ML20202D849

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Rev 0 to Topical Rept 116, Leakage Assessment Methodology for TMI OTSG Kinetic Expansion Examination
ML20202D849
Person / Time
Site: Three Mile Island Constellation icon.png
Issue date: 11/06/1997
From: Keasler R, Odonnell E
GENERAL PUBLIC UTILITIES CORP.
To:
Shared Package
ML20202D831 List:
References
116, 116-R, 116-R00, NUDOCS 9712050169
Download: ML20202D849 (44)
v • d • e


Text

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LEAKAGE ASSESSMENT METHODOLOGY FOR 1MI OTSG KINETIC EXPANSION EXAMINATION TOPICAL REPORT # 116 REV.0 AU1110R:

A.  ;

DATE: _/] I/

APPROVALS:

/ U Sff.*?

DEPARTMENT DIRECTOR DATE IdRECfol( ESGINEERING l

6ATs f

(SIGNITICAW IMPACT REVIEV!)

9712050169 971126 PDR ADOCK 05000289 P PDR

uncer ass:swrur wrecaecuca rw one == re reserm CwruDet0 LEAKAGE ASSESSMENT METHODOLOGY FOR TMI OTSG KINETIC EXPANSION EXAMINATION 1.0 I NTRO D U CTIO N A N D B AC KG RO U N D ....................................................... 3 2.0 OVE RVI EW O F M ETH ODO LOGY................................................................ 4 3.0 t /. AIN STEAM LIN E B REA K A N ALYSIS..................................................... 6 3.1 OVERVIEW..........................................................................................6 3.2 S H O RT T E R M ANALY S I S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

3. 2.1 B a nis o f Dura tion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6
3. 2. 2 M a o dolo.rv. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ......7 ..............
3. 2. 3 A s s ump tion s. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ....8
3. 2. 3.1 I n itial C on d itions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8
3. 2 . 3. 2 B r e a k M od cli n g . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ...8
3. 2.3.3 Re a cto r Ve s sel M ixin g . . . . . . . . . . . . . . . . .. .. . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . .. . . . . . . ... . . . 9 3.2.3.4 Rea ctor Kinetics Pa ra meters .. . . . . . ..... . .. ..... ... ... ... . . . .. ... ... ........ 10
3. 2. 3. 5 R e a ct o r Tri p . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 3.2.3.6 initial Steam Generator Mass . .. ......... . ...... ......................... ... 12 3.2.3.7 Main Feedwater and Emergency Feedwater Flow ......... . . ... ..12
3. 2.3.8 Hig h P re ssu re I nje ction . . . . . . . . . . . . . . . . . . . .. . . . . . . . . .. . . . . . . . . . .. . . . . . . . . . . . . . . . 15 3.2.3.9 Steam Generator Dowocomer Modeling . ............ .... ... ..... .. .. 15
3. 2. 4 Summa ry af Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 3 . 2 . 4 .1 P o w e r . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 5 3.2.4.2 Loo p Te m pe rat u re s . . . . . . . . . . . . . . . . . .. . .. . . . . . . . . . . . . . . . . . . . ... .... . ...... 1 6 3.2.4. 3 OTS G P re ss u re . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17
3. 2.4.4 R C S P re ss o re . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17
3. 3 LO N G TE RM AN ALYSIS . . . . . . . . . . . . . . . . . . . . . , . . . . . . . . . . . . . . . . . . . . . . . . . .. . . ... 1 9
3. 3. 1 Appro a ch . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .............................19
3. 3. 2 A s s ump tion s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 3.3.3 UTSG Cooldown Analysis..... ... . . .. ... ........................20
3. 3. 4 Re s ults . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 1
4. 0 O T S G T U B E L O A D S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . . . . . . . . . ... . . . . . 2 3 4 .1 I NT R O D U CTI O N . . . . . . . . . . . . . . . . . . , , . . . . . . . . . . . . . . . . . . . . . . . . . ... ... ....... ..... .. ..... 2 3 4 . 2 M ETid O DO L O GY . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 3
4. 2. I G PUN Me th odology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23
4. 2. 2 F TI Meth odology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24 4 . 3 RE SULT S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .......... ........ ............. ........ 2 5
4. 3.1 GPUN O TS G Tubo L oads.. . . . . .. . . . . . . . .. . . ... ..... . . . . .... . .. . . . . . .. . .. . . .. . . . ... 25
4. 3. 2 F Tl O TS G Tub e L o a ds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 5 4 . 4 ANALY S I S O F LOAD S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 7 5.0 C RAC K AR EA D ETE RMIN ATIO N ............................................................. 2 8 l

1

LLAKAGE ASSESSMENT METH000t0CY FOR TMt OT50 kwtTC EXPANSOV ExAMnATION utaso10

5.1 INTRODUCTION

....... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ..... . ... .. 2 8 5.2 METHODOLOGY (KINETIC EXPANSION REGION) ........ .... . . ......... ..... . . . 29 5.2.1 Circumferential Through-Wall Crak in Tension... ............ ..... ... .. . 29 5.2.2 Axial Through-Wall Crack Subjected to Intemal Pressure .. . ... . . 29 6.0 C R.AC K A RE A LEAKAG E AN ALYSl S....................................................... 30 6.1 OVERVIEW................................................................................ . 30 C O D E D E SC RI PTI O N . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31

$. e LEAKAG E RE DUCTION FACTOR . , . . ... . . . . . . . . . . . . . .... . . . . .. . . . . . . .. . . . . . . . . . .. . . . . . 31 0 . . CONTACT PRE SSU RE D ETERMINATIO N . . . .. . . . . .... . . . . .. ... .. . . . . . . . . . . . . . . . . . . . . . . . . 3 3 7.9 TOTA L LEA KAG E EVA L U ATIO N .............................................................. 3 6 7 .1 OVE RVI EW . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 6

7. 2 LEAKAG E R E SU LT S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 6 7.3 AFFECTED OTSG VS UNAFFECTED OTSG........... .. .. ................. ... .. ..... 37 8.0 S U M M A RY A N D C O N C LU S t O N S .............................................................. 41

9.0 REFERENCES

............................................................................................42 2

a~nusuar ---rgormm,c enn au-~

1.e .JRODUCTION AND BACKGROUND GPUN agreed to do inspections within the upper tubesheets during the 12R refueling outage and to define acceptance criteria for these inspections. Four sets of criteria were developed for use during the 12R eddy current testing (rep) inspection as follows:

1. Structural Integrity A. Minimum intact expansion determination t B. Maximum acceptable def9ct determinatioa
2. NDE performance evaluation
3. Material condition assessment
4. Primary-to-secondary leakage analysis The first three of the above items were discussed in an August 08, 1997 NRC submittal (Reference 1) . The purpose of this report is to describe the methodology that is being used to evaluate the total primary-to-secondary leakage that may occur during a guillotine rupture of a main steamline as a result of assumed through-wall (>67.4% ECT indication) cracks in the kir.;ic expansion region of the OTSG tubes.

In Reference 17 it was demonstrated that the limiting accident scenario which reaults in the largest tube loadings is that which results in a large SG tube to shell temperature differential (AT). The most restrictive limits were deterriced to be when the tubes are colder than the steam 91 are nr shell. Consequently, in Reference 1, it was concluded ' rot the MSLB accident results in the largest tube to shell at In order to establish the total primary-to-secondary leakage that would be acceptable during the MSLB event from assumed through-wall cracks in the kinetic expansion region, a calculation was done to determine the maximum leakage that would meet the offsite dose criteria of 10% of 10 CFR 100 limits for the 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> Exclusion Area Boundary (EAB) and 30 day Low Population Zone (LPZ) (Reference 2). The revised dose consequences for the FSAR MSLB analysis were Pubmitted to the NRC for approval (Reference 3). The results are as follows:

1. Integrated Primary Coclant Leakage @ 2 hrs (gallons @

579 F) = 3228

2. Total Integrated Primary Coolant Leakage (gallons @

579 F) = 9960 3

LEAKAGE A55E5$htEn{T neETMCCOLCCY FC@ TD OT5C KINETC Exh*.NSON EKLwNATION twum0 The methodology used to determine if thene leakage limits are met is discussed in the following sections. Section 2.0 provides an overview of the methodology and the subsequent sections provide some additional details.

2.0 OVERVIEW OF METHODOLOGY The methodology involves the following activities that are depicted in Figure 2-1:

A. Develop the time varying thermal hydraulic (T-H) information from the design basis Main Steamline Break (MSLB) event analysis.

B. Determine the OTSG tube tensile and differential p essure loads from the T-H data. The loads will vary as a function of time throughout the transient and as a function of radial distance from the center tube to the peripheral tube.

c. Calculate the theoretical crack opening area separately for circumferential and axial cracks. This will vary with the applied load and crack length.

D. Determine the theoretical leakage flow as a function of the crack area. Based upon the expansion contact pressure at the flaw location, determine if leakage reduction factor can be applied to the crack area leakage flow. The total mass released from the crack is obtained by integrating the leakage flow over the first 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and over the entire transient interval.

