions -l l

4

~

1.0 , , , , , , ,

2 Fertile B I

' o

~ Intact OPyC I

7 l.

D 4 0.8, - CT6B-932, 4.7% FIMA / .

! m 26.5 hr CIIST / .

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Predicted f

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s /

O' > a - " --T~~~ o -- - -

a i . 900 110 0 , '1300 1500 1700 1900 2l00 2300

! TEMPERATURE (*C) ,

Figure 7. Comparison of Observed and Predicted Total Coating Failure ,

Fractions for FSV Core I Production Fuel Tested under Hypothetical DBA il1. Conditions ; .

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3 I -

F Predicted j y O.8 , g ,

x

u. /

/ .

g Fissile A i l w i /

Failed OPyC -

C / ;

O.6 b / CU6A-6328, 18.2% FIMA -

h 26.5 hr CHST I

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. .J Q,2 '. .

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p . observed .

i l 0 m a ' - '

~

900 110 0 1300 1500 1700 1900 2l00 2300 i

i TEMPERATURE (*C)

! Figure 8. Comparison of Observed and' Predicted Total' Coating" Failure j Fractions for FSV Core I Production Fuel Tested Under *

Hypothetical DBA #1 Conditions 1

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~

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e / CT6A-2399, 4.7% FIMA i z ' -

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. Observed .

-C O l' - i i ' . .

900 110 0 - 1300 1500 1700 1900 2l00 2300 i '

(

l TEMPERATURE ('C) i i

l Figure 9. Comparison of Observed and Predicted Total Coating Failure ,

l Fractions for FSV Core I Production Fuel Tested Under l

Hypothetical DBA f1 Conditions , ,

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Predicted /

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g/ / Fertile B Id /

/ Failed OPyC .

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/ CT6B-932, 4.7% FIMA

h. / 27.0 hr CHST w / s E O.4 -

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900 libO 13b 1500 i OO 1900 210 0 2300 (-

TEMPERATURE (*C)

Figure 10. Comparison of Observed and Predicted Total Coating i'ailure Fractions for FSV Core I Production Fuel Tested Under Hypothetical DBA il Conditions

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OPyC Failure f

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900 110 0 1300 1500 1700 1900 2l00 2300 i

TEMPERATURE (*C) l Figure 11. Comparison of Failure Model Used for FSV DBA i1 Analysis l Assuming no OPyC Failure with High Confidence Bound Model Based -

on Experimental Observations l

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OPyC Failure Fraction

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. 2 7 * '

.O Fissile = 0.056 f g ,

Fertile = 0.34 / ,

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900 110 0 1300 1500 1700 1900 2l00 2300 TEMPERATURE (*C)

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Figure 12. Comparison of Failure Model Used for FSV DBA #1 Analysis Assuming Expected OPyC Failure Fractions with High Confidence Bound Model Based on' Experimental Data e

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f 2 Original FSV Model *~~ [y

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-(Ref. 6) p Models Used on Refs. 4,5 o 0.8 - No OPyC Failure ~ ~ ~ ~ " ~ *

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4 OPyC Failure l k Model Based on CllST*~ ~ ~ ~

Results j, f*

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-jf O i 900 110 0 1300 1500 1700 1900 2l00 2300

- TEMPERATURE (*C)

Figure 13. Comparison of Models for FSV Fuel Failure versus Temperature for DBA #1 Conditions with Model Based on CHST Data. The Model Based on CIIST Results is Independent of Microporosity * *

Induced OPyC Failure .

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T.' ele-1--Description-of FSV Fuel-Samples Used in CHST Studies

. 1 Particle Type Fissile A Fertile B Fertile A Data Retrieval No. CU6A 6328 CT6B 932 CT6A 2399, Particle Parameters Kernel .