E. The integrated leakage flow for each of the identified cracks is summed based on crack size and radial position and the total is compared against the leakage limits specified in the offsite dose calculation (Reference 2) based upon 10% of 10 CFR 100 limits.

If the total leakage exceeds the limits established in Reference 3, then a decision wil1 be made as to which tube (s) will be repaired such that the leakage contribution from the repaired tube (s) can be eliminated from the total to meet the allowed leakage limits.

Additional details and references regarding each of the activities discu. sed above will be provided in the sections which follow this overview.

4

unace usessurm unmocrg ors mue emm uaun rm FIGURE 2-1 LEAKAGE EVALUATION METIlODOLOGY OVERVIEW CALC LEAKAGE DETERMINE OTSG FLOW AREA DEVELOP T-il TUBE LOADS FOR CRACK CONDITIONS FOR . ,

MSLB TRANSIENT FROM MSLil F(TUBE LO4D. CRACK F(TIAIEj T li CONDITIONS LENTIL, & CK4CK F(TIAIE.POSITIONIN OTSG) ORIENTA TIONJ A

INCREAIEAT TO NEXT CR.4CK h

CALCULATE CONTACT PRESSURE AS F(llEIGIIT) CALC INTEGRATED IN KINETIC LEAKAGE FLOW .

Toi,.L LEAKAGE EXPANSION REGION FOR EACil CRACK ^ '

COMPILATION F(.4 REA. T-H)

APPLY v LEAKAGE REDUCTION DEVELOP LEAKAGE REDUCTION FACTOR -

COMPARE TOTAL FlCONTACTPRESSURE)

^^ ^ A 7

! EAB AND LPZ s

LlAKACE AS$255 MENT METH000LC Tw OTSC KsNEnc E9AN5JO4(KAMhAEDN Ytev FC,sson 0 3.0 MAIN STEAM LINE BREAK ANALYSIS 3.1 Overview The MSLB analysis is used to generate the transient thermal hydraulic parameters that are needed as input to define the OTSG tube loads, to calculate the leakage from each crack and to determine the contact pressure as a function of axial position above the kinetic expansion transition region in the tube sheet for determination of the appropriate leakage reduction factor.

The analysis is accomplished in two phases: a short term and a long tern. The short term phase duration is 10 minutes (600 sec) and utilizes the transient systems analysis code RETRAN-02, Mod 5 (Reference 5). The long term phase thermal hydraulic conditions are developed by applying assumed operator actions, based upon TMI-1 Anticipated Transient Procedures (A1Ps), to recover from the event and to calculate the OTSG shell metal cooldown rate in order to develop a technical basis for cooling down to DHR conditions without violating tube to shell differential temperature limits. The long term analysis begins at 10 minutes and extends to the end of the transient (approximately 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />).

Details of these evaluations are provided below.

3.2 Short Term Analysis 3.2.1 Basis of Duration As indicated above, this portion of the MSLB thermal hydraulic analysis includes the first 10 minutes (600 sec) of the event. There are multiple reasons for choosing this time frame. First, this portion of the transient is characterized by the most complicated and dynamically changing thermal hydraulic attributes. The affected OTSG is blowing down, the Heat Sink Protection System initiates a closure of the Main Feed Water (MFW) control valve and the MFW block valve and also initiates Emergency Feed Water (EFW) on low OTSG level. The RCS is depressurizing and cooling down, the pressurizer is. emptying and refilling, an RPS trip occurs , ESFAS is initiated, etc. Because of the complexity of this portion of the transient, a sophisticated systems analysis code (RETRAN 02, Mod 5) is used to ectablish the thermal hydraulic parameters during this period (Reference 5).

Another reason for this duration is that no operator recovery actions are assumed to take place until after 10 minutes has passed. This is a licensing basis for TMI.

6

~ - - . . . . - - -. . ._ . . --. - . - - . - - _ ~

isacceassessww wrmocnor tw orso mere temsa tuusarm etytpon 0 Following the first 10 minutes, credit for operator actions

- is permitted.

The peak loads _for this event also occur within the first 10 minutes and the thermal hydraulic conditions at the end of- '

this duration are important since they represent the end of the peak _ load period and the transition to reduced OTSG tube loads.

In this manner, the first 10 minutes of the MSLB analysis-

- sets the stage for the entire leakage determination effort.

At the end of this period, the system is not characterized by-rapid changes in thermal hydraulic conditions and is-in transition to the recovery from the event.

3.2.2 Methodology The RETRAN-02 MOD 005 computer code and a TMI plant model were used to perform this analysis (Reference 4). The TMI RETRAN model has been extensively benchmarked against plant data and previously approved licensing codes. The benchmarks demonstrate the adequacy of the TMI RETRAN model for performing safety analysis. The TMI RETRAN model has also been approved by the NRC for referencing in licensing applications (Reference 5). The TMI base deck (Reference 6)

STEAM YJ209 B h

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@ THREE WILE saAND UNIT 1 res g b"*n%"i" 2 res O .

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LELKQCf ASSESSMENT METrQ2DC10GY FC2 iMt OTSG ENETC LK9th$ ON (KDMINaTION tevsoor10

, as shown in Figure 3-1 was used for this analysis.

3.2.3 Assumptions The analysis assumptions and initial conditions as discussed below were chosen to provide a conservative RCS overcooling and pressure history for the MSLB event and the resulting tube loads.

3.2.3.1 Initial Conditions The reactor is assumed to be operating at rated power prior to the accident (2568 MWt). The initial pressurizer liquid level is set to 220 temperature-compensated inches, which is the typical hot full power (HFP) pressurizer level. The initial cold leg temperature was 555.0'F, which is the nominal cold leg temperature. The initial RCS pressure was 2170 psia in the hot leg, which is the normal operating value. The TMI design basis MSLB assumes that offsite power is available and that is the assumption in this analysis.

The effect of high RCS loop flow is to minimize the OTSG tube average temperature during the initial phase of the event. Thus, OTSG tube axial loads are maximized.

3.2.3.2 Break Modeling The initiating event is assumed to be a double-ended rupture of a 24 inch steam line on one steam generator.

This is the largest possible break which results in the maximum cooldown rate. The faulted steam generator steam line was nodalized as shown in Figure 3-2, so as to model each steam line individually. The flow area of the two break junctions are consistent with the 24 inch steam line piping.

A Moody choking model is used for these break junctions with a contraction coefficient of 1.0 to maximize break flow rate.

The break occurs in the int'ermediate building upstream of the MSIV. This is an appropriate break location because it results in a ground level release of coolant activity.

8

unce assessaw urnmootocnor w orso mrre exoamov exauanxw Revneon 0 FIGURE 3' #

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\@ MVEFu VALVt Fu 3.2.3.3 Reactor Vessel Mixing The amount of mixing that occurs within the vessel is a ratio of the difference in hot lag temperatures to the

'ddfference in cold leg temperatures.

T,ot Nnf aulted - T 3 ,, (f auhed RATIO - =

T,,1, (unfaulted) - T,oi, (faulted)

-A value of-RATIO = 0.0 im' plies perfect mixing while 1.Oc. implies-no mixing. For the purposes of this-analysis, a target value of RATIO = 0.5 was chosen to conservatively bound the. analyses at an upper value.

To simulate-this mixing in RETRAN, the reactor vessel-was modified;to include two equal' parallel flow paths by splitting the downcomer, the lower plenum, the core, 9

1

-, , ,_la. .,_[. '

umcr asseswivr urwxxxocuca rut orse merc imwsa ruwerm 20vm 0 and the upper plenum as shown in Figure 3-3. For the most part, these parallel flow paths behave independently, with the exception of common connections with the bypass and upper head volumes. These common flow paths keep the loop pressures in balance but contribute little to mixing of loop flows.

3.2.3.4 Reactor Kinetics Parameters To minimize the power increase response to the core temperature decrease, the moderator temperature coefficient (MTC) used was a value of zero. This is conservative since it will not increase the power prior to trip and results in lower RCS temperatures. Post trip, the MTC determines the extent to which the core energy generation is increased by sub-critical multiplication. An MTC of zero will assure that the post trip reduction in temperature will not lead to increases in power generation above the normal decay heat power. The absence of a return to power after the trip results in a greater cooldown.

Decay heat is based on the ANS5.1 1979 decay heat standard. In order to maximize RCS cooldown following reactor trip, a 0.95 multiplier on decay heat was used.

The 5% reduction was chosen since it is greater than a 2c uncertainty for thermal fission of U* under equilibrium operatir.g conditions .

i 10

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samaat assesswur wrmoovne ru orso mere tensov tunnu tow -

awsen o .

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uence ysessww wr.emocygorso mr= remsa rmawrm 3.2.3.5 Reador Trip -

With1an'MTC-of zero, the power will not. increase with  !

y~ the decrease in-moderator temperature, so the reactor -  ;

-will' trip _onglow RC pressure. Since this analysis is primarily interested in tube temperatures, a trip setpoint of 1900 psig plus a 30. psi error was used.

This limits the_ amount cf energy-the core model generates, resulting in a lower primary system l temperature during-the-event. It should be noted that -

this setpoint results.in an earlier trip which is

, conservative for_ tube temperature calculations. For the steam line break event, the trip setpoint will be--

reached rapidly due to the dramatic overcooling which 4 is occurring, ,

3.2.3.6 initial Stearn Generator Mass The_ initial steam generator inventory provides a measure of the heat removal capability of the secondary system. For a steamline break, a larger initial secondary system inventory in the steam generator associated with the break will lead to a higher integrated heat removal. The larger the heat removal, the lower the resultant reactor coolant temperature.

The-(OTSG) design has the maximum inventory at full power conditions. Thus,.the event should start from full power to maximize the heat removal capability of the steam generator. The steam generator inventory can increase.if fouling of the SG tube bundle region occurs. The inventory predicted for full power and fouled _ conditions has been conservatively determined to

, be approximately'55,000 pounds per SG. In addition, the mass of feedwater between the isolation valves and the affected' steam generator which was calculated to be 35,500.lbm was also modeled and available to cool the i

affected steam generator, a 3.2.3.7 Main Feedwater and Emer0ency Feedwater Flow The MSLB accident in this calculation assumes the worst

' single failure, which is the failure of the feedwater

regulating valve-to close on the affected generator.

This_ maximizes the-overcooling of the event by maximizing the main feedvater i.

12

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- umor essasswur wrmocnor rm orsa aure temsm uumorew hmsson 0 FIGURE 3 ,

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> uemwso - 1 A

I I I I i e m m m m = .a -'

no eein. tow) i (MFW) flow to the affected generator as a result of the preferential' feeding'to the broken- depressurized, ,

side. Feedwater f. Low to the affected steam generator is shown on Figure 3-4 above. MFW flow is terminated to the affected. steam generator after-the MFW block valve

- closes in about-30 seconds after a low SG_ pressure of 600psig is reached.

For this transient, the EFW system would be initiated by_a low OTSG lev'el signal. The OTSG low .

' level initiation signal of 10 inches is measured by the

.startup: range-instruments. The setpoint is calculated

- in the RETRAN-model as-the collapsed-liquid level in the tube region. Zero inches indicated level .is 6 inches above the-lower tube sheet. EFW controls level at:25.. inches-indicated. Due to the continued MFW flow

- to the broken SG until the-MFW block valve closes, the 4

13 4

.3

- s,,, a _.-a. _ .---_ _ __m ,m.- a- -- - -- ._ _____eo n __ . _ _ _ _ _- - -

m.

y MMAGf A55f55WWF WDCDROCY NW fM OT3C mrTC fpw$m fumWTM -

etw> cts 3 :

1 Figure 3-51 asse,srp eista nn amm E -

s,  ;

I, .

j y 1 -

ameso E-W E .

I  :- ineso O

i i i i i

< u. m m- m m m s w ees u ne (seel OTSG level'does not drop below the low level initiation signal until.about 67 seconds.

The start of the motor driven EFW pumps (MDP) -is delayed by 5 seconds after the initiation signal having a coastup time of 10 seconds. Subsequent to the EFW initiation-signal, the steam admission-valve to the-

. turbine - driven = pump- (TDP) , MS-V13A,- receives an immediate open signal and.is fully;open in 24 seconds.

Turbine: testing shows the TOPS are at_ full speed in-11 seconds lafter the steam admission' valves are full.open.

JAn additional 8' seconds for flow coastup is modeled resulting-in TDP flow. delivery at 43 seconds.

For-this analysis, 2-MDPs and TDP.were! conservatively F ' assumed to deliver flow instantaneously to the steam

- generator following an EFW initiation signal (See 1-Figure-3-5 above).

14

uance assassums urr,emocrggrsc menc exmsm ruum rm 3.2.3.8 High Pressure Injection The high pressure injection (HPI) system is actuated during the cooldown period following a large area steam line break. The system supplies borated water to the RCS to recover the RCS shrink and to provide core cooling if necessary, and to increase the core shutdown margin. Baron addition to the reactor coolant, during the controlled cooling to atmospheric pressure, will prevent criticality at lower temperatures. For this analysis, no credit was taken for boron addition resulting from HPI actuation, since the BOL kinetics and best-estimate rod worth will result in keeping the core ahutdown. To minimize the primary system temperature, and thus tube temperatures, full HPI from three trains was conservatively assumed. The HPI will be initiated on a signal of 1600 psig plus a 30 psi error at the pressure measurement tap location. This is conservative, since a rapid actuation of HPI will maximize the overcooling.

3.2.3.9 Steam Generator Downcomer Modeling The RCS cooldown is maximized by minimizing the amount of liquid carried over from the steam generater out of the break. To minimize the liquid carryover, the downcomer was modeled with a single bubble rise volume and a large bubble velocity (1x10E6 ft/sec) which produced less liquid carryover.

, 3.2.4 Summary of Results 3.2.4.1 Power The results of the MSLB analysis for the first 10 minutes (600 sec) are provided in this section. The reactor scram occurs on low reactor pressure in about 10 seconds as shown in Figure 3-6. This reflects a trip setpoint of 1900 psig plus a 30 psi error.

The reactor power in Figure 3-6 also indicates-that there is no return to power as a recult of the absence of a negative moderator temperature feedback. This is a conservative result with respect to the cooldown.

15

. . . , . - . . ...- ~. - , . - .- - . -..

s &-

umcs usasseur wrea ocnor to onc awrc romsew annammw

  1. wesen 0 -)

i 3 .

_ Figure-3-6l

_ _ eeuw new

.5 y

-j g.

1 l- :_

I k

i i i i i i, n. . a. m m I

slapsed flee teac) 4 k

3.2A.2 Loop Temperatures -

The hot and cold leg' temperature response to the MSLB-are shown in Figure 3-7. A rapid overcooling results from the ,

event with the cold leg temperature reaching about 435 deg F in aboutc70 seconds after'the break. After the OTSG blowdown

'is completed,-the primary to secondary. heat' transfer is reduced and the-cold ~1eg and hot' leg temperatures are essentially the same. . The-temperature:-is about 450 deej F at this point and is maintained for the duration of this portion-of the event.-The final temperature for~this phase of the-event-reflects the' fact-that the' intact OTSG acts'as

-a, heat source as discussed below.

l W

16

<m,. .m.-* < -,. a ,,-- =w.e, *--u W

L uus,r urss==r wrenmoonee rm orso nvere r,=scw ruamerm -  ;

Nmor10 {

-Figure-3-7 d

.. f

- bcf feeltet tasy T.sp.r.tese. j e- .

s_ ,

g

> mg i l a.

Cad t.o e.

t 8 -

i i i

m

. m .i .i .= m it a tw i i 3.2.4.3 OTSG Pressure The: pressure response for both the faulted and unfaulted OTSG'is shown in Figure 3-8. The faulted OTSG is fully depressurized-in about 100 seconds.

The unfaulted OTSG responds initially in a normal post trip manner, increasing to the MSSV setpoint, but is_ slowly reduced _in pressure as a-result of_ reverse heat transfer to the RCS.

- 3.2.4.4 RCS Pressure .

The RCS pressure is depicted in Figure 3-9 and reflects a rapid drop in pressure due. to the ^ initial cc,oldown. The drop -

in.~ pressure results in-a reactor trip,-ESAS_ actuation and a sma11 influx of, core flood tank' flow. After the cooldown has 1.:

stabilized, the RCS repressurizes in response to HPI injection-flow refilling the pressurizer. At the end of 10 minutes, the RCS subcooling margin is less than 100 deg F. ,.

17

f' .- __ _ _ . -. _ . . . _ . _ - . _ . -. ._. _ _ _ _ - - . - .

truact assessueur ucraaocs roe tw orso entre uxnma sausarm ermon o Figure 3-8

e nta naar.anv==

.3 I~ ,, sed S0 2

5 E- ,

1 _

g I-

> emers E.

I i i i i i, ,i,, m i., .a w **

31 4 et= tuo Figure 3-9 n== tin naar.

E ',

I g.

A g.

I-e ;se ses see 4:a m Sc3 Blaseed time I.Mi 18

ummserwraaocnor tw enc mcwasa naarm Nmwn0 1

' 3.3 Long Term Analysis '

3.3.1 Approach Following the first-ten minutes, it was assumed that '

operator _ action would be.taken to terminate EFW to the affected OTSG and to begin a controlled cooldown and depressurization to DHR conditions using the unaffected

'OTSG..The limitations imposed by the various cooldown P-T limits- and tube to shel2 dif ferential temperature limits would be observed. The following assumptions reflect this approach,

- 3.3.2 Assumptions

1. The operator will control the NSSS such that the tube to shell differential temperatu.' ret.eile limit of

--70'F (tube temp minus shell temp,in observed

-(Reference 9).