Density (gm/cm ) 9.19 8.90 8.89 Diameter (um) 180(23.1) 437(24.4) 374(35.1)

Buffer Density (gm/cm3 ) 1.10 *1.14 , 1.19 Thickness (pm) 54(11.5) 49(14.7) 50(14.5)

IPyC

' Density (gn/cm3 ) 1.89 1.91 1.89 Thickness (pm) 23(3.0) 35(7.7) 29(3.7) sic -

Density (gm/cm3 ) 3'.21 3.19 3.20 Thickness (pm) 28(2.5) 26(2.7) 27(2.7)

Ohc l 3 l Density (gm/cm ) 1.84 1.80 1.80 Thickness (pm) 39(4.5) 48(8.4) 41(6.2)

(*) Numbers in parentheses are standard deviations.

( . .. . .

I *

. 30 m

W .

Table 2. Description of FSV Segment VII Reload Fuel Fissile Fissile Fertile Fertile Particle Type A B A B :

Particle Parameters (a,b)

Kernel .

Density (gm/cm ) 8.90(.10) ' 9.01(.10) 8.83(.05) 8.81(.07)

Diameter (pm) 189.2(25) 251.2(16) 365.1(36) 446.0(36)

Buffer

~ '

Density (gm/cm )~ 1.13'([07)~ 1.19(.07) 1.11(.06) 1.06(.05)

Thickness (um) 54.8 (10) 51.6(8) 55.5(13) 53.3(12)

~-

IPyc Density (gm/cm ) 1.89(.02) 1.90(.03) 1.89(.03) 1.88(.03)

Thickness (pm) 25.5(4) 24.4(4) 27.1(5) 27.4(5) sic Density (gn/cm ) 3.20(.005) 3.21(.008) 3.20(.004) 3.20(.005)

Thickness (um) 25.1(3.6) 25.0(2.9) 24.7(3.1) 25.0(3.6)

OPyC ,

Density (gm/cm ) 1.86(.01) 1.81(.01) 1.81(.01) 1.83(.01)

Thickness (pm) 60.2(7) 41.?,(6) 51.8(8) 46.2(8)

(*}Hean values.

bambersin'parenthesesare~stindirddeviations.

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31

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l Table.3.._Ixradiation_ Conditions _of FSV__ Fuel..CilST :.imples.

Irradiation Conditions Fast Kernel Test Sample Burnup Exposure Fuel Data (%.FIMA)

Description Retrieval No. (1025 ,j,2) Fissile Fertile Fissile A CU6A-6328 9.1 18.2 NA(*}

Fertile A Cr6A-2399 9.1 NA 4.7 Fertile B CT6B-932 9.1 NA 4.7 FSV( )

4 Cord avg (*) NA 2.8 14 1.9 Avg. 6 yr fuel (d) NA 4.9 19 4.5 Maximum (* NA 8.0 22 7.4 (a) NA = not applicable.

(b) Irradiation exposures expected at equilibrium.

(c) Average over all fuel in the FSV reactor.

(d)__' Average conditions . for a. segment .remov.ed af.ter 6 y.rs_ residence.... *

(e) Maximum conditioni experienced after 6 yrs residence.

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Table 4. Key Stages of the CHST Method

. l -

Stage 3 Comments

1. Sample characterization e determine initial fission product! content I

. e contact x-radiograph samples I !

I j o -

4

2. Remove OPyC Layer (a) e heatinairto900C,holdfor2 hrs ;

e monitor fission product release i

- . l

/

w :

- (

'w . e contact x-radiograph samples s %.,s i t

3. Conduct CHST e heat approx. 100 particles from 1100 toI 2500'C in approx. 8 or approx. 30 hrs a I.

4 monitor fission product release i l

i

4. Post-test Characterization e determine final metallic fission produc,: content e

l j e visual assessment of particle condition. ,

I !

e contact x-radiograph samples ,

i i,

i 5. - Data Analysis . e' relate Kr-85 release data to total coating failure fraction j l:

I 3

' (a) Stage 2 was used to' prepare samples with " failed" OPyC layers.

i i

l w %9 .

Table 5.. Test Conditions During FSV Fuel CHST Studies i

1 l

Particle Type !FissileA Fertile B Fertile A Data Retrieval No. .