2 RCS temperature will not be allowed to increase to reduce the tube to shell differential temperature (Reference 10). Procedure guidance has the operator minimize the RCS reheat following an overcooling event.

Increasing RCS temperature'for this analysis would reduce (make less negative) the tube to shell differential temperature and reduce the tube load.

Reduced tube load would lead to reduced tube leakage.

3 RCS pressure will be maintained at a subcooled margin of 75'F. Reference 10 directs the_ operator to minimize the RCS pressure increase-following an overcooling event. The minimum SCM limit is 25'F (Reference 9).

An RCS pressure control value of 75*F SCM is reasonable. Higher RCS pressure _ leads to greater tube leakage.

4. As RCS temperature _and pressure decrease, additional pressure limitations are established. The operator will maintain RCS pressure in-excess of the emergency RCP_NPSH limit (Reference 9). A margin of 50 psi is considered.to be adequate. A high margin maintains RCS pressure high,_ increasing tube leakage. However, a

. large margin to the NPSH curve could prevent initiation of DHR. Therefore, a margin.of 50. psi is reasonable, Additionally, the operators will maintain RCS pressure such that the minimum RCP seal differential pressure (275 psid) is maintained (Reference 12). Seal return can be dumped to'the sump instead of being sent to the

. Makeup Tank.___A margin of 25 psig is maintained to the 19

trau,r usesswur urr,emem ca to orso mene causa reumra tms2n 0 limit of 275 psid. Therefore, a minimum RCS pressure of 300 psig is established.

5. The transier.t after 600 seconds is quasi-steady-state.

Therefore, large time steps can be used. A time step size of 600 see was cho.en as reasonable.

6. Operator action is assumed to take place at 10 minutes.

The following actions would be taken by the_ operator for a MSLB event (References 9 and 10):

a. Terminate EFW to the broken OTSG (MFW is already isolated).
b. Control / terminate HPI to the RCS to control RCS pressure,
c. Adjust the TBV on the Unbroken CTSG to prevent RCS temperature from increasing.
d. Terminate EFW and initiate MFW to the Unbroken OTSG.

3.3.3 OTSG Cooldown Analysis As indicated above, the operatir will control the NSSS such that the tube to chell differential temperature tensile limit of -70 F is ooserved. The maximum possible cooldown rate that meets this criterion is established by the rate at which the affected OTSG shell cools down.

To determine the shell cooldown rate, the GOTHIC computer code, version 5.0e, was used with a six(6) volume model as shown in figure 3-10 (Reference 11). Two volumes (volumes 1 and 2) represent the primary (tube) side of the OTSG, two volumes (volumes 3 and 4) represent the secondary (steam side) side of the OTSG shell inside the shroud, and two volumes (volumes 5 and 6) represent the secondary side of the OTSG outside the shroud (between the shroud and the shell metal). The volumes are divided to correlate with the division of the downcomer region into upper downcomer and lower downcomer regions.

The analysis began at 10 minutes and allow S the RCS to cool down as the shell cooled down to preserve the -70 deg limit and thus acccunt for the impact of the cooler RCS tube temperature on the cooldown rate of the shell.

The shell cooldown rate results from this analysis are shown in Figure 3-11 in section 3.3.4 below.

I 20

tsaxe asstssuum utvecnocr vor rm onc mirx rxasov sxcawwx Revno10 Figure 3-10 GOTHIC Model For Shell Cooldown Analysis 4F

. AT

  • a b

,A

._ 9_, g_-_, -

g,

-l -

I I 1 I I 1 I I I I I I I I . I l

.I l 1 M i1 I I, 1 I

i 1

l l I I l I I sp I 1 I l , I I:

ICE" _", h h :q 5, b  : :. d E M IG

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$$d - I T$O$ I 1

l i I I i l I i i

l i I l i I I I l t i i I l u_q;f J u _ _ _ _: ^

m _ _I BP I

l 3.3.4 Results Figures 3-11 and 3-12 below provide the results of the long term analysis. The figures also include the first 600 seconds as well. The results reflect the application of the criteria described above. The average shell temperature is a weighted average of the upper and lower shell temperatures

at the outside metal surface of the OTSG. The RCS l temperature is the average of the hot and cold leg temperatures for the affected OTSG.

21

armcs erstssurvr urrectocr m to orse awrre rmwsov txmaarov reseno Figure 3-11 MSLB Temperature Response

  • *
  • RC51emperatse su l Average $heO

.00 . . . . . . . . .. ,,,

sao ', t n . _ _ _ - -- . - .. _

s,_ . - .

7

.._.._C__%....,.,

q %e

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m - .- - - . .- ..-.- - ---

u ase _ _ - _ _ _ _ _ . _ . _ . _ _ . . . _ _

100 - - - - - - - - - - - - - - - - - - - - - - - - -

gsq -.--- - - _ . . . ------- - --.- - -.-. - - ---- -

300 1 10 100 1000 10000 100000 in. a..o Figure 3-12 MSLB Pressbre Response stoo

,...****. * *

  • RCS Presswo 2000 - - - - - - - - - - - - - - - -* * . - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - ~ -

js@Q ___,.-___-.__4.__,_-.____._._.-.. .

1800 - - - - - - - - - - - - - - - - - - - - - - - -- - - - - - - - - - - - - - - - - - - - - - - - - - - - -

,3cc _ _ _ _ _ . . . . ~ _ _ _ . . . _g__ _ , _ . - . _ . . _ _ _ _ - _

i ,a e - _ . _ - .

f.

~~..

800 ,

--.---.-g

$0$ ~ ~ - - - -

L

?O -

230 - - - - - - - - - - - - - . - - -- -

0 1 to 100 1000 10000 100000 Tem. (s.c) 22 I

l i auccr assessym urwemocy no rv, orse merc emusa ravema tevtsson 0 4.0 OTSG TUBE LOADS 4.1 Introduction The OTSG tube loads are determined from the T-H parameters provided from the analysis presented in Section 3.0 above.

The method for calculating the tube loads evaluates the theoretical tubesheet deflection under a differential pressure and tube axial load and as a function of the different OTSG tube, tubesheet and shell metal temperatures.

The resulting pressure and tensile loads are used to determine the leakage area that develops for a given crack length and crientation as described in Section 5.0 below.

Since the thermal hydraulic conditions are changing with time, the resulting tube loads are also changing accordingly. As a consequence of the tubesheet deflection from the center to the periphery, the tube loads will vary as a function of thc radial distance from the center of the OTSG, In this way, a plot of the tube loads as a function of radial distance from the OTSG center to the OTSG periphery would be different for each set of consistent T-H conditions. The discussion below provides an overview of the methodology used by both GPUN and FTI to independently determine the OTSG tube loads using the T-H data in Section 3 and a presentation of the results.

4.2 Methodology 4.2.1 GPUN Methodology The methodology that was employed by GPUN in the determination of the tube lcads is described in Reference 16 and comprises the following steps:

  • Establish the tubesheet behavior as a function of applied load and material properties as a function of temperature.
  • Establish the tube loading (pre-load) in the OTSG as a function of the measured gap between the separated sectiona of a failed tube at the temperature at the time of measurement.. The calculation will be based on the assumption that very few tubes have parted so that the loading on the balance of the intact tubes is unchanged.
  • Separate the three major OTSG components (tubes, shell, and tubesheet) to free components (bodies),

remove all loads acting on them and find their unloaded geometry.

i 23

um:t assessueur uno<mocqrgrsc mac emsm enuwm

  • Establish the physical variables that will result in deformation of the free bodies and calculate these deformations, including an accounting for the Poisson effect on the tubes and on the shell.
  • Re-combine the deformed free components by pulling the tubes until they meet the final tubesheet location. The final tubesheet location must simultaneously satisfy both of the following conditions:
  • The tubesheet periphery must be at the same location as the shell.
  • The tensile load from all of the tubes must be equal to the shell compressive load.

4.2.2 FTl Methodology An ANSYS finite element model of the OTSG was used to determine the tube load contribution for various system operating parameters. The ANSYS model is basically identical to the NASTRAN model used in the 'OTSG Tube Topical Report,' Reference (17). The NASTRAN model was converted to ANSYS due to some extra features ANSYS possessed at the time.

The model is an axisymmetric thermal and structural model of the OTSG. The model includes the steam generator shell sections, upper and lower heads, upper and lower tubesheets, support skirt, and twelve beams representing twelve effective tube regions. The tubesheet model accounts for the material properties which were adjusted to account for the tubesheet temperature and the effects of the perforated plate.

Several different load cases (parameter study) were executed to establish the variation in tube loads due to changes in primary pressure, secondary pressure, tube to shell AT (both tubes hotter and cooler than the shell) and average tube temperature. The end result was a series of equations, as a function of average temperature and tubesheet radius, that provides the load in the tubes for each of the pertinent system parameters.

Using the postulated MSLB system transient parameters discussed in Section 3.0 above, the total tube loads for the transient, as a function of transient time and tubesheet radius, were determined.

24

uuxu usessww wrmaocv na rw msc twc tame u1mera

  1. wascio
4.3 Results

~ 4.3.1 GPUN OTSG Tube Losds The GPUN analysts results are provided in Figure 4-1. This shows the.OTSG tube loads for three radial-positions in the OTSG -(Center, Average and-Periphery) as a function'of time

- from the start of the MSLB transient. The peak load of 1310

--lbs occurs 60 seconds _.into'the transient at_the periphery of-the-0TSG. The smallest loads occur at the center of the OTSG '

as:was discussed earlier.  ;

4.3.2 FTl OTSG Tube Loads

~

The FTI results (Reference 22) are provided in Figure 4-2.

As can be seen,-they are very similar to the GLUN load results. The peak load is 1135 lbs at 60 seconds and also occurs at the OTSG periphery with.the smallest loads at the center as well.

A comparison of the GPUN and FTI results are provided in section 4.4 below with an explanation for the loads that were used to perform the subsequent tube to tubesheet

-interface pressure and the leakrate analyses which are described in sections 5.0 and 6.0.

4

=

25-

uncca tssessutw utunxxxcznor tw ors uncre exxusov tx4use r:c.v revsen o Figure 4-1 GPUN Tube Loads 5400 1

^Z%si 1 00 _ _ _ _ _ _ _ _ _ _ _ ._ _ _- __ -..____...-*^*9* ._

"**Y"_,

1000 _ _ _ _ . _ _ _ _ _ __ ._. -

800 -- ~-- _- - ---

f3 00 _ _ . . _ . , . . _ _ _ _ _ _ _ _ _ . _ -__ ___ --

q h400. -- - -- _ _ - - - -

200

/

ir 0 --- ---- - -

300 1 10 100 1000 10000 100000 Time (sec)

FIGURE 4-2 FTl Tubo Loads 1400 G Raeus o'55 4*

1200 . . _ - . . - _ . _ - - - - - - - _ _ -- Raous s 36* ' ,_

y Radius e 6",

1000 - - - . - . - _ . - - - - __ . _ -

800 4- - _ _ - _ _ . _ - - _ - ___.

i ,,

j 2 .==-

f400 . _ _ _ _ _ _ _ m . _ _ _ _ _ #_ -

,0. 2 - - .

0 . _ _ _ _ _

1 10 100 1000 10000 100000 Time (sec) 26

awcu nsuswim wrnwacer rea rw orse mere somsa twurm

  1. cvss:r10 4.4 Analysis of Loads Figure 4-3 provides a comparison of the FTI and GPUN OTSG tube load results. Results are presented for three points in time as a function of radial distance from the OTSG center (R-0.0) to the periphery. While the results are very close, y_toure 4-3 Tube Load Comparison [" Lg 50 **co, Long-Tenn Mawnum 24600 sec

$400 .- 1.1

+ FTI Dela (60 secs 1300 -

"' "I - - - - - - - - - - - - - - - _ - , - - - - - - -- 1.0

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

+ F fi Data (2e600 sec)

-+-- GPUN Data (24600 sec) 08 9,

_=.-- Area Rase _ _j-------~--..---

a*'

i oco . _ _ _ . . _ . _ _ _ . _ _ _ . _ _ - . _ _ _ -. __ .-

o.7 fu m _ _ _ . _ _ _ _ , ..__ _ _ . - . ___. ._ _- _- oa t,*' ooo _ _ . - _ _ _. _ ____ _ . _ __ _ _ _ _ _ ._ . _ _ _ - - . os sco  ; _ _ _ _ _ _ . _ _ _ _ _ . _ _ _ _ _ _ _ _ _ - . _ . - - - - - 3.3 500 - - - - - - - - - - - - - - - - - - - - - - - - - - - - 0,2 m , ,

~

3co ' - oo 0 $ 10 il 2o 28 30 35 de 48 80 55 to Tubesheet Rmkws (in) it can be seen that the GPUN results tend to be more conservative than the FTI results as R increases. Similarly, for smaller R, the FTI results are slightly more conservative. The plot of area ratio vs. radial position (right side ordinate axis is the area ratio) shows that there are substantially more tubes at the higher R values than at the lower R values. It was judged that the GPUN results would be more conservative since they would result in higher loads to a greater number of tubes. As a result, for this study, the GPUN loads were used to perform the crack area and crack leakage analyses described below.

The two sets of independent analyses are confirmatory and demonstrate that the calculated OTSG tube loads are reasonable.

l 27

ancer asussww waaocuenm onc mem comm cumo revis*0n 0 5.0 CRACK AREA DETERMINATION 5.1 Introduction The crack area determination is based upon the methodology provided in Reference 13 and establishes a method for calculating the crack opening area for through-wall cracks in tubes. Primary-to-secondary leakage is calculated using two potential crack orientations in combination with a specific applied load (Reference 14). These are:

1. Circumferential Through-Wall Crack in Tension (Note: The contribution of primary pressdre is included in the applied tension load)
2. Axial Through-Wall Crack Subjected to Internal Pressure Using these methods, the user can calculate the crack opening area (COA) for a crack given the specified conditions and use that area to determine the tube leakage (see Section .0).

There are conditions particular to the capture of the tube within the kinetic expansion region that separate the COA within the kinetic expansion from the COA for a defect in the free span.

It is arguable that any COA occurs at all within the expansion because the tube will not slip or rotate within the expansion. Within any expansion region, the tubesheet, due to its proximity alone, guides the tube and prevents rotation at the elevation of a defect that could result in increasing COA. In addition, remaining contact pressure on the tube OD surface further provides a friction reaction that prevents bending of the tube that could result in increasing COA.

Therefore, for the purpose of leakage assessment from expanded tubes, COA depends on applied axial tension only because there is no rotation at the elevation of a defect due to remotely applied tension. COA develops because of asymmetry local to the section as the symmetrically distributed load comes into equilibrium with the asymmetrical section containing the defect.

NUREG/CR-3464 (Ref.13) provides the solution for COA for circumferential defects in OTSG tubes for applied axial tension. The COA for axial defects is also provided. This reference has been widely used in the nuclear industry and, in particular, is the source for COA for leak-before-break analysis of RCS piping in B&W plants (Reference 18).

28

lfAKACE A$5E$5 CENT ECTrVXICN,0CY 500 Th0101%C K9YETC (KPANSJCW EXAhthATON 9tvosot10 5.2 Methodology (Kinetic expansion mglon)

Reference 13 provides the equations necessary to establish the methodology to calculate the crack opening area for circumferential through-wall cracks in tension and axial through-we'.1 cracks subjected to internal pressure. The methodolcgy was implemented in Reference 14 and summarized herein.

5.2.1 Circumferential Through Wall Crack in Tension The crack opening area for the tensile load is calculated based on the applied axial stress (o ) ,

Young's Modulus (E) for the tube material, and a non-dimensional function (I t (0)) formulated from the stress intensity factors.

A,=E(rRjf,(g) 2 E

The applied stress is calculated given the axial tensile load (P) and the mean tube radius (R) with the tube wall thickness (t) or the inner and outer tube radius (R and R , respectively): i P P

=

cr' =

2xRt x(R,' - R )

5.2,2 Axial Through Wall Crack Subjected to Internal Pressure The crack opening area for a axial through-wall crack with internal pressure is calculated based on the membrane stress (c), Young's Modulus (E) for the tube material, mean tube radius ,(R), tube thickness (t), and a non-dimensional function (G ( A) ) formulated from the stress intensity factors.

A = E(2xRt)G().)

E 29

u sace nssessus wr uernocaw

, rgrsc worre tmo t awron The applied stress is calculated given the differential pressure (p), tean tube radius (R), and tube thickness (t): i cr = d I

This methodology can be used to calculate the crack opening area for through-wall cracks of tubes with an outer radius te, wall thickness ratio (R/t) of less then or equal to 10.0 with no bending moment applied. The crack opening area for R/t ratioe of less than 10.0 are conservatively large.

J 6.0 CRArK AREA LEAKAGE ANALYSIS 6.1 Overview The leakage flow for a given crack area (from section 5.0) is determined by the PICEP (Eipe . Crack Evaluation Erogram code) computer code developed by EPRI (Reference 15). A briof description of the code is provided in this section.

The crack area as a function of time for a given crack length and crack orientation is provided from the analysis oescribed in Sr. ion 5.0 above. The T-H parameters are provided in Sect.un 3.0 above. The PICEP code utilizes a crack area, the RCS pressure, RCS temperature and OTSG pressure at a single point in time and calculates a leakage rate through the crack for that cpecific time. In order to develop a leakage rate as a function of time, the code has to be run numerous times throughout the transient duration.

Currentiv, the PICEP analyais la run at the data intervals.

. gl l

The resuit is a leakage rate as a function of time which can be integrated to provide a total leakage. This process must be repeated for each crack indication (sce Section 7.0).