CU6A 6328 CT6B 932 CT6A 2399

~ l --- - - - - - - ~ ~ ~ ~ ~ ~ -

CHST Condition No. of Particles 88 100(*) -

101--- ~100 -- 99 (*} ~ 101 Temperature Range ( C) 1015- 1141- 1120- 1134- 1117- 1107-2400 2364 2475 2395' 2323 2410 Duration of CHST (HRS) 9.0 26.5 26.5 27.0 24.75 26.0 OPyC Condition Intact IIissing( } Intact liissing(b) Intact Missing (

(a) 33 particles removed from test at approximately 2000 C.

(b) Simulates fuel with failed OPyC layers.

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Table 6. Failura Fractions Suggested by CHST Data Collected on Individual Batches of FSV Cora 1 Production Fissile and Fertile Fuel ,

900 C 1100-1750 C 2050 C 2175 C *

'2300 C Data Kr-85 Kr-85 Kr-85 Kr-85 Failure (,) Release Failure (,) Release Failure g Release Kr-85 Phal Retrieval OPyC' Release Failure Release Fail (

Type No. Condition Fraction Fraction Fraction Fraction Fraction Fraction Fraction Fraction Fraction Fract Ficoilo A CU6A-6328(b) Intact <0.001 0 <0.001 0 <0.001 0. 0.02

_. 0,.018_ 0.095 0.1 7srtile A CT6A-2399I *) Intact <0.001 0 <0.001 0 0.070 0.07 0.85 0.85 1.0 1.C 7ertile B CT6B-932I ") Intact <0.001 0 <0.003 0 0.005 0.01 0.55 0.55 1.0 1.0 ,

l !

CU6A-6328 IC) Failed ( )

Ficcilo A <0.001 0 <0.001' 0 O.12 0.12 0.41 0 41 0.95 0.9 Fcrtils A CT6A-2399 IC) Failed I) <0.025 0.03 0.044- 0.06 0.092 0.09 0.35 0.35 0.95 '.9 0.058 hetilo B CT65-932I ") Failed I) 0.011 0.01 U.032- 0.05 0.076 0.08 0.47- 0.47 0.95' O.

0.046 (a) Total coating failure fraction.

(b) 9.0 hr CHST. c ,

(c) 24.75-27.0 hr CHST's. I (d) Removed after irradiation by burning in air at 900 C.

V _

's .

s .

I.

i

, Fable 7. Fatture TractionsI ") Dpected F37 Fissile and Fertile Fuel Irradiated to Average Conditions Fredicted for a

' 6 Yr Old Fuel Segment I .

a Failure Fraction Failure Fract on Failure Fragtion Failure Frgetion Failure . Fraction Number at 90 C .at 1100-175 C at 205 C at 217 C at 2300*C g Fuel Chc Particles Lage at Range at Range at Range at Range at Type Condition Tested Expected 90% Confidence Expected 90% Confidence Expected 90% Confidence Expected 90% confidence Expected 90% confidence Fissile Failed 100 0 0-0.03 0 0-0.03 0.12 0.06-0.18 0.41 0.30-0.55 0.95 0.80-1.0 i

- Fartile II Failed 201 0.020 0.01-0.05 0.050 0.03-0.09 0.085 0.05-0.13 0.41 0.38-0.55 0.95 0.80-1.0 t

M Ficcile & Intact 200 0 0-0.02 0 6-0.02 0.04'O 0.02-0.07 0.70 0.60-0.80 1.0 0.85-1.0 Fartile (b.d di 8

l a

h (2) Based on 24.75-27.0 hr CHST's.

, (b) Assumes A and B size fuel performance vill be the same.

. (c) Assumes fissile and fertile fuel with intact OPyC layers will be the same; does not include fissile fuel CHST data since they were obtained in a 9.0 hr CHST.

(d) Empirical f ailure model for a FSV core assuming no OPyC failure. .

4

i .* -

Table 8. Empi'rical Description of Expected, FSV Fuel Failure in a 6 Yr Old Segment i Failure Fraction (*)

Fissile (b') Fertile (c) Combined Fissile / Fertile (d)

G Temperature Expected Range (*} Expected Range (*) Expected Range (*)

( C) 900 0 0-0.02, 0.007 0.003-0.03 0.004 0.002-0.03 1100-1750 0 0-0.02 0.02 0.01-0.04 0.009 0.005-0.03

2050 0.04 0.02-0.08 0.06 0.03-0.09 0.05 0.03-0.08 2175 0.68 0.58-0.79 0.60 0.53-0.72 0.64 0.55-0.75 2300 1.0 0.85-1.0' O.98 0.83-1.0 0.99

0.84-1.0

)

(a) Normalized to expected OPyC failure fractions.