The coitact pressure between the expanded tube and the tubusheet causes a significant reduction in leakage. This was determined empirically and the evaluation for this

' leakage reduction f actor' (LRF) is provided below. A discussion of the method used to calculate the contact pressure is also provided in this section.

m 30

(fMAGI Assessa,f avt 4,f fH000LOCY F0f 72 Of$0 #Witt ipAV5dO% ( AMf4AT:04 twnton 0 6.2 Code Description The PICEP program (Reference 15)can calculate the crack opening area, the critical crack length and the flow rate through cracks in steam generator tubes. options are available to calculate the leakage with a crack area that is supplied by the user. For subcooled or saturated liquid discharge, the critical flow equations are based on the Henry /Fauske homogeneous non equilibrium critical flow model with modifications to account for fluid friction due to surface roughness, crack turns and non equilibrium

' flashing' mass transfer between liquid and vapor phases.

The flow is assumed to be isenthalpic and homogeneous with non equilibrium effects introduced through a parameter, N, which is a function of equilibrium quality and flow path length-to-diameter, L/D, ratio. The model was validated with experimental data.

6.3 Leakage Reduction Factor The primary-to-secondary leak rate test results report (Reference 20) provided the basis for identifying a minimum leakage reduction factor due to contact prese'.re between the expanded tube and tubesheet as well as a justification for neglecting any contribution to leakage from potential defects located further into the expansion than the minimum required inspection distance to assure structural integrity.

Primary-to-secondary leakrate tests were cunducted using a bolted split clamp assembly. Increasing tube to clamp contact pressure was achieved by applying increasing torque to the bolts in the assembly. Additional experimental components provided the capabilities to achieve a very wide range of primary and secondary temperatures and pressures as well as to develop an axial tube load. Each OTSG tube specimen contained a small, through-wall, Electro Discharge Machine (EDM) circumferential notch. By sliding the tube specimen within the clamp, primary-to-secondary leakage could be measured with the notch within the clamp or with the notch outside the clamp representing the free span condition.

The general trend of results showed a dramatic reduction in leakage for a minimum applied contact pressure equal to 500 psi and a minimum lea 4 path length equal to 0.25 inches.

There is little or no benefit derived from increasing tne contact pressure above 500 pai to as much as 3000 pai. It was necessary to remove the influence of tube internal pressure from the test results and isolate only the effect of increasing contact pressure in the derivation of the leakage reduction factor (LRF). This was accomplished by using as a basis for comparison the zero applied contact 1

31

anaa csussueur uernanocgsgrso nnerc smus<n unumrm pressure results which are representative of the effect of tube internal pressure alone. With this as a basis, the effect of applied contact pressure alone was determined using results obtained for increasing applied contact pressure since che same tube internal pressure was used throughout the tests. It was also necessary to remove the effect of thermal tightening from the test results. This waa accomplished by testing at cold conditions.

The leakrate for a contact pressure equal to 500 psi is about (1/36) of that for zero contact pressure. Using ASME Code guidance for faulted conditions as a basis for establishing a safety factor, only 70% of maximum capability should be used. The LRF is (1/25) or 4 x10 . The LRF is associated with a reference notch location 0.25 inches from the edge of the clamp, or, effectively a 0.25 inch leak path length. The LRF as a function of contact pressure alone will be different for other notch insertion depths because of the influence of the leak path length. The reduction in leakage due to co.itact pressure alone will be less for greater notch insertion because there is less leakage to begin with due to increased leak path length. Leakage decreases by about 20% for 0.125 inches additional insertion without applied contact pressure. Correspondingly, reduction in leakage due to contact pressure alone will be more for notch insertion less than 0.25 inches because of greater leakage without applied contact pressure. The engineering resolution of this issue was to assume that the apparent reduction in LRF due to increased insertion is compensated by the effect of increased leak path length alone and to take no credit for the apparent LRF increase due to decreased insertion. The LRF associated with a reference notch location equal to 0.25 inches was judged to be a practical gauge of overall leakage reduction due to contact pressure remaining in the expanded tube to tubesheet joint.

The results also suggest that less remaining contact pressure than that used in the test will be equally effective in reducing leakage since there is no benefit from increasing contact pressure to reduce leakagc. Leakage reduction ie not prorvrtional to the magnitude of remaining contact pressure but is achieved by establishing and maintaining minimal contact pressure, independent of magnitude. In order to account for uncertainties in the method used for calculating contact pressure, it was decided that a threshold for using a LRF should be 250 psi at a minimum.

The criteria for application of the LRF in the field are that the defect must be located at an elevation at which structural analysis results identify a remaining contact pressure at least equal to 250 psi and a leak path length of at least 0.25 inches from the expansion transition. Defects 32

1 1

I umce cunwor mnmocrgona rme tamsat ta===

that are not clamped by a least 250 psi over a leak path l length of at least 0.25 inches were evaluated without a LRF.  ;

The structural analysis results that identify remaining .

contact _ pressure for leakage assessment purposes are based j on applied axial tube load, tubesheet deflection, tube d internal pressure and thermal tightening that is adjusted .

for tube to tubesheet temperature differences where appropriate. l t

Any leakage contribution due to possible defects located further into the expansion than the minimum inspection

- dd stance was considered negligible. Established calculation '

methods for leakage through cracks show that leakage is inversely proportional to the length of the leak path. The experimental results discussed above show a 20% leakat  ;

reduction for an additional leak path length of 0.125 s :hes  ;

without applied contact pressure. The minimum inspectic.1 length was 1.8 inches from the transition for peripheral i tubes. There is both a theoretical and experimental basis ,

for assuming that the flow resistance due to 1.675 inches of additional leak path length with applied contact pressure >

would effectively prevent additional leakage.

6.4 Contact Pressure Determination 7 The application of both the leak reduction factor (LRF) discussed above and the crack opening area (COA) solution (Section 5.0) requires an assessment of remaining contact pressure within the kinetic expansion. It is necessary that ~

contact pressure be established and maintained so that both LRF and COA are correctly applied.

For the 17 inch kinetic expansions, for purposes of leakage assessment, recent analysis results (Reference 19), show that the minimum contact pressure equal to 250 psi is established and maintained at all times throughout the  ;

expansion regardler, of location within the tube bundle.

Because.of these' conditions, use of the LRF is appropriate -

without exception. For the 22 inch kinetic expansions, the analysis results (Reference 19) also show that minimal contact pressure is established and maintained at all times beginning at the center of the u' nit at an elevation above- ,

the transition equal to 0.86 inches. The tubesheet radial

- location at which minimal contact pressure is established >

at the transition begins at 0.2R and the radial location at  :

which 250 psi contact pressure is maintained at the transition begins at 0.36R.

The use'of the LRF begins'at that location. The application ofithe leakage model using the COA solution is appropriate s 33

- - ,- a- -.-,.-.,i_.-.._. ,.-i..s.<- . . _ , . - - - - - - . . - _ - . . - . . - - - - . . - _ _ . . ~ - - - - - - _ . - . - . . - . - -- - - , - _'

unactcsosswar m.comocr om ewere roavsm tuvem everywhere except for the 22 inch expantion between 0.0R and 0.2R for defect location up to 0.86 inches from the expansion transition.

This guidance is captured in Table 6-1 which is to be used for leakage assessment determination.

By using this table, a determination can be made whether an LRF can or cannot be used to modify existing leakage calculations or whether an additional calculation of free-span leakage is required. Leakage assessment of each flaw indication can be accomplished given the length of the tube expansion, the radial position of the tube and the elevation of the' flaw indication with respect to the transition location.

9 34 t

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

,,,,,,,,m , ,. . .-m. ,

umce estis=w wr.a.acerg grsc <arre eme rawara l TABLE 6-1 I i

LEAKAGE ASSESSMENT DETERMINATION f i  :

TUBE RADIAL LENGTH OF KINETIC EXPANSION LEAKAGE ASSESSMENT  :

POSITION METHOD  :

.. 17 inch KE l 1  ;

ETL TO ETL MINUS CALCULATIONS REQUIRED OR to 1.OR ETL TO ETL+0.25 USE TABLES ETL+0.25" TO ETL + REQUIRED USE TABLES AND LRF i i

ABOVE REQUIRED no tggggag 22 inch KE ETL MINUS TO ETL+0,86" (ALCULATIONS REQUIRED t

OR to 0.2R ETL +0.86" TO ETL + 5.25" USE TABLES  ;

f ETL +5.25" TO ETL + REQUIRED

  • USE TABLES AND LRF ABOVE REQUIRED NO LEAKAGE  ;

ETL TO ETL MINUS CALCULATIONS REQUIRED 0.2R to 0.36R ETL TO ETL v5.0 USE TABLES ,

ETL +5.0" TO ETL + REQUIRED

  • USE TABLES AND LRF ABOVE REQUIRED NO LEAKAGE ETL TO ETL MINUS CALCULATIONS REQUIRED 0.36R to 1 OR ETL TO ETL +0.25 USE TABLES

, ETL +0.25" TO ETL + REQUIRED

  • USE TABLES AND LRF ABOVE REQUIRED NO LEAKAGE
  • Required for structural integrity t

+

35

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7.0 TOTAL LEAKAGE EVALUATION 7.f Ovendow  !

This section describes the approach taken to determine the total leakage for the purposes of comparison against the leakage limits. A calculation methodology was developed that t integrates the OTSG tube loads with the thermal hydraulic data and analysis needed for leakage through the cracks and combines the results into leakage assessment tables. This is basically _ implementing the_ methodology discussed in Sections 1

- 3.0 through 6.0 above. Also discussed in this section are l the-ways in which the unaffected OTSG can be treated since i the loads are quite different (smaller) and the steamline is l intact.

7.2 Lenkage Results A calculation which applied the methodology discussed in i carlier sections of this report to calculate the leakrate through a crack in a tube in the tubesheet region of an OTSG

was implemented (Reference 21). The crack could be either circumferential or longitudinal (axial). If a crack was indicated to be volumetric, it was treated as two cracks.

Each dimension was treated independently so that there would be one axial and one circumferential cracks with the appropriate lengths. This is a conservative representation of a volumetric indication for the purposes of total leakage evaluation. The crack opening area was calculated based on the tube tensile load or the differential pressure depending on the orientation of the crack. The mass flux was calculated using the PICEP computer program given the crack geometry and the fluid properties as discussed in Section 6.0. The mass flux was converted to a volumetric leakrate ,

based on a reference density (579 deg F and 2200 psi) and the crack opening area. This reference density corresponds to the same value used in determining the leakage limits.

The leakage is integrated over a period of 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and for the auration of the MSLB transient. The results of this calculation can be provided by 'binning' of integrated leakage from cracks in a range of sizes for circumferential F

-and axial-leakage. The crack size bins for a given radial position in the OTSG are the same, but the integrated leakage for a given crack size is different as a function of radial position. This is true for circumferential crack leakage-but not axial crack leakage which is not sensitive to radial position, only differential pressure, ,

The circumferential crack integrated leakage results  ;

presented as leakage tables according to crack size for 10 radial positions are provided in Table 7-1. For axial  :

36

su<cer asu nwer mocoaangorso mm twmm rum rum

-cracks, the leakage is provided as crack size bins. The bins for all of the circumferential crack tables range from 0.05 inch crack size (.05 inch leakage is used for all cracks from 0.01 to 0.05 inches) through 0.65 inches. The ,

lonoitudinal (axial) leakage table covers cracks up to 1 inch in length. In the field all circumferential and axial extents are ' rounded up' to the next 0.05 inch increment.  ;

The leak volumes given in Table 7-1 do not include the t application of the Leakage Reduction Factor (LRF). This factor of 1/25 is to be applied on a case-by-case basis as per the guidelines discussed in Section 6.0 (i.e. a minimum contact pressure and minimum leak path length are required).

7.3 ANected OTSG Vs UnaWected OTSG Since both.the affected OTSG and the unaffected OTSG will experience tube loads, leakage is expected to occur in both generators. Since either of the two OTSGs might be the affect-d one, it is necessary to assume that the OTSG tith the greater number of crack indications is the affected one.

Tae leakage from each of the crack indications has to be summed, and the total leakage for the OTSG can then be compared against the total leakage limits of 3228 and 9960 ,

gallons (e 579F, 2200 psia) for the 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> SAB and 30 day LPZ, respectively, discussed in Section 1.0. A ' running total' after each crack is evaluated will actually be maintained as shown in Figure 2-1. In this way, if the leakage is found to exceed the limits, a decision can be  ;

made regarding which tubes will be repaired to bring the postulated leakage back within allowable limits.

Since OTSG tube loads were not specifically determined for the unaffected OTSG, it will be necessary to treat the unaffected generator as if it had the same loads as the  ;

affected generator. Thus, the same process used for the affected OTSG will be used for the unaffected OTSG with the -

difference that the unaffected OTSG will experience a decontamination factor of 133 as a result of the intact steamline and the availability of the flow path through the ,

-turbine bypass and main condenser. This would be accounted for'in the total leakage evaluation. It would be very-conservative to assume that both generators would leak as if '

they were both affected generators and determine the leakage to be based on the sum of the cracks in both generators

. without taking credit for the intact steamline of the unaffected generator, t

(,

37 N - _

o

?

LI ACCf C81(15WWr Mf7HODOLOG7 FDP e 48 Or$C rM7C spAWSOY i AAM40 ROV tre mon 0 Table 7-1 Leakage Assessment Evaluation Data (R values refor to inches from center of OTSG)

(' Longitudinal' refers to an ' axial' crack)

Center: 0.0 < R <a 0 10%: 0 < R <a 6.7426 C6rcumferential 2 Hour Duret60n Circumferent6at 2 Hour Duration Single Crack Single Crack Single Crack Single Crack  !

Length (in) Leakage (gal) Leakage (gal) Length (in) Leakage (gal) Leakaos (gal) 0 01 0 00 0 00 0 01 0 00 0 00 0 05 0 00 0 04 0 05 0 00 0 04 0 10 0 02 0 26 0 10 0 03 0 27 0 15 0 07 0 79 0 15 0 07 0 80 0 20 0 16 1 79 0 20 0 17 1 82 0 25 0 32 3 48 0 25 0 32 3 54 0 30 0 55 6.15 0 30 0 57 6 25 0 35 0 91 10.17 0 35 0 93 10 34 0 40 1 44 16 51 0 40 1 47 16 84 0 45 2 32 27.20 0 45 2 38 27.75 0 50 3 64 42 94 0 50 3 74 43 76 0 55 5 54 65 40 0 55 5 67 66 59 0 60 8 19 96 73 0 60 8 38 98 41 0 65 11 85 139 69 0 65 12 11 142 00 Hot Lenkage Based on a densny of 0 7094 gm/cc Hot Leakage Based on a densny of 0 7094 gm/cc 20%: 6.7826 < R <= 11.626 30%: 11.626 < R <= 17.2876 C6rcumferer.tml 2 Hour Duration Circumferent6al 2 Hour Duration Single Crack Single Crack Single Crack Single Crack Length (in) Leakage (es.1) Leakage (gal) Length (in) Leakage (gal) LeekaDe (gal) 0 01 0 00 0 00 0 01 0 00 0 00 0 05 0 00 0 05 0 05 0 01 0 05 0 10 0 03 0 28 0 10 0 03 0 31 0 15 0 08 0 85 0 15 0 09 0 94 0 20 0 18 1 93 0 20 0 21 2 14 0 25 0 35 3 77 -

0 25 0 40 4.16 0 30 0 61 6 66 0 30 0 70 7.34 0 35 1 01 11 00 0 35 1.15 12 15

~

0 40 1 61 18 21 0 40 1.88 20 61 0 45 2 63 29 93 0 45 3 05 33 69 0 50 4 10 47.04 0 50 4.74 52 66 0 55 G 21 71.34 0 55 7.13 79 45 0 tiO 9 14 105 10 0 60 1044 116 44 0 65 13.17 151.16 0 65 14 95 166 66

- Hot Leakage Based on a densny of 0 7094 gm/cc Hot Leakage Baned on a densny of 0 7094 gm/cc 38

__, - . _ , . _ . _ _ . . , _ . - ~

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revnum 0 l

Table 7-1 (Continued) f 40%: 17.2875 < R <= 23 06 60%! 23.05 < R <= 28.8125 C6ccumferent6al 2 Hour Duratinn Circumferenttal 2 Hour Durat6on Single Crack Single Crock Single Crack Single Crack Length (in) Leakage (gel) Leakage (gel) Length (in) Leakage (gal) Leakage (gal) 0 01 0 00 0 00 0 01 0 00 0 00 0 05 0 01 0 06 0 05 0 01 0 07 0 10 0 04 0 35 0 10 0 04 0 41 0 15 0 11 1 07 0.15 0.13 1.23 0 20 0 24 2 42 0 20 0 28 2.78 0 25 0 46 4 70 0 25 0 55 5 40 0 30 0 81 8 30 0 30 0 96 9 54 0 35 1 34 13 98 0 35 1.66 16 68 0 40 2 28 24 02 0 40 2 79 28 43 045 3 66 38 99 045 4 44 4578 0 50 5 64 6054 0 50 6 80 70 55 0 55 8 42 90 73 0 55 10 07 104 95 0 60 12 25 132 12 0 60 14 53 151.75 0 65 1743 187 93 0 65 2053 214 41 ,

Hot Leakage Based on a C'snsRy of 0 7094 gm/cc Hot Leakage Based on a densay of 0 7094 gm/cc 60%: 28.8125 < R <= 34.378 70%: 34.675 < R <= 40.3375 C6ftumferent6al 2 Hour Duratson C6rcumferenttal 2 Hour Duration Single Crack Single Crack Single Crack Single Crack Length (in) Leakage (gal) Leakage (gal) Long'.h (in) Leakage (gal) Leakage (gal) 0 01 0 00 0 00 0 01 0 00 0 00 0 05 0 01 0 08 0 05 0 01 0 09 0 10 0 05 0 47 0 10 0 06 0 55 ,