(b) Assumes 5.6% OPyC failure.

i (c) Assumes 34% OPyC failure.

(d) Accounts for fraction of fissions in fissile and fertile fuel.

(e) Assumes that all fuel failure fractions equal 5% or 95% confidence bounds.

i 2

. .~

p.- e w w .

e APPENDIX A CHST METHOD )

I l

. FSV production fuel was tested using standard methods developed on l the DOE sponsored accident condition test program, which is bettern known as the core heatup simulation:/ test (CHST) program, (Ref. Al- A3). Unbonded irradiated particles were heated from 1100 C to approximately.2500 C over j a period of approx. 8 or 30 hours3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months . Particle performance was monitored by measuring the fraction of Kr 85 and Cs-137 released as a function of time and temperature. Since fiss:Lon gas release is indicative of total coating failure, particle failure fractions were determined from Kr-85 release frac-i tions. Tests were conducted in a maximum of five stages. These stages are

. sussiarized in Table Al, and described below.

Stabe 1 includes collection and cha'racterization of the irradiated i ' fuel samples. Initial gamma counting is used to determine activities and activity ratios of key fission products (Cs-137, Ce-l' 4, 4 Ru-106) . The initial samma-count results are used to predict the Kr-85 inventory using the FISPROD code since Kr-85 is not detectable when gamma-counting intact irradiated fuel l

particles. Samples are contact X-radiographed to provide a permanent record of particle appearance prior to heating. - -

During stage 2, samples to be tested with missing (failed) OPyC layers are heated for two hours in air at 900 C in order ~to remove (burnoff)

~

f their OPyC layers (Ref. A3). During the burnback operation, the furnace l

atmosphere is periodically purged through a liquid nitrogen cold trap and the air replenished to provide an additional oxygen supply; subsequent gamma counting of the cold trap provides a measure of Kr-85 release due to pressure l

vessel failure resulting from removal of the OpyC coatings. Post-burnback saama counting and sample radiography is done to characterize the sample prior to the ,CHST and provide a measure of Cs-137 release that occurred i during the burnback operation.

G S e

S A-1 l . ..,.,_.m _ .,.y.. . . , _ - , . , . , - . . . . _ . . _ _ - _ - , _ . . . , , , . . . . - . . -

.,~,,,.m,- - .,,_...._,,,,.._--m..

l

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@ W .

During stage 3, samples are heated out-of pile in standard, resis-tance heated graphite tube (King) furnaces. Each test sample is separated into'3 equal groups of particles; each group is then placed into a. type H-451 sample holder prior to insertion in the furnace. A schematic of the test

. furnace / fission product release sampling system is shown in Fig. A1. Four Ta tubes are inserted into each furnace (only one is shown in' Fig. Al for simplicity). One tube ~ extends approximately half way through the furnace, is sealed on one end, and contains a temperature control thermocouple. The other 3 Ta tubes are open ended and extend through the furnace. One sample holder is placed in each of these tubes; a mullite Cs trap is also placed in each tube. Sample temperatures are monitored optically during testing.

~

Samples are heated to approximately 1100 C for 10-20 hours to simulate normal operating conditions prior to a core heatup and then heated from approxi-mately 1100 to 2500 C in approx. 8 or 30 hours3.472222e-4 days 0.00833 hours 4.960317e-5 weeks 1.1415e-5 months . A continuous flow of He (50 cc/ min /open ended Ta tube) is maintained during testing. Fission pro-duct gases released during heating are swept up by the flowing He and passed through a series of traps to remove tritium and radon. The remaining

- gases are then passed through two ionization chambers (vibrating reed electro =e-ters);the cold traps are changed at approximately 100-200 intervals and analyzed for Kr-85 by gamma counting. Results from the ionization chambers are normalized'to cold trap results to obtain a continuous measure of Kr-85 release as a function of tim'e and temperature'.

4 A mullite tube is placed at 1100 C on the downstreau side of each Ta tube to collect cesium released during testing. The mullite tubes are changed at approximately 70 C int'ervals during testing and gamma counted to

~

monitor Cs release as a function of time and temperature.

An optio occasionally chosen is to remove a portion (approx.1/3) ,

of the test sample at approx. 2000 C. Post-test characterization of the

- sample provides information needed to define failure mechanisms at elevated temperatures.

O e

M

'. A-2 .

4

.. . 9 9 .

t

'Camma' count results obtained during stage 4 are compared with pre-heating data to provide a measure of the release of additional metallic fission products (i.e. Ce, R, Zr, Eu). Post-heating radiography and metallography are used to illustrate the appearance and phase distributions within particles after heating.

g. . .

4

( -

l A-3 l

I

- -

W W . ,

6 .

APPENDIX A - REFERENCES A-1. "HTCR Fuels and Core Development Program Quarterly Progress Report for the Period Ending November 30, 1976," GA-A14180, December 27, 1976.

" Test Plan for FY-79 Core Heatup Simulation Tests", Document No.

k-2.

~~

903856, Issue A, December 7, 1978.

A-3. " Procedure for Burning Back Outer Pyr 'arbons of Irradiated Fuel

' -~ ~

Particles", Document No. 903878, Issur A, in review.

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. a. . .

MM

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66. . 9 @

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9 e e e e

A-4 4

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c .&". -

9 %9 .

o . .

(a) test configuration .

. i

^ x x x ' ' '

Test Cs He in >- Sample trap

~

"Y gfap c

_x x s .

x x x s ' x s .

resistance heated graphite tube furnace Ta tube (b) component description (c) test conditions e test sample N100 particles in e 1100 to 2500 C H-451 graphite crucible e linear increase in e Cs trap - mullite at N1100 C temperature uith time e Kr trap - liquid N C ld tr8P e 8 to 30 hrs per test 2

Figure A-1. Schematic of Test Configuration used for CHST Studies 4

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I A-5

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Table A-1. Summary of Steps Involved in FSV CHST Series Stage Description 1 Sample Characterization

1. Recover Test Sample

2. Camma Count Test Sample

3. Radiograph Test Sample

4. Do FISPROD calculationgs to Estimate Inventories of Fission Products Not Detected During Gamma Counting 2 OPyC Burnoff

. (only those samples to be tested with missing (failed) OPyC's) 1.

. Load Sample

2. Heat Sample in Air for 2 Hours at.900 C

3. Nbnitor Fission Product Release

4. Radiograph Test Sample 3 Core Heatup' Simulation Test

1. Load Sample into H-451 Holder

2. .

Load Furnace

3. Heat .

4. Monitor Fission Product Release !

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4 Post-Test Sample Characterization

1. Gamma Count Test Sample

2. Radiograph Test Sample

3. Perform Meta 11ographic Examination i 5 Analyze and Summarize Results - l e

A-6

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I ATTACHMENT 4 ENGINEERING EVALUATION OF LINER C0OLING WITH FIREWATER FOLLOWING A HIGH ENERGY LINE BREAK FROM 35 PERCENT POWER OPERATION EE-EQ-0019

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l A'l"l'AC i t M r.:NT 4 to I ,< ' t t o r l> -- 8 '3 4 (,()

PROTO-POWER c o a y a g , , , C. N

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v t e Letter Sequence Other
TAC:59787, (Approved, Closed)
Initiation Request, Request, Request Acceptance... Administration Meeting, Meeting, Meeting Questions 26 December 1985 11 October 1985 5 November 1985 Answers Results Approval, Approval, Approval, Approval Other: ML20137H372, ML20137L926, ML20137Z363, ML20138B512, ML20138L407, ML20141E527, ML20141H534, ML20141H544, ML20141H549, ML20151R074, ML20154N150
MONTHYEAR1985-12-31 [Table View]

ML20138B512

Text

SUBJECT:

REFERENCE:

Dear Mr. Berkow:

SUMMARY

SUMMARY

SUMMARY

SUMMARY

2.0 INTRODUCTION

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