0 15 0 15 1 42 0.15 0 18 1.65 0 20 0 34 3 22 0 20 0 40 3.73 0 25 0 65 6 25 0 25 0.77 7.24 0 30 1.15 11 08 0 30 1 42 13.30 0 35 2 05 20 02 0 35 2.51 23 95 0 40 3 41 33 81 0 40 4.15 40 09 0 45 5 40 53 99 0 45 6 51 63 50 0 50 8 19 82 54 0 50 9 80 96 28 0 55 12 04 121 84 0 55 14 30 141.04 0 60 17,25 174 88 0 60 20 32 200 99 0 65 24 18 245 40 0 65 28 29 280 20 Hot Leakage Based on a densty of 0 7094 gm/cc Hot Leakage Based on a densay of 0 7094 gmice 39

u AOCf AS$(55WW W7NODCt0CW0t fr 0f5C DV(TC ( Amsm in AwCrm  :

revsen o l

l Table 7-1 (Continued) 80W 40.3375 < R <= 48.1 80W 48.1 < R <= $1.8625 Circumferential 2 Hour Durstson Circumferent6al 2 Hour Duration Single Crack Single Crack Single Crack Single Crack Length (in) Leakage (gal) Leakage (gal) Length (in) Leakage (gal) Leakage (gal) 0 01 0 00 0 00 0 01 0 00 0 00 0 05 0 01 0 11 0 05 0 01 0 12 0 10 0 07 0 63 0 10 0 08 0 73 0 15 0 21 1 90 0 15 0 25 2 18 0 20 0 47 4 30 0 20 0 55 4 94 0 25 0 92 8 35 0 25 1 10 9 69 0 30 1.74 15 93 0 30 2 09 1888 0 35 3 05 28 43 0 35 3 64 33 39 0 40 4 99 47 19 0 40 5 91 54 96 0 45 7.76 74 12 0 45 9 12 85 65 0 50 11 60 111 51 0 50 13 53 127 90 0 55 16 79 162 14 0 55 19 45 184 70 0 60 23 69 229 50 0 60 27 28 259 81 0 65 32 78 318 02 0 65 37.51 358 47 Hot Leakage Based on a density of 0 7094 gm/cc Hot Leakage Based on a density of 0 7094 gm/cc Peripheral: 61.8626 < R <= 57.62f All Tubes Circumferental 2 Hour Duration Longitudinal 2 Hour Duration Single Crack Single Crack Single Crack Single Crack Length (in) Leakage (gal) Leakage (gal) Length (in) Leakage (gal) Leakage (gal) 0 01 0 00 0 00 0 01 0 00 0 00 0 05 0 02 0 14 0 05 0 01 0 02 0 10 0 10 0 84 0 10 0 04 0 13 0 15 0 29 2.50 0 15 0 12 0 45 0 20 0 64 5 65 0 20 0 31 1.19 0 25 1.32 11.51 0 25 0 70 2 73 0 30 2.50 22 24 0 30 1.53 5 72 <

0 35 4 31 38 98 0 35 3 14 11 21 0 40 6 94 63 64 040 $ 81 20 49 0 45 10 64 98 41 045 9 87 36 51 0 50 15 67 145 89 0 50 15 64 61.20 0 55 22 37 209 32 0 55 2345 96 42 0 60 31 18 292 75 0 60 33 61 144 31 0 65 42 65 401 91 0 65 46 45 206 92 Hot Leakage Based on a densty of 0 7094 gerece 0 70 62.33 286 28 0 75 81.64 384 43 0 80 104 81 503 54 0 85 132 33 646 46 0 90 164 68 815 41 1.00 245 97 1238 97 Hot Leakage Based on a density of 0 7094 gm/cc 40

aamumaswaaawrggnemememumm 8.0

SUMMARY

AND CONCLUSIONS i A methodology is described which allows for a determination 4

of the leakage-which may occur during a Main Steam Line ,

Break (MSLB) event from conservatively assumed through-wall cracks in the kinetic expansion region of the upper tubesheet of both OTSGs. Cracks.which have eddy current through-wall indications greater than an estimated value (67.4%) are assumed to be 100% through-wall cracks that will leak during the MSLB.

The. amount of leakage is determined by calculating the leakage area resulting from the event induced. tube loads (differential pressure only for axial cracks) and the subsequent leakage flow rate and total event integrated  !

leakage for each applicable, crack indication based upon the- i thermal hydraulic conditions associated with the event. The total leakage for-all cracks is compared against 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and event duration leakage limits. The leakage limits are associated with exclusion area boundary and 30 day low population zone doses which would not exceed a small fraction of 10 CFR 100 requirements if the event were to  ;

occur.

The implementation of this methodology using OTSG Eddy ,

Current data provides reasonable assurance that the leakage that could occur during a design basis MSLB from indicated cracks in the kinetic expansion regicn may be conservatively ,

determined.

t J

s 4

t

'41 w .. - -be.. + . , - -E-,--r-.. , - . - - + ..-*w.-.. w, ..r., ..- --- + - . - - < <, .- . e.-',-= . . - - . . - = _. - --____.,.--w e.,- --

-r-- - - - - - --v-r----we r

acucs cssessutur un,emccrg gne mnc amwsa tawwm

9.0 REFERENCES

1. Letter from James W. Langenbach to USNRC , 'Once-Through Steam Generator Kinetic Expansion Inspection Criteria', dated August 08, 1997 (6710-97-2348),
2. GPUN Calculation No. C-1101-224-6612-057, 'Offsite Doses from OTSG Tube Leakage Due to Main Steam Line Break', August 1997.
3. Letter from James W. Langenbach to USNRC , 'TMI-1 License Amendment Request No. 269 ' Revised Steam Line Break Accident Analysis Dose Consequence', August 14, 1997 (6710-97-2345)
4. GPUN Calculation No. C-1101-224-E610 061, 'MSLB Analysis for OTSG Tube Integrity', dated August 21, 1997
5. Letter, USNRC to Mr. C.R. Lehmann, ' Acceptance for Referencing of the RETRAN-02 MOD 005.1 Code', April 12, 1994.
6. GPUN Calculation, C-1101-202-5412-114, Rev. 2. 'TMI-1 RETRAN Base Deck'
7. Weimer, J. A., ' Thermal Mixing in thc Lower Plenum and Core of a PWR', EPRI NP-3r45, May 1984.
8. GPUN Calc C-1101-224-E610-060, 'Long Term MSLB Transient Data for Tube Leakage Calculations', September, 1997 9 GFUN Procedure .sTP 1210-10, ' Abnormal Transients Rules, Guides and Graphs,' Rev. 32, 10 GPUN Procedure ATP 1210-3, ' Excessive Primary to Secondary Heat Transfer,' Rev. 20, 11 GPUN Calculation C-1101-224-E610-058, ' Post MSLB Cooldown of a Voided OTSG,' Rev. O, September,1997.
12. GPUN Procedure OP 1103-6, ' Reactor Coolant Pump Cperation', Rev. 62.

42

i uanact assissurvr utr,<xaocr,rgryrseswarca4Ssmsnauwarow 13 NUREG/CR3464, 'The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulating Circumferential Through-wall Cracks',

September 1983

14. GPUN Calculation C-1101-224-E610-054 Rev 1, ' Tube Crack Opening Area Calculation Methodology', August 14, 1997
15. EPRI NP-3596-SR, Revision 1, Special Report, 'PICEP:

Pipe Crack Evaluation Program (Revision 1)' December, 1987.

16 GPUN Calculation C-1101-224-E520-061,Rev 0, ' Methodology for Calculating OTSG Tube Axial Loads', September 30, 1997

17. Babcock and Wilcox Report No. BAW 10146, ' Determination of Minimum Required Tube Wall Thickness for 177-FA Once-Through Steam Generators,' dated October 1980.
18. Leak-Before-Break Evaluation of Margins Against Full Break for RCS Primary Piping of B&W Design NSS, BAW-1847, Rev. 1, 77-1153295-01, 9/85.
19. Analysis of Remaining Contact Pressure for TMI OTSG Kinetically Expanded Tubes for Leakage Assessment Purposes, enclosure to letter from H. W. McCurdy (MPR) to S. D. Leshnoff (GPUN), dated 9/29/97, 'Three Mile Island Generating Station Development of OTSG Kinetic Expansion Inspection Acceptance Criteria.'
20. TMI-1 EDM Notch Sample Hot / Cold Leak Tests Results, Framatome Technologies Inc (FTI), document number 51-1264463-00, 8/15/97.
21. GPUN Calculation C-1101-224-E610-059 Rev 0, 'OTSG Tube Leakage Methodology for Tubesheet Region', September 1997
22. GPU MSLB Tube Load Summary, Framatome Technologdes Inc.

(FTI) Document 51-5000542-00, dated September 19, 1997

23. GPUN Calculation C-1101-224-ES20-063, Rev 0 'TMI OTSG Analysis of Tube Axial Loads During Plant Specific MSLB' 43 l

- _ _ _ _ - _ _ _ _

